[ { "image_filename": "designv11_5_0002497_j.ijleo.2021.166917-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002497_j.ijleo.2021.166917-Figure2-1.png", "caption": "Fig. 2. The LEW; a) Schematic of process condition, b) The welded specimen.", "texts": [ " Sample number p (W) v (mm/s) f (mm) 1 1600 35 50 2 1600 30 50 3 1600 25 50 4 1600 30 40 5 1750 30 50 6 1750 25 40 7 1900 35 50 8 1900 30 50 9 1900 35 40 Table 2 Stainless steel 316L chemical composition (wt%). Cr Ni Mo Mn Si C P S Fe 17.0 12.0 2.5 1.5 0.5 0.03 0.03 0.05 Balance S.F. Nabavi et al. Optik 241 (2021) 166917 5 Stainless steels is one of the most beneficial materials due to its corrosion resistance [15]. AISI 316L stainless steel sheets with a thickness of 1 mm and chemical composition shown in Table 2 are used in the current experiments. All AISI 316L sheets were cut into 130 mm \u00d7 250 mm \u00d7 1 mm specimens and then bent 90\u25e6 shown in Fig. 2(a). The specimens were clamped along the intersections to ensure a no-gap region along the edge. Prior to the LEW process, all specimens were rinsed with water and washed with alcohol to remove any contamination. In order to analyze microstructure properties, specimens were prepared using SiC grit paper with grit sizes ranging from 360 to 2000, and later polished with alumina powder. The specimens were etched in Glyceregia etchant for macrography tests to reveal weld width and penetration. Oxalic acid 10% etchant was used for metallography test to analyze the microstructure. The microstructure was analyzed using optical microscopy with imaged obtained using an IMM420 microscope. Mechanical properties including the fracture load were measured by the T-peel test based on ASTM D1876 [16]. After the LEW process, the specimens were sectioned by water jet for geometrical, mechanical and microstructural examination (Fig. 2(b)). The distortion produced along the weld bead during LEW process was measured by the indicatory clock Mitutoyi FJY229. The clock was put on the flatbed to calibrate a zero point. Each weld edge was marked into 1 cm regions as shown in Fig. 3(a). The height of each marked point was measured by the indicatory clock as shown in Fig. 3(b). It should be noted after finishing the measurement of each specimen, the indicatory clock was put on the flatbed to ensure the correctness of calibration. Moreover, the flatness of all specimens before welding was measured to ensure that the measured distortion was a result of the LEW process" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001557_s12555-020-0110-9-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001557_s12555-020-0110-9-Figure2-1.png", "caption": "Fig. 2. Tilt tri-rotor UAV coordinate system.", "texts": [ " The rear rotor is used to compensate for the moment generated by two front rotors to stabilize the pitch. The yaw moment is created by the difference of tilting angles between front rotors. Besides, the altitude is controlled by adjusting the thrust of all rotors. 2.2. Nonlinear equations of motion The dynamic model of the tilt tri-rotor UAV is complex due to its time-varying structure. The design of the controller and allocator is based on the system model so that the dynamic and kinematic equations should be derived first. Fig. 2 shows the coordinate systems of the aircraft. The term \u03a3e represents the world coordinate system, \u03a3b denotes the body coordinate system, and \u03a3ri represents the rotor coordinate system. It is noted that the right rotor, the left rotor, and the rear rotor are labeled by 1, 2, 3, respectively. The tilting angles [\u03b11,\u03b12,\u03b13] T are defined as negative when the two front rotors tilt forward, or the rear rotor tilts backward. The transformation from the rotor frame to the body frame is denoted using the matrix Rb ri as follows: Rb ri = cos\u03b1i 0 sin\u03b1i 0 1 0 \u2212sin\u03b1i 0 cos\u03b1i , i = 1,2,3" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002356_978-3-030-50460-1-Figure9.2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002356_978-3-030-50460-1-Figure9.2-1.png", "caption": "Fig. 9.2 Graphic illustration of planar position and rotation fields.", "texts": [ " The position vector of the centerline, the tangent vector and its derivative with respect to the arc length parameter s for the reference and current configurations are given by r0(s) = X(s)e1 + Y (s)e2 , r(s, t) = x(s, t)e1 + y(s, t)e2 , r\u20320(s) = X \u2032(s)e1 + Y \u2032(s)e2 , r\u2032(s, t) = x\u2032(s, t)e1 + y\u2032(s, t)e2 , r\u2032\u20320(s) = X \u2032\u2032(s)e1 + Y \u2032\u2032(s)e2 , r\u2032\u2032(s, t) = x\u2032\u2032(s, t)e1 + y\u2032\u2032(s, t)e2 , (9.71) using the coordinate functions X,Y : I \u2192 R and x, y : I\u00d7R \u2192 R. Planar rotations are given by rotations around D3 = d3 = e3. As depicted in Fig. 9.2, the reference and current director triads are D1 = cos \u03b80(s)e1+sin \u03b80(s)e2 , D2 = \u2212 sin \u03b80(s)e1+cos \u03b80(s)e2 , d1 = cos \u03b8(s, t)e1+sin \u03b8(s, t)e2 , d2 = \u2212 sin \u03b8(s, t)e1+cos \u03b8(s, t)e2 , (9.72) where \u03b80 : I \u2192 R parameterizes the absolute angle of the reference director D1 with respect to the vector e1 and \u03b8 : I \u00d7 R \u2192 R the absolute angle of the current director d1. 112 Eugster, Harsch The rotation of the reference configuration is given by R0 = R0 ijei \u2297 ej , with the components R0 ij = ei \u00b7 Dj . The current rotation field is given analogously by R = Rijei \u2297 ej with Rij = ei \u00b7 dj " ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000745_iet-map.2018.6170-Figure10-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000745_iet-map.2018.6170-Figure10-1.png", "caption": "Fig. 10 Details of (a) The prototype, and (b) All components", "texts": [ " (a) Perspective view, (b) Lossless layer, (c) Top view of the lossy layer, (d) Bottom view of the lossy layer (b) The currents Za = Ra(\u03c92LbCb \u2212 1)2 \u03c92Ra 2Cb 2 + [\u03c92(La + Lb)Cb \u2212 1]2 + j \u03c9Ra 2Cb(\u03c92LbCb \u2212 1) + \u03c9La(\u03c92LbCb \u2212 1)[\u03c92(La + Lb)Cb \u2212 1] \u03c92Ra 2Cb 2 + [\u03c92(La + Lb)Cb \u2212 1]2 \u2212 1 \u03c9Ca (5a) Zb = Za Ra = 0 = j \u03c9La(\u03c92LbCb \u2212 1)[\u03c92(La + Lb)Cb \u2212 1] [\u03c92(La + Lb)Cb \u2212 1]2 \u2212 1 \u03c9Ca (5b) IET Microw. Antennas Propag., 2019, Vol. 13 Iss. 11, pp. 1777-1781 \u00a9 The Institution of Engineering and Technology 2019 1779 To verify the proposed band-absorptive FSR, a prototype of the dual-polarised one is fabricated, assembled, and tested. The assembled prototype and its\u2019 all components are shown in Fig. 10. The prototype is composed of 3 \u00d7 30 unit cells, with a size of 24 mm \u00d7 240 mm \u00d7 8.5 mm. The lossy layer and the lossless layer are supported and fixed by 8 nylon columns and 16 nylon screws (\u025br: 3.1\u20133.7). The substrates of two layers are F4BM220 with a dielectric constant of 2.2. There are 180 ordinary lumped resistors melded on the top and bottom surfaces of the lossy layer. As shown in Fig. 11, the prototype is measured in a parallelplate waveguide set-up, which is similar to that used in [20]" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002503_j.scitotenv.2021.145978-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002503_j.scitotenv.2021.145978-Figure1-1.png", "caption": "Fig. 1. The (a) schematics of air cathode MES stack with normal cathode module and the (b) schematic showing the connection of each anode brush and air cathode.", "texts": [ " Themass and energy balanceweremade under three different operation conditions of MES, including open circuit condition, low current condition, maximumpower generation and high current condition. The stacked air-cathodeMESwith the separately accessible alternating anode and cathode electrodes has the same design and dimensions (L: 22 \u00d7 W: 7 \u00d7 H: 13 cm, empty bed volume of 2 L) as previously applied (He et al., 2016c). The dual air cathode module equipped with pre-optimized wire spacers was plugged in the middle of the reactor (Fig. 1a, Fig. S1). Individually connected with one cathode layer, the two anode arrays separated by cathodemodulemade up two duplicated MES modules labeled as \u201cA\u201d and \u201cB\u201d. For each anode compartment, the empty bed volume was 0.86 L, while the working liquid volume was 0.7 L. The graphite fiber brushes (2.5 cm diameter, 12 cm long; MillRose, USA) were applied as anodes by holding carbon fibers (PANEX 35 50 K, Zoltek) between two twisted titanium wires (He et al., 2016c). The rolling-pressed stainless-steel-mesh air cathodes with activated carbon catalyst were applied as cathodes (He et al., 2016b). One anode array, which contained six anode brushes, was inserted into one MES module (module \u201cA\u201d and \u201cB\u201d). The external resistances were connected between each anode brush and the matching air cathode. The height of the flow cross-sections connected to adjacent chambers was 4.5 cm. Six anode brushes (\u03c6: 2.5 \u00d7 L: 12 cm, Mill-Rose, USA) were inserted into each chamber from the top cover plate of the MES stack (Fig. 1a, b). The upstream to downstream anode brushes in two anode arrays (array \u201cA\u201d or \u201cB\u201d) were record as \u201cA-An1\u201d or \u201cBAn1\u201d and so on (An1 to An6). Two Ag/AgCl reference electrodes (BASi, RE-5B, 210 mV versus a standard hydrogen electrode, SHE) were used to measure anode potentials directly (the cathode potential was obtained by the difference between anode potential thewhole-cell potential). All potentials reported here were versus the Ag/AgCl reference electrode. The anode communities in theMES stackwaswell acclimated to domestic wastewater under the resistance of maximum power density output", " Wastewater stored for more than a week was disposed. Before being pumped into the MES reactor, the influent wastewater tank was temporarily stored in an insulation can with an ice-bath and the influent COD was checked every 12 h. The wastewater was pumped separately into the two anode compartments by a two-channel peristaltic pump (Cole-Parmer, model 7523- 90, Masterflex, Vernon Hills, IL) from the bottom influent tubes on one side and flowed out through the outlet pipes on the opposite side (Fig. 1a). To prevent the effects of dissolved oxygen in influent, all the wastewater tanks were with an airtight cover and the lowoxygen-permeable fluororubber pump tubing (precision pump tubing, Masterflex) was applied. Therefore, the wastewater was in an anaerobic state before pumped into the MES reactor. To make the influent wastewater could recovery to room temperature (20 \u00b1 3 \u00b0C) before being pumped into MES, the pump tubing was warmed up by a water bath (30 \u00b1 1 \u00b0C) controlled by a submersible heating-rod", " During the startup and stabilization period, all the brushes in one anode compartment twisted together by copper wires to make sure the maturation and uniformity of anodes biofilm. Then, the MES was operated in the continuous flow with the HRT regulated from 0.5 to 6 h (0.5, 2.0, 4.0 and 6.0 h). The current densities were adjusted by different resistors (Open circuit voltage condition (OCV), 2000, 1000, 500, 200,100, 60, 30 and 20 \u03a9) connected between each anode brush and the matching air cathode (Fig. 1b). The terminal voltages (U) between each electrode pair were obtained at the same time. The external resistances of anode brushes were changed synchronously. Unless otherwise stated, the value of external resistance in this study represented the external resistances between each brush and the air cathode. Under each resistor value, theMES stackwas first operated for 6 HRTswith the HRT of 0.5 h and then operated for 4 HRTs with the HRT of 2.0, 4.0 and 6.0 h in sequence to stabilize the reactor under each HRT" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000200_s00202-019-00882-x-Figure11-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000200_s00202-019-00882-x-Figure11-1.png", "caption": "Fig. 11 Magnetic field distributions at different angle positions of the magnetic gear rotors", "texts": [ " The values of the magnetic flux density are greater for NdFeB and SmCo magnets. The magnetic flux density in the gear parts has minimum values, when the ferrite magnets are used for the coaxial magnetic gear construction. Maximal flux densities are 1.4\u00a0T, 0.85\u00a0T, 2\u00a0T and 1.95\u00a0T for AlNiCo, ferrite, NdFeB and SmCo magnets, respectively (Figs.\u00a07, 8, 9, 10). These values are above magnets\u2019 residual flux density because of CMG flux focusing. 1 3 Results for the magnetic field distributions at different angle positions of the magnetic gear are shown in Fig.\u00a011. A complete pole movement of 45\u00b0 is presented. Comparative bar graphs for the FEM calculated torques (5) of the magnetic gear rotors are shown in Figs.\u00a012 and 13. It is interesting to note that there are very close results for NdFeB and SmCo torques 1736 to 1534\u00a0Nm, respectively. Results are in accordance with the residual flux density and the magnetic energy density of each material as they are introduced in Table\u00a02. Results for the steady-state torque transmission are obtained and presented in Fig.\u00a011. Results for input and output torques acquired by the model are shown in Table\u00a03. Because of small torque ripples (Fig.\u00a012) due to modulating steel segments, average torque values are presented. Torque ripples are below 0.5% of the steady-state output torque. Input torque is set to be constant. In real application, motor driver ripples must be taken into account. These results are obtained taking into account the electric conductivities for permanent magnet domains only, i.e., the eddy current effects in other domains are neglected, so that the permanent magnet material influences over the CMG efficiency are precisely estimated (Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001304_s00170-020-05395-7-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001304_s00170-020-05395-7-Figure1-1.png", "caption": "Fig. 1 Examples of an infill (left) and overhang (right) structure", "texts": [ " Thus, it is important to understand the heating and cooling effect of the gases being used. Additionally, since wire arc MBAAM is capable of a continuous deposition over an extended period, the process economics plays an important role in selecting an appropriate shielding gas. The recommended selection of shielding gases in wire arc additive manufacturing originated from welding practice, but it does not address the new complexity of sizes and shapes. Rendering such geometrical features as infill layers and overhang structures (Fig. 1), overhang structure requires a weld pool fluidity with low surface tension and a high degree of wetting that is not common in other welding applications. In addition, most of the previous welding studies are based on single bead-on-plate experiments and varieties of weld joints, which is a very different environment than wire arc MBAAM. One of the predominant challenges in wire arc MBAAM is to fabricate dimensionally accurate and equivalent to computeraided design (CAD) geometry. Additionally, most of the past research on wire arc AM has focused on high-cost alloys [6, 7]" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000489_s00170-016-9165-4-Figure8-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000489_s00170-016-9165-4-Figure8-1.png", "caption": "Fig. 8 Finite element (FE) model and the boundary conditions for heat transfer analysis", "texts": [ " The heat generation can be expressed as heat flow rate (Q\u0307) under a thermal load at the boundary: Q \u0307 \u00bc \u03b7 f cv f \u00f05\u00de where fc is the friction force and vf is the velocity. \u03b7 is a factor that specifies how much friction is transformed into heat; this factor varies between 0 and 1. In this study, we assumed that friction energy was completely converted into heat (\u03b7 = 1). Also, the ball array along the LM block was represented by the partial surface of a cylinder for the contact area between the balls and race. Figure 8 shows the finite element model representing the LM guide system (THKHSR-30R); the mesh contains 51,059 solid elements and 180,390 nodes. The material specified by the manufacturer is stainless steel; its thermophysical properties are listed in Table 4. There are four boundary conditions to be considered corresponding to convection, conduction, ambient temperature, and the heat flow rate. The convection region, air exposure under constant temperature, and humidity conditions can be represented by a convective heat transfer coefficient of 20 W/m2 \u00b0C [15]" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002127_j.sna.2020.112301-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002127_j.sna.2020.112301-Figure5-1.png", "caption": "Fig. 5. Experime", "texts": [ " A magnet bar with the same diameter and height is used for comparison (the distance between l and the surface of the magnet bar is 3 mm), and the results are also plotted in Fig. 4. As can be seen from Fig. 4, larger magnetic 4 W. He and C. Qu / Sensors and Actuators A 315 (2020) 112301 F fi s 3 i s g t circuit and load voltages. The experimental output powers are ig. 3. Magnetic force on the suspending part as a function of the displacement. eld variation can be obtained by the proposed suspending magnet tructure when the ME transducers are suitably placed. .2. Experimental setup The experimental setup is shown in Fig. 5. The external mechancal vibrations applied on the prototype are provided by a vibration haker (ESS-015) which is driven by an amplifier (PA-1200). A signal enerator (DG4162) is used to drive the amplifier. An acceleromeer sensor (DL100) and a vibrometer (HY5932) are used to measure the acceleration. The open-circuit and load voltages are measured by a digital storage oscilloscope (UTD2052CEX or TDS2022B). 4. Results and discussion The output performances of the proposed vibration generator are evaluated by theoretical and experimental results" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001844_s12541-020-00333-9-Figure11-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001844_s12541-020-00333-9-Figure11-1.png", "caption": "Fig. 11 Von-Mises stress of flexspline with circular spline and torque", "texts": [ " The simulation results (17)P = 2KT cos bd1( cos t cos \ufffdt ) cos at (18)Ft = 2KT d1( cos t cos \ufffdt ) (19) 1 = d1 sin t 2 cos b (20) 2 = u \u22c5 1 (21) H = \u221a\u221a\u221a\u221a\u221a 1 \u22c5 Ft cos cos \ufffd t b cos2 t \u22c5 1 1\u2212 2 1 E1 + 1\u2212 2 2 E2 \u22c5 u + 1 u \u22c5 2 cos b d1 sin at 1 3 are compared with the calculated values using Hertz contact theory. A comparison between theoretical calculation and simulation is shown in Fig.\u00a010. We can see that the maximum error between simulation and theory-based results is less than 10%, which means the model is quite accurate. Subsequently, we focused on the circular spline and applied a torque of 10\u00a0N\u00a0m. The stress at the root of flexspline suggests the presence of partial load (Fig.\u00a011). We then extracted the stress for the root of the ten meshing teeth in the meshing area. The simulation results are shown as Fig.\u00a012. The results in Fig.\u00a012 indicate that the stress value of 1th\u20134th teeth decreases, but that of 4th\u201310th teeth increases with the meshing process. The 4th tooth is the position with the smallest stress, which is due to the large meshing contact area and without the stress concentration. The value of stress fluctuates at the 9th and 10th teeth, which indicates that there exists the meshing impact at this position" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000720_tmag.2019.2894739-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000720_tmag.2019.2894739-Figure1-1.png", "caption": "Fig. 1. (a) FE mesh, where the additional winding can be seen at the bottom of the slots. (b) Computed flux density distribution in the motor at rated torque and levitation conditions.", "texts": [ " In the simulations, the rotor is considered at a constant speed and a fixed lateral position and the airgap is remeshed at every time step. \u23a1 \u23a2 \u23a2 \u23a2 \u23a3 \u2329 DT mag \u2202H \u2202B Dmag \u232a \u2329 DT mag \u2202H \u2202\u03b5 Dmech \u232a 0 \u2329 DT mech \u2202\u03c3 \u2202B Dmag \u232a \u2329 DT mech \u2202\u03c3 \u2202\u03b5 Dmech \u232a C + 1 t M 0 1 t I \u2212I \u23a4 \u23a5 \u23a5 \u23a5 \u23a6 . (7) The described models are applied to a prototyped BSynRM, the parameters of which are given in Table I. The mesh used in both models and the computed flux density distribution of the motor at the rated torque and levitation are presented in Fig. 1. In this case, the main winding and additional winding fluxes add to each other in the positive y-axis and oppose each other in the negative y-axis. As a result, the stator is more saturated in the upper part than in the lower one. The prototype machine is presented in Fig. 2. AMBs are used on both sides of the shaft, so that the machine can be operated without any active control of the levitation. The prototype machine has been simulated with both models under the same conditions as it has been measured" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001994_lra.2020.3003862-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001994_lra.2020.3003862-Figure1-1.png", "caption": "Fig. 1: Design of 3D maneuverable robotic fish.", "texts": [ " Then, by creating a virtual orifice (circle) around this mobile agent, and maneuvering the robotic fish through this virtual orifice, the robotic fish can get itself into proximity with the mobile agent. Such a maneuver needs to necessarily take into account the velocity vectors of both - the robotic fish as well as the mobile agent, and this is achieved by developing the guidance and control scheme using a relative velocity framework. Preliminary work, including initial design, fabrication, modeling, and testing, was presented in [12]. In this paper, we extend our previous work and focus on modelbased control for passing through a moving orifice. The overall design of the fish is shown in Fig. 1. An IPMC enabled buoyancy control device is used for depth control. Since we mainly focus on investigating the feasibility of the on-board water electrolysis method, we choose a simple propulsion method for 2D maneuvering. A single-joint caudal fin design is adopted because it is a practical, reliable, and straightforward propulsion mechanism for forward swimming. A servo motor is attached at the end of the fish body to actuate the tail to generate flapping motion. Upon being activated, the servo motor drives the tail in a sinusoidal pattern" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000420_17452759.2015.1136868-Figure18-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000420_17452759.2015.1136868-Figure18-1.png", "caption": "Figure 18. The relation chart of gully and layer height.", "texts": [ " Thus, the neck length between the filaments of the adaptive part is between the maximum generated in the smallest thickness and the minimum detected in the largest thickness. And if the neck length is much closer to the maximum value, the part has good mechanical properties. The neck length between adjacent filaments also has a significant impact on surface roughness. Figure 17 is the ultra-depth pictures of a cylinder\u2019s surface with different layer thicknesses (Table 6). The size of the gully between two filaments is related to layer thickness and bonding degree, as shown in Figure 18. The height H and layer thickness have the same trend, but the neck length 2y moves in the opposite direction according to Table 5. So the gully will become more obvious with the increase of layer thickness. But every layer of adaptive slicing has the same accuracy C0 no matter what layer thickness is. According to the core of the adaptive slicing algorithm, a part fabricated with the adaptive algorithm should in principle have the same surface roughness with the one that is built with the smallest layer thickness among the thickness values calculated by the adaptive algorithm" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001398_s12540-020-00779-6-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001398_s12540-020-00779-6-Figure4-1.png", "caption": "Fig. 4 Schematic of welding pool, for calculation of MZ, PMZ and HAZ [4]", "texts": [ "\u00a01) [17]: In this equation, To is the pre-welding sample temperature, T: temperature, Q: input heat, K: thermal conductivity, \u03b1: thermal permeability, R: radial distance from the source and V: is the welding speed. Various references [17] have considered the thermal efficiency of the source on which Q is calculated to be 0.1. To facilitate the solution of the (1) 2 ( T \u2212 To ) KR Q = exp [ \u2212V(R \u2212 x) 2 ] equation, the Rosenthal equation is computed as one-dimensional and summarized as 2 equation: For better understanding of numerical computation, Fig.\u00a04 shows the schematic of computing the lengths of the MZ, PMZ and HAZ zones. According to Fig.\u00a04, the values of each of the zones MZ, PMZ and HAZ are given as follows and corresponding to one of the temperatures mentioned: In the above equations, X2 = TL, X1 = TE, Y2 = TL, Y1 = TS, Z2 = TL, Z1 = TD, X1: The difference between the center of the weld to the beginning of the welding boundary, (2)X = PAVG \u00d7 2 K ( T \u2212 To ) = Q 2 K ( T \u2212 To ) (3)MZ ( x2\u2212x1 ) = TL\u2212 TE (4)PMZ ( y2\u2212 y1 ) = TL\u2212TS (5)HAZ ( z2\u2212z1 ) = TL\u2212TD 1 3 X2: The difference between the welding center distance to the end of the welding pool boundary, Y1: The difference between the welding center and the fusion line, Y2: The difference between the welding center line to the end of the fusion boundaries, Z1: The difference between the weld center line and the fusion line, Z2: The difference between the weld center line to the solution line of precipitates in the structure" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000777_j.energy.2019.04.050-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000777_j.energy.2019.04.050-Figure3-1.png", "caption": "Fig. 3. The schematic of the double-cylinder fuel cell.", "texts": [ "0mg cm 2 Pd/C, Pt/C and Pt40Ru20/C with 5wt% PTFE as a binder were incorporated into MEA-1 to MEA-3, which were listed in Table 2. Three different catalyst contents were loaded into the cathode CL of MEA-3 to MEA-5 to explore the influence of catalyst loading, which was shown in Table 3. The double-cylinder AEM-DGFC is the core component of the electrochemical photobioreactor for consuming the dissolved oxygen, which consists of two chambers, a quaternary ammonia polysulfone anion exchange membrane, two electrodes, two upper and bottom cover, and four bolts. As shown in Fig. 3, the peripheral walls of the two chambers were composed of two coaxial plexiglass cylinders with the thickness of 5mm and the height of 80mm. The inner cylinder with 132 circular holes, which were evenly arranged in 11 rows and 12 lines, was designed to ensure the transmission of the ions between the cathode chamber and the anode chamber. The diameter of each circular hole was 4mm. The QAPS AEM was in the middle of the cathode electrode and anode electrode to form the MEA with an active area of 78" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001812_1077546320912109-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001812_1077546320912109-Figure1-1.png", "caption": "Figure 1. (a) Geometric representation of the roller bearing (b) geometrical deformation between the roller and races.", "texts": [ " The outer race is kept constant, and an NR + 2 degree-of-freedom model (NR = number of rollers) is considered. It is assumed that except for the local surface defects there are no other geometric errors due to slipping, skidding, or skewing motions. The elastic contact deformation between the roller and raceways satisfies the Hertzian contact theory. Nonlinear stiffness and material damping are considered, neglecting the effect of lubrication in the contact areas. A set of nonlinear equations, using theNR + 2 independent coordinates, are formed from Lagrange\u2019s equation. From Figure 1(a) and (b), the geometrical deformations between rollers and races can be expressed as \u03b4ir \u00bc \u00f0r \u00fe Rr \u00fe cl\u00de \u03c7j \u03b4or \u00bc R \u03c1j \u00fe Rr \u00fe cl (1) As the deformations at the contact points are defined using the Hertzian contact theory, so both \u03b4ir and \u03b4or can actively participate when the springs are compressed because of contact or some positive values. Hence, the contact force between the roller and raceways using the contact deformation theory can be written as Fj \u00bc k\u03b410=9j U\u00f0t\u00de \u00bc Kj\u03b4jU\u00f0t\u00de (2) where U\u00f0t\u00dex 1; if \u03b4j < 0 0; if \u03b4j > 0 ", " Using Lagrange\u2019s theory, the contribution of each element of the roller bearing system in the form of kinetic and potential energy written in the form of equations of motion with respect to the generalized coordinates \u03c1j; xir, and yir can be written as follows mr\u20ac\u03c1j \u00fe mr\u03c1j _\u03b8 2 j \u00fe mrg sin \u03b8j \u00fe 1 2 \u2202 h Kir h \u03b410=9 1 ir i \u00fe i \u2202\u03c1j \u00bd\u03b4ir 2\u00fe h Kir h \u03b410=9ir i \u00fe i \u2202\u03c7j \u2202\u03c1j \u00fe 1 2 \u2202 Kor \u03b410=9 1 or \u00fe \u2202\u03c1j \u00bd\u03b4or 2\u00fe \u00fe Kor \u03b410=9or \u00fe \u00fe 10 9 XNR: j\u00bc1 ( Cir\u00f0Kir\u00de\u03b410=9ir\u00fe \u03c7j q\u2202 _\u03c7j\u2202 _\u03c1j ) \u00fe 10 9 XNR: j\u00bc1 Cor\u00f0Kor\u00de\u03b410=9or\u00fe _\u03c1j q \u00bc 0 (6) mir\u20acxir XNR j\u00bc1 h Kir h \u03b410=9ir i \u00fe i \u2202\u03c7j \u2202xir \u00fe 10 9 XNR j\u00bc1 CirKir h \u03b410=9ir i \u00fe _\u03c7j q \u2202 _\u03c7j\u2202 _\u03c7ir \u00bc Fu sin\u03c9t (7) mir\u20acyir \u00femirg XNR j\u00bc1 Kor \u03b410=9or \u00fe \u2202\u03c7j \u2202yir \u00fe 10 9 XNR j\u00bc1 CirKir h \u03b410=9ir i \u00fe _\u03c7j q \u2202 _\u03c7j \u2202 _yir \u00bcW \u00feFu cos\u03c9t (8) From Figure 1, \u03c7j can be calculated as \u03c7j\u00bc \u00f0xor xir\u00de2 \u00fe R2 r \u00fe 2Rr\u00f0xor xir\u00decos \u03b8j \u00fe 2Rr\u00f0yor yir\u00desin \u03b8j \u00fe \u00f0yor yir\u00de2 \u2202\u03c7j \u2202\u03c1j \u00bc \u03c1j \u00fe \u00f0xor xir\u00decos \u03b8j \u00fe \u00f0yor yir\u00desin \u03b8j \u03c7j \u2202\u03c7j \u2202xir \u00bc \u00f0xor xir\u00de \u03c1j cos \u03b8j \u03c7j \u2202\u03c7j \u2202yir \u00bc \u00f0yor yir\u00de \u03c1jsin \u03b8j \u03c7j \u2202 _\u03c7j \u2202 _xir \u00bc \u00f0xir xor\u00de \u03c1j cos \u03b8j \u03c7j \u2202 _\u03c7j \u2202 _yir \u00bc \u00f0yir yor\u00de \u03c1jsin \u03b8j \u03c7j (9) The above set of equations (6)\u2013(8) are, \u00f0NR \u00fe 2\u00de, secondorder nonlinear differential equations. Rotor mass and its unbalanced effect have been added and no other external load interactionwith the system has been considered" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002414_012043-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002414_012043-Figure1-1.png", "caption": "Figure 1. An example of rotor system with a cantilever arrangement of the impeller.", "texts": [ " Futhermore, a special module is required I order to perform the variation of individual object parameters or their sets. The combination of these two capabilities provides the basis for determining the sensitivity of stress-strain characteristics and critical rotational speeds to varying individual parameters. And this, in turn, enables purposeful variation to ensure the criteria of rigidity, strength and stability of the rotor systems. The proposed approach can be considered as an example of a rotor system, which is shown in Figure 1. HERVICON+PUMPS 2020 Journal of Physics: Conference Series 1741 (2021) 012043 IOP Publishing doi:10.1088/1742-6596/1741/1/012043 This rotor system is defined by many parameters. They are: the shape and dimensions of the shaft and the blade disc, in addition, it is the material properties, angular rotation speed, distance between the bearing supports etc. Some of these parameters are specified and some are variable. In different cases, these may be different parameters. We denote the vector of variable parameters by TNppp ,", " , (6) where 0 is a predetermined interval 0 around the operating speed of the system. Thus, the proposed approaches make it possible, on the basis of a single parametric model, to set and solve the problems of analysis of stress-strain state and critical rotation speeds of rotor systems. In addition, the basis for solving the problems of substantiating the rational parameters of these rotor systems by the criteria of strength, stiffness and stability of motion is created. Let us set some nominal parameters of rotor system under study (see Figure 1): shaft length 0L = 0,24 m, shaft diameter 0d = 0,019 m, disk diameter 0D = 0,18 m, length l = 0,12 m, distance between bearing supports 0l = 0,1 m, shaft material is steel (elasticity modulus \u0415 = 2\u221910 11 N/m 2 , Poisson's ratio =0,3); disk with blades made of aluminum alloy (elasticity modulus \u0415 = 7,1\u221910 10 N/\u043c 2 , Poisson's ratio = 0,33); nominal rigidity of bearing supports \u0441 = 1,96\u221910 7 N/m, nominal operating speed 0 = 20 thousands rpm. First, a quasi-static analysis of this rotor loaded by centrifugal forces has been performed with a commercial finite element package ANSYS Workbench" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000951_tro.2019.2936302-Figure11-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000951_tro.2019.2936302-Figure11-1.png", "caption": "Fig. 11. Collision between the simplified model of a dual-arm robot and a space station cabin handrail. In this model, the body is equivalent to the torso of the robot astronaut, joint 1 is equivalent to the shoulder joint of the robot astronaut, joint 2 is equivalent to the elbow joint, and joint 3 is equivalent to the wrist joint. The total robot mass is equal to the mass of the astronaut.", "texts": [ " This section introduces the control of stable parking of a robot astronaut based on human dynamic characteristics. From its physiological and histological aspects, the astronaut parking process is an interactive process involving the arm muscles, ligaments, and bones. In this process, the torso, upper arm, forearm, and hand make up a planar moving structure with six linkages, including two sets of three linkages symmetrically connected to the torso [27]. Referring to this, the robot parking model was simplified, as shown in Fig. 11. The symbols used in this model are defined in Table I. In the table, the subscript i represents the rod number; i = 0 represents the robot body itself, and 1, 2, and 3 indicate the joint numbers of the left (right) arm of the robot. 1) Dynamic Mapping Analysis: The astronaut parking process can be interpreted as a process in which the human body\u2019s momentum gradually decreases to zero due to the pressure of a contact force. During this process, the momentum of the system composed of the human and the space station is conserved", " In general, the closer \u0393 is to 1, the better the humanoid robot parking is; the larger the difference between\u0393 and 1, the worse the humanoid robot parking is. When \u0393 1 or \u0393 1, the robot parking has failed. Four types of parking simulations under microgravity are carried out in this section, i.e., humanoid parking, parking with different velocities, parking with different masses, and parking with XY initial velocity. The simulations are used to verify the effectiveness, robustness, and adaptiveness of the parking control method. A virtual prototype model of the robot astronaut parking simulation is designed, as shown in Fig. 11. The mass of the tester providing the parking data is 80.0 kg; based on the adult body weight ratio [29], the torso mass of the robot astronaut is set to 67.0 kg, the forearm mass is 3.25 kg, the combined mass of the arm and hand is 3.25 kg, and the following assumptions regarding the robot astronaut are made. 1) The robot is bilaterally symmetrical, and the mass of each part is evenly distributed. 2) The robot in the parking remains in the same plane. 3) The wrist joint is a slave joint and does not participate in torque control", " In the virtual prototype, the contact model between the end of the robot arm and the handrail is preset in ADAMS as follows: fcontact = \u23a7 \u23a8 \u23a9 0 q > q0 ka(q0 \u2212 q)ea \u2212 ca ( dq dt ) \u00b7 q\u2212q0 d q \u2264 q0 (30) where q0 is the initial distance between the end of the robot arm and the handrail; q is the actual distance; ka is the stiffness coefficient; ea is the collision index; ca is the maximum damping coefficient; and d is the maximum mutual penetration depth. The contact parameters are determined according to the material properties of the robot astronaut and the handrail and are shown in Table II. To verify the effectiveness of the parking control algorithm, the following three types of parking simulations were designed: humanoid parking experiments, parking experiments with different velocities, and parking experiments with different masses, as shown in Table III. The coordinate system directions are defined with reference to Fig. 11, and the contact force and velocity in the figures are all the X-direction data unless otherwise indicated. 1) Verification Simulation of Humanoid Parking: To verify the effectiveness of the control theory of humanoid robot astronaut parking, this article conducts a parking simulation with a robot astronaut with the same velocity and mass as a human. The initial velocity of the robot astronaut is set to 1.0716 m/s based on the observed initial velocity of the tester before the parking. Based on the parking time of the tester, the simulation time is set to 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001443_rpj-11-2019-0287-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001443_rpj-11-2019-0287-Figure5-1.png", "caption": "Figure 5 Description of different surfaces of the printed samples", "texts": [ " Besides, simulations were performed using ANSYS (3DSIM) software to compare their displacements in the case with and without support structures. The surface roughness (Ra) of the as-built samples was measured by a Form Talysurf-120 profilometer. The surface connected to the substrate and the surface with the support structures could not be measured due to the interference of the support structures. The surface roughness of the remaining four surfaces was measured. The description of these four surfaces is shown in Figure 5. Similarly, each sample was measured three times, and three samples were measured at each building angle with and without the support structures, with the average value taken finally. The Vickers microhardness (HV) of the supporting surface was performedwith SHIMADZUHMV-2Micro Vickers Hardness Tester (Kyoto, Japan). Prior to the measurement, the support structures were removed, and the as-built samples were then polished using a Wirtz Buehler (Dusseldorf, Germany) polishing machine. Each sample was polished to meet the requirements for the hardness measurement", " Without support structures, when the building angle approaches to 90\u00b0, the dimension of the sample accuracy is higher, as revealed in Figures 10(b), 10(d) and 10(f), the dimension becomes more precise as the built angle increases. The dimensional accuracy of building angle at 75\u00b0 in Figure 10 (f) was better than 60\u00b0in Figure 10(d), and 45\u00b0in Figure 10(b) was the worst when there were no support structures. The surface roughness (Ra) of the outer surface, top surface, left surface and right surface (described in Figure 5) was measured. The results of the measurements are demonstrated in Figure 11. The horizontal axis represents different building angles, and the vertical axis represents the measured value of surface roughness. Conspicuously, the surface roughness of these four surfaces has no much difference with or without support structures at the same building angle. The surface roughness is highly depended on the process parameters at the time of printing, such as laser power, scanning speed, spot diameter and so on (Koutiri et al" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001241_j.ijfatigue.2020.105632-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001241_j.ijfatigue.2020.105632-Figure6-1.png", "caption": "Fig. 6. Force analysis diagram and the slight deformation of the planetary gear, (a) force analysis diagram, (b) slight deformation.", "texts": [ " However, because of the interference fit, the friction force on the interference fit surface between the planetary gears and corresponding bearings satisfy Amonton's law. In this study, the limit torque Tj is defined as the maximum torque to ensure that the interference fit surface does not slip. The limit torque met by the interference fit surface is shown in Eq. (5). =T \u03c0r P\u03bcl2j c g 2 (5) where \u03bc is the friction coefficient and lg is the tooth width of the planetary gear. The instantaneous force analysis diagram of the three planetary gears in the planetary gear system is shown in Fig. 6(a). The carrier contrarotates around the axis of the sun gear. FCB is the force on the bearing when the planetary carrier rotates under the action of torqueT . FIG and FSG are the dynamic meshing forces of the planetary gear and the ring gear and the planetary gear and the sun gear, respectively. The superscripts n and t represent the radial and tangential components. Under the action of these three forces, a slight elastic deformation, as shown in Fig. 6(b), will occur on the planetary gear. This slight elastic deformation will reduce the amount of local interference and even separate in some extreme cases, and the limit torqueTj will also reduce. When the meshing error is ignored, the meshing frequencies of the planetary gear, the ring gear, and the sun gear are the same. According to the study of Wei and Zhang [5], the amplitude of the dynamic meshing force between the planetary gear and the inner ring is slightly larger than that between the planetary gear and the sun gear. Therefore, it can be considered that the amplitudes of the dynamic meshing force FIG and FSG are different. Suppose >F FIG SG, an instantaneous torque TGB, as shown in Fig. 6(a), will be generated on the interference fit surface because of the existence of the difference of \u2212F FIG SG. Therefore, according to the magnitude relationship between TGB and the limit torque Tj, the force conditions can be divided into two cases. 2.1. = R\n(67)\nr\n\u03c32 \u03c30 =\n{\n0 r < R\n2 C 1 a 2 / b 2 + 1 \u2212\n\u221a\nC 2 1 + M\n2\nr 2 r > = R\n(68)\n\u00af =\n3 M\u0304\n8 ( 1 \u2212 a 2\nb 2\n)[ 1 R\u0304 2 \u2212 1\nB\n+ ln\n( B\nA\n) + 1\nR\u0304\n2 \u2212 8\n3\n( a\nb\n)2\n+\n5\n3\n] (69)\nhere\n\u00af =\n3 E\u03d5\n4 \u03c30\n( 1 \u2212 a 2\nb 2\n) (70)\n.2.1. Verification of accuracy\nIn this sub-section, the annular membrane is analyzed with he new wrinkling model and the results are compared with he theoretical solutions obtained from Eqs. (63) to ( 65 ), ( 67 ) o (69). The finite element model is shown in Fig. 20 . Its outer adius b and inner radius a (shown in Fig. 19 ) are 1250 mm nd 500 mm, respectively. The thickness t is 1 mm. The Young\u2019s\nodulus E = 10 0 0 MPa and the Poisson\u2019s ratio \u03c5 =1/3. The inner dge is restrained in the radial direction and the circumferential OFs of nodes on the edge are coupled to have the same torsional otation. The outer edge is firstly pretensioned to introduce a", "u c t\nr r a m p M i a s c b p r o a p a\nw fl v t\n6\nm ( 0\n8 As for the numerical examples in Section 6.2 , their geometrical and physical parameters are the same as given in Section 6.2.1 except the quantities assigned specifically.\nniform radial stress \u03c3 0 and then fixed along the radial and ircumferential directions. With a twisting moment M applied to he inner edge, wrinkles arise and their radii increase gradually.\nThe relation between the twisting moment and the torsional otation is given in Fig. 21 . In the figure, the horizontal axis repesents the dimensionless torsional rotation (please see Eq. (70) ) nd the vertical axis denotes the dimensionless twisting mo-\nent (please see Eq. (66) ). Contours of the major and minor\nrincipal stresses and the wrinkling strain corresponding to\n/ (2 \u03c0\u03c30 a 2 t) = 2 . 9 in Fig. 21 are illustrated in Figs. 22\u201324 , where t can be found that the principal stresses and the wrinkling strain re all uniformly distributed along the circumferential direction. It ignifies that the variation of stresses and strains along the radial oordinate may be typical of their distribution over the mem-\nrane. Fig. 25 illustrates the distributions of the major and minor rincipal stresses along the radial direction with R\u0304 = 1 . 2 and 1.6 espectively. The horizontal axis in the figure represents the ratio f the radial coordinate r to the inner radius a and the vertical xis denotes the major or minor principal stresse ( \u03c3 1 or \u03c3 2 ) to the restress \u03c3 0 . Figs. 21 and 25 indicate that the wrinkling model is ccurate and the results agree well with the theoretical solutions.\nFig. 21 also manifests that the major principal stress in the inkled region is much larger and its distribution curve becomes at in the taut area. Conversely, the minor principal stress almost anishes in the wrinkled region while it increases gradually with he radial coordinate in the taut zone.\n.2.2. Influence of the prestress\nThe annular membrane shown in Fig. 19 subjected to a twisting\noment M = 10,0 0 0 N \u2022mm is analyzed 8 with different prestresses \u03c3 = 0 . 1 \u03c4, 0 . 2 \u03c4, 0 . 3 \u03c4, 0 . 4 \u03c4 and 0.5 \u03c4 , where \u03c4 is the uniformly", "o u m\n\u223c s i r\ndistributed shear stress along the circumferential direction at r = a and obtained from the boundary condition with \u03c4 = M\n2 \u03c0a 2 t ).\nThe influences of the prestress on the major and minor principal stresses and the wrinkling strain are illustrated in Fig. 26 (a) and (b), respectively. The vertical axis in Fig. 26 (a) represents the ratio of the principal stress \u03c3 1 or \u03c3 2 to the shear stress \u03c4 and the vertical axis in Fig. 26 (b) denotes the ratio of the wrinkling strain \u025b w to the shear strain \u03b3 ( \u03b3 = \u03c4 G , G is the shear modulus). The solid\nr dashed lines in Fig. 26 (a) represent the major principal stresses nder different prestresses while the discrete symbols denote the\ninor stresses.\nFig. 26 (a) shows that when the prestress increases from 0.1 \u03c4 0.5 \u03c4 , the major and minor principal stresses are augmented lightly in the taut area. Fig. 26 (b) indicates that the prestress s able to decrease the wrinkling strain notably in the wrinkled egion." ] }, { "image_filename": "designv11_5_0001422_j.ijsolstr.2020.07.013-Figure21-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001422_j.ijsolstr.2020.07.013-Figure21-1.png", "caption": "Fig. 21. The test setup.", "texts": [ " 18 and 19 show that the proposed model also simulates the frictional behavior of rough interfaces in the low plastic region, as w \u00bc 4:72, according to Table 2. In the next case, the contact interface vertical force was not constant where the experimental setup reported by (Rajaei and Ahmadian, 2014) consists of a clamped-frictionally supported steel beam. A suspended mass block at frictionally support provides the desired value for preload, as shown in Fig. 20. Roughness characteristics of the contact surfaces shown in Fig. 21 were obtained from surface roughness measurements and are reported in Table 3. The surface roughness parameters are calculated from the measured surface topography (McCool, 1986). According to Table 3, the values of skewness and kurtosis parameters (Rsk and Rku) show that the probability density distribution function of asperity heights is approximately symmetric and is consistent with the Gaussian distribution function (Shi et al., 2019). The roughness characteristics of the contact surfaces are obtained from surface roughness measurements; the calculated radius of asperity summits is R = 201 mm" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002259_j.measurement.2020.108723-Figure16-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002259_j.measurement.2020.108723-Figure16-1.png", "caption": "Fig. 16. Sketch map of plastic deformation of the bearing under radial load considering the Hertzian contact.", "texts": [ " Torque load directly increases the amplitude of stator current; Radial load may cause some misalignment of the rotor when the load is significant high. On the other hand, the motor usually has some inherent eccentricity due to the assembly accuracy, which is inevitable in a real motor. Therefore, it is necessary to discuss about the influence of initial static eccentricity to FHD in the motor. According to Hertzian contact theory, when radial load is added on the bearing, the bearing inner ring, outer ring and rolling elements bring forth plastic deformation in the contact area as shown in Fig. 16 [51]. This may lead to a static eccentricity along with the radial force direction. The relative displacement between inner ring and outer ring can be expressed as: \u03b4 = \u03b4ib + \u03b4ob (32) where \u03b4ib is the displacement between inner ring and balls, \u03b4ob is the displacement between outer ring and balls: \u03b4 = e\u03b4 \u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305 ( \u2211 \u03c1)Q23 \u221a (33) where e\u03b4 is Hertzian contact coefficient, \u2211 \u03c1 is the sum of main curvature, Q is the load on the ball and it can be expressed as the following formula in deep groove ball bearing: Q = 5Fr Zcos\u03b1 (34) where Fr is the radial load, Z is the number of rolling elements, \u03b1 is initial contact angle" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000063_rnc.4690-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000063_rnc.4690-Figure3-1.png", "caption": "FIGURE 3 The coordinates frame of robot manipulator with 3-DOF", "texts": [ " (ii) Due to the sliding model surface s = 0, ||z2|| \u2264 \ud835\udefc|\ud835\udf0c| can be ensured. Therefore, on the approximation domain of U0 = {X|||X|| \u2264 |\ud835\udf0c|\ud835\udefc}, the RBFNNs, as shown in Figure 1, is utilized to approximate the unknown model of robot manipulators. Accordingly, we have the conclusion that the tracking errors and the states of the closed-loop system can achieve GUUB whatever go outside or stay inside the approximation domain of RBFNNs by the above two steps. In this simulation studies section, the coordinate system of robot manipulator with 3-DOF is shown in Figure 3 to illustrate the RBFNNs adaptive controller. The dynamical robot manipulator system can be written as (1) with the parameters in Table 1. M(q) = [ 11 12 13 21 22 23 31 32 33 ] , F\u0303 = [ \ud835\udc5311 \ud835\udc5321 \ud835\udc5331 ] ,C(q, .q) = [ 11 12 13 21 22 23 31 32 33 ] ,G(q) = [ 11 21 31 ] ,with 11 = m3q2 3 sin(q2 2) + m3r2 1 + m2r2 1 + 1 4 m1r2 1. 12 = m3q3r1 cos(q2), 13 = m3r1 sin(q2), 21 = M12, 22 = m3q2 3 + 1 4 m2r2 2, 23 = 32 = 0, 31 = 13, 33 = m3. 11 = m3q2 3 sin(q2) cos(q2) .q2 + m3q2 3 sin(q2 2) .q3, 12 = m3q2 3 sin(q2) cos(q2) " ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000640_humanoids.2016.7803433-Figure9-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000640_humanoids.2016.7803433-Figure9-1.png", "caption": "Fig. 9 End-effector trajectory", "texts": [ " Then, after being placed on the ground, the body leaves the ground and proceeds forward. By repeating these two phases, the robot moves forward. In terms of the trajectory of the end of the limb, when it propels forward, the collision between the limb and ground should be avoided as much as possible. Consequently, to move the limb end forward and higher than the ground, we provide a trajectory for it. Similar to the limb end, the body likewise first rises, moves forward, and contacts the ground (Fig. 9). By repeating these motions, the robot is propelled forward. The movement in terms of the limb end is similar to rotating along the trajectory. We can thus handle the crawling motion of the four-limbed robot as the motion of a wheeled robot. The difference between these two motions is the timing of contacting the ground. For wheeled robots, the wheels are always contacting the ground, whereas for legged robots, the legs are in a swinging phase and contacting phase; the legs do not contact the ground at all times" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002142_b978-0-08-102965-7.00006-0-Figure6.6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002142_b978-0-08-102965-7.00006-0-Figure6.6-1.png", "caption": "Figure 6.6 Schematic of scanning paths: (A) linear path and (B) fractal path [56].", "texts": [ " Consequently, it will affect the mechanical strength due to the poor geometric relationship between the sintered powder particles [42]. Therefore, optimum scan path generation is extremely important for fabrication of a fully dense and dimensionally accurate part using SLS. There are certain conditions on which the scan pattern depends, foremost of which is that the path should not intersect itself, which means that during laser scanning of a given layer, every powder particle must be sintered only once. The scanning path should be simple and convenient to access real-time control. Fig. 6.6 shows an illustration of different types of scanning paths followed in SLS [56]. Another important scan parameter is hatch spacing. Hatch space is also known as the scan line spacing or the distance between two single scan lines, as shown in Fig. 6.7 [59]. In the SLS process, scanning is done line by line, beginning from the start of the uppermost or lowermost line, continuing till the end of the line and then moving on to the adjacent line in the pattern illustrated. It is also observed that the end portions develop a denser structure compared to the other portion on a single line" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002536_j.apor.2021.102597-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002536_j.apor.2021.102597-Figure4-1.png", "caption": "Fig. 4. The 3D model of a torpedo-shaped AUV that refers to REMUS 100 (Lab, 2017) and suffers additive actuator faults f = [fX , fY , fZ, fK, fM, fN]T where the forces (fX , fY , fZ) are parallel to the axes of {b} and torques (fK, fM, fN) follow the right-hand screw rule along the axes of {b}.", "texts": [ " (25) equals u[n+1] = B\u2212 1\u03c4\u2605 \u2212 B\u2212 1f ( uj[n] ) + ej[n] 1 + bT j k[n] B\u2212 1k[n]. (27) Since f(uj[n]) is fixed and ej[n] is 0 if uj[n] has converged, the convergence of the rest elements in u[n+1] is affirmative. Remark 12. Previous analysis was given by treating the jth equation of (21) as a synthesized design of FTC. The convergence of (24) could also be proved with a similar way. The analysis is obvious and thus omitted. Simulations on a torpedo-shaped AUV were taken to illustrate the effectiveness of the proposed methods. The torpedo-shaped AUV shown in Fig. 4 contains propeller blades that rotate with angular velocity \u03c9p and four rudders that provide attitude control torques under angles \u03b4u, \u03b4l, and \u03b4e. The global coordinate {n} and local body-fixed reference frame {b} are both given. Unexpected factors could cause actuator faults, such as the deformations shown in Fig. 4, where additive forces and torques would be produced. The dynamic model of an AUV given in (Fossen, 1994) is M\u03bd\u0307 + C(\u03bd)\u03bd + D(\u03bd)\u03bd + g(\u03b7) = \u03c4, (28) where \u03bd \u2208 R6 represents velocities in {b}; \u03b7 \u2208 R6 contains generalized positions and heading angles in {n}, which satisfies \u03b7\u0307 = J(\u03b7)\u03bd with J(\u03b7) being a revolution matrix between {n} and {b}; M, C(\u03bd), D(\u03bd) \u2208 R6\u00d76 are inertia, Coriolis, and damping matrices, respectively; g(\u03b7) includes forces and torques of gravity and buoyancy to the center ob; \u03c4 \u2208 R6 consists of forces [\u03c4X, \u03c4Y , \u03c4Z] T and torques [\u03c4K, \u03c4M, \u03c4N] T along body axes" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002210_j.robot.2020.103670-Figure11-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002210_j.robot.2020.103670-Figure11-1.png", "caption": "Fig. 11. Action of the stopper. (a) Robot moving forward, (b) Forearm links in contact with the wheelchair stopper, (c) Lifting the wheels of the robot, (d) Preventing the robot from falling.", "texts": [ " The wheelchair has stopper under the push handle mechanism that is composed of ront and rear bars (Fig. 9). During the robot climbing process, the anipulators are spread, and the right and left manipulators are nserted into the stopper (Fig. 10(b)). As described above, the robot\u2019s dual manipulators have touch ensors that permit it to detect contact with the stopper (Fig. 6). his allows the robot to push against the wheelchair\u2019s front bars using the manipulator forearm links in order to lift its front wheels (Fig. 11(a)\u2013(c); see Section 3). The rear bars prevent the robot from tipping over backward when the robot angle inclines and its mass position shifts behind the contact point between its middle wheels and the ground (Fig. 11(d)). Fig. 12 shows a diagram of the system configuration. Here, the solid lines show the wired network, while the dotted lines show the wireless network. In this study, the robot and wheelchair move automatically to perform step climbing. Thus, although the caregiver is able to monitor the video imagery, he or she does not need to operate the system. An accelerometer system (MMA7361, Cixi Bouri Technology Co., Ltd.) is installed on the robot to detect its incline, while an ultrasonic sensor system (HC-SR04, SainSmart) installed on the front part of the robot (Fig", " [Stage 3] <9> The robot folds its front and rear wheel mechanism (Fig. 4) and spreads its manipulators (Fig. 10(a) and (b)). The touch sensors on the forearm links (which are covered with bumpers) detect contact with the wheelchair (Fig. 6), after which the manipulator forearm links are inserted into the stopper (between the front bars and rear bars)(Fig. 9). The wheelchair then stops, and the robot moves forward until the manipulator forearm links come into contact with the front bars of the wheelchair stopper (Fig. 11(a)\u2013(c)). <10> If the system detects the step height, it can then determine the angle needed to place the robot\u2019s front wheels on the step (See Section 4). The robot continues to drive forward, pushing against the wheelchair using the forearm links, until the robot\u2019s front wheels are lifted off the surface and its accelerometer system detects the inclination needed to place its front wheels on the step. <11> As the robot\u2019s tilt increases, the robot center of mass shifts behind the contact point between the robot middle wheels f and the ground and the robot begins to tip backward. However, the manipulator forearm links come into contact with the rear bars of wheelchair stopper, which limits the extent of rotation (Fig. 11(d)). As a result, the wheelchair prevents the robot from tipping over. <12> The wheelchair and the robot then move forward, and the front wheels of the robot are placed on the upper level of the step. [Stage 4] <13> The wheelchair and the robot continue to move forard. <14> Next, the robot\u2019s middle wheels come into contact ith the step as the wheelchair pulls the robot forward. At this oint, the robot\u2019s rear wheels are lowered and provide additional upport as the robot\u2019s middle wheels begin to climb the step" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000151_s10999-019-09479-5-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000151_s10999-019-09479-5-Figure3-1.png", "caption": "Fig. 3 The diagram of a lubricated revolute joint", "texts": [ " However, with regard to the lubricated revolute joint in Fig. 2, the centre of the journal is almost free to move relative to the centre of the bearing in a circle with the radius as the clearance size. A force constraint replaces the kinematic constraint in lubricated revolute joint. The interaction forces depend on not only the kinematic interaction of the journal in the bearing but also the hydrodynamic behaviours of the lubricant oil. The eccentricity vector that indicates the relative position of the centres of a journal and bearing in Fig. 3, is evaluated as e \u00bc rPj rPi \u00bc \u00f0rj \u00fe Tjr 0P j \u00de \u00f0ri \u00fe Tir 0P i \u00de \u00f04\u00de where rn (n \u00bc i; j) is the global coordinate vector of the centre of mass in XOY, while r 0P n is the local coordinate vector of the centre Pk of the lubricated revolute joint in okxkyk. Tn is a matrix. The eccentricity direction is necessary to describe the force direction of the lubricated joint, and it can be expressed as r \u00bc e ej j \u00bc effiffiffiffiffiffiffi eTe p \u00f05\u00de The eccentricity ratio e is always calculated by dividing the vector e by the size of the radial clearance C \u00f0C \u00bc Rj Ri\u00de. The rate of change with respect to time of the eccentricity ratio is obtained as _e \u00bc _ej j C \u00f06\u00de The angle of the vector e in the global coordinate system XOY, which is shown in Fig. 3, is given by a \u00bc arctan ey ex \u00f07\u00de Then, the rate of change with respect to time of the angle of the eccentricity vector can be written as _a \u00bc _eyex _exey e2 \u00f08\u00de Lubricated joints are widely used in machine tools, industry robots, aircraft engines and rolling stock, and among others, to reduce the undesirable friction and wear phenomena. For lubricated joints in mechanical systems, the lubricant film is subject to dynamic loads and provides hydrodynamic force to prevent surface contact between the journal and bearing (Bannwart et al. 2010). The hydrodynamic force model should certainly contain the density, viscosity and flow rate of the liquid, in addition to all parameters of journals and bearings (Flores et al. 2004). Hydrodynamic lubrication is based on the lubricating film, which is automatically generated by using a relative movement between the friction surface. The schematic diagram of lubrication is shown in Fig. 3. For a lubricated revolute joint with dynamic loads, the dynamic pressure in isothermal Reynolds\u2019 equation is typically expressed as Pinkus and Sternlicht (1961) o ox h3 l op ox \u00fe o oz h3 l op oz \u00bc 6Rix oh ox \u00fe 12 oh ot \u00f09\u00de where x and z are the coordinates of the radial direction and length direction, respectively. h, Ri, x, p and l are the film thickness, the radius of the bearing, the rotation speed of the journal relative to bearing, the dynamic pressure and the dynamic viscosity of the lubricant, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001584_s40815-020-00940-8-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001584_s40815-020-00940-8-Figure3-1.png", "caption": "Fig. 3 The schematic of the quadrotor along with the assigned coordinate systems", "texts": [ " If more than one expert are consulted, the average weight will be considered. In other words, Eqs. (13) and (14) are used to compute the weights for each expert, and then the average value of weights is inserted into Eq. 11. In conventional FCMs, weights ~w1;3 and ~w2;3 are determined in advance, and they do not change throughout the procedure. But in our approach, new weights are obtained at each sample time. The coordinates of the quadrotor model are described with a body-fixed frame FB \u00bc fxB; yB; zBg and a ground-fixed inertia frame FI \u00bc fxI ; yI ; zIg. In Fig. 3, P \u00bc \u00f0x; y; z\u00deT represents the position of the quadrotor\u2019s center of mass, and V is its linear velocity in the inertia frame. The angular velocity of quadrotor is indicated by X \u00bc \u00f0p; q; r\u00deT , its total mass by m and the acceleration of gravity by g. The governing equations of the quadrotor are as follows [5]: _P m _V \u00bc V FRaz \u00fe mgaz \u00fe F d ; \u00f014\u00de _R J _X \u00bc Rsk\u00f0X\u00de sk\u00f0X\u00deJX\u00feM \u00feMd : \u00f015\u00de The orientation of quadrotor in the body frame relative to the ground-fixed (inertial) frame is specified by the rotational matrix R FB " ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000129_tia.2019.2946525-Figure14-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000129_tia.2019.2946525-Figure14-1.png", "caption": "Fig. 14. SCC with remaining turns in the same phase winding shorted.", "texts": [ " 1) Since this method is based on calculation of highfrequency winding input impedance in comparison to a normal value, operational factors like different steady loads and unbalances will have little impact on the effectiveness of this approach; 2) Other common winding insulation failures like winding open circuit or terminal short circuit feature either no current feedback or fairly high current response, which will not cause ambiguities when using the proposed approach. IV. POSTFAULT REMEDIATION METHOD The primary approach for fault mitigation discussed in existing literature is to short the remaining turns of the same phase winding. Injecting a negative d-axis current for the purpose of fault mitigation will also be evaluated in this section. Figs. 14 and 15 provide a comparison of the faultmitigating methods. Shorting the remaining turns is applied in Fig. 14 while applying negative d-axis current (current amplitude equals machine characteristic current ic) is used in Fig. 15 (both \u201csine current\u201d and \u201csine + third-harmonic\u201d are discussed for the latter method). Fig. 16 is a demonstration of the angle and amplitude of fundamental and third-harmonic components of the negative d-axis current applied in Fig. 15. It should be noted that the negative d-axis lags the q-axis by 90\u00b0 for the 1st component and 30\u00b0 for the 3rd harmonic, resulting in a sharp-top synthetized current waveform" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000403_0954405416640171-Figure8-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000403_0954405416640171-Figure8-1.png", "caption": "Figure 8. Optimal selection of initial data points.", "texts": [ " Extraction of optimized tooth data points According to tool trace distribution on the CNCmilling model, the intersection points constitute the boundary curves of the tooth flank which can be extracted as the initial data points. With the application of the detailed extraction method, as represented in Figure 7, whole tooth profile part including the work tooth profile, fillet and tooth root portion is selected as an object to make coordinate information of the corresponding selected points. In order to get optimal selection of data points on cubic NURBS curves, it employs a process method as shown in Figure 8. As the fillet at Middle East Technical Univ on May 21, 2016pib.sagepub.comDownloaded from process on the tooth profile curve is a key part and its knife pattern becomes denser than other areas, the extraction of the data points need to be larger. Here, the tooth profile is assumed as v-line and the fillet curve is assumed as the u-line. Two lines have same quantity of data points. If the data point Pn in the initial sampling can constitute a curve equation F(u), the corresponding parameter of each point is u and can be obtained by the following formula ui = PiPi+1j j P0P1j j+ P1P2j j+ + Pn 1Pnj j 2 0, 1\u00bd (i=1, 2, " ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002298_tec.2020.3045063-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002298_tec.2020.3045063-Figure2-1.png", "caption": "Fig. 2. Part of the calculation model (air) for numerical ventilation calculation.", "texts": [ " The standard k\u2013\u03b5 model uses the following transport equations for k and \u03b5: t t k \u03c1k \u03bc \u03c1k (k) 2\u03bc \u03c1\u03b5, \u03c3 U ij ijS S t 2 t 1\u03b5 t 2\u03b5 \u03b5 \u03c1\u03b5 \u03bc \u03b5 \u03b5 \u03c1\u03b5 (k) 2\u03bc \u03c1 , \u03c3 k k U ij ijC S S C t where 2 t \u03bc k \u03bc \u03c1 , \u03b5 C is turbulent viscosity (in m2/s); ijS are components of deformation rate (in m/s); \u03bc k \u03b5 1\u03b5 2\u03b50.09,\u03c3 1.00,\u03c3 1.30, 1.44, 1.92C C C \u2014 adjustable constants for standard k\u2013\u03b5 model. The computational domain parameters: rotor length is 1900 mm, radial ducts number \u2013 24, the ratio of a sub-slot section to the total area of radial ducts section \u2013 0.43; rotation speeds 485, 1000, 1500 rpm. Figure 2 shows a part of the numerical model. The air flow considered is stationary, isothermal, with constant density. Simulation was done using the Reynolds equations based on polyhedral mesh by the finite-volume method in ANSYS Workbench [19]. The Reynolds stresses are defined using k\u2212\u03b5 turbulence model. Rotation was set by the MRF method (Multiple reference frame). This method directly reflects the flow character at the edge of the divide between the rotating and the stationary zones. MRF approximation in the steady state allows separate cells to rotate and move at different speeds" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002210_j.robot.2020.103670-Figure18-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002210_j.robot.2020.103670-Figure18-1.png", "caption": "Fig. 18. Model showing the lifting of the robot\u2019s front wheels.", "texts": [ " However, while contacts between the robot\u2019s front wheels nd the step do not always cause a step-climbing failure, there s a limit to the extent of an acceptable wheelbase transformaion. (The minimum value of WBf is 190 mm. See Appendix A, able A.1.) In addition, the robot cannot begin lifting itself if it is significant distance from the step, because it is connected to the heelchair. In this section, we show the results of a theoretical analysis hat clarifies the distance from a step that is necessary to prevent ailure when lifting the robot\u2019s front wheels during step climbing. .1. Moving distance of basic coordinate system for the robot inclines In Fig. 18, O is the contact position between the robot\u2019s middle wheels and the road surface, which is also the basic coordinate system for the vehicles before lifting the robot\u2019s front wheels, 0\u03a3B. O moves to O\u2032 as the robot inclines, so the basic coordinate ystem for the vehicles is expressed as \u03a3B. pi (i = 1\u20135) are the joints (p1: axes of the robot\u2019s middle heels, p2: shoulder joints, p3: elbow joints, p4: robot hands location where the hands hold the wheelchair push handle), p5: xes of the wheelchair\u2019s rear wheels, and 0pi (i = 1\u20135) when the obot incline is zero", " 17 and 18, it can be seen that when \u03a3i is parallel o \u03a3 in the local coordinate system, Bp = [0 R ] T , 1p = B 1 B 2 [lLB hLB] T , 2p3 = [l2 0]T , 3p4 = [l4C 0]T , and 4p5 = [lLA \u2212 hLA] T . Moreover, \u03c6i is the angle of \u03a3i formed by \u03a3i\u22121. In Stage 3, the wheelchair stops on the step and the robot continues driving forward as the system lifts that robot\u2019s front wheels (see Section 3). Since the robot\u2019s angle increases as it moves forward, the position of the basic coordinate system for the vehicles moves from 0\u03a3B to \u03a3B (Fig. 18), and the robot\u2019s center of mass shifts behind the contact point between the robot\u2019s middle wheels and the ground as the angle of the robot increases. During this process, the positions of p1, p2 and p3 are changed, but the positions of p4 and p5 are maintained because the wheelchair remains stopped while the robot\u2019s front wheels are lifted. The homogeneous transformation matrices BT 4 are given below. BT 4 : [cos\u03c61234 \u2212 sin\u03c61234 X4 sin\u03c61234 cos\u03c61234 Y4 0 0 1 ] (1) Here, \u03c61234 = \u03a34 i=1\u03c6i is the incline of the wheelchair against the ground. Therefore, \u03c61234 = 0 in Stage 3. The position vector of push handle of the wheelchair before lifting the robot\u2019s front wheels, Bp = [ X Y ] T in \u03a3 , is as 0 4 0 4 0 4 0 B w b o T h ( r w \u2206 S \u2206 B 0 \u03c1 t B H s given below (Fig. 18). B 0p4 : [ 0X4 0Y4 ] = [ cos 0\u03c61 \u2212 sin 0\u03c61 sin 0\u03c61 cos 0\u03c61 ][ lLB hLB ] + l2 [ cos 0\u03c612 sin 0\u03c612 ] + l4C [ cos 0\u03c6123 sin 0\u03c6123 ] + [ 0 RB ] (2) here 0\u03c612 = 0\u03c61 +0 \u03c62, and 0\u03c6123 = 0\u03c61 +0 \u03c62 +0 \u03c63. When the robot center of mass shifts behind the contact point etween its middle wheels and the ground, the angles of the links f the manipulators, 0\u03c61, 0\u03c612, and 0\u03c6123, shift to \u03c61, \u03c612, and \u03c6123, respectively. Thus, the position vector of hand rim of the wheelchair, Bp4 = [X4 Y4] T in \u03a3B, is as given below", " The wheelchair emains stopped until the system finishes lifting the robot\u2019s front heels. p4 does not change. Thus, x = \u2212(X4 \u2212 0X4) (4) ubstituting (2) and (3) for (4), x = l4C (cos 0\u03c6123 \u2212 cos\u03c6123) + l2(cos 0\u03c612 \u2212 cos\u03c612) +lLB(1 \u2212 cos\u03c61) + hLB sin\u03c61 (5) The position vector of \u03a3B in 0\u03a3B is described below (Fig. 19). p\u2206x = [\u2206x 0]T (6) 4.2. Robot front wheel height needed to climb a step The angle of the robot necessary to place the front wheels on a step is clarified in Stage 3. In Fig. 20, q1 is the position of the robot\u2019s middle wheel axes (Here, q1 = p1, Fig. 18), q2 is the position of the robot\u2019s front wheel axes, and q3 is the tread of the robot\u2019s front wheels. \u03a3I and \u03a3II are the local coordinate system. (Here, \u03a3I = \u03a31, Fig. 18.) \u03c11 and \u03c12 are the angles of \u03a3I and \u03a3II formed by \u03a3B and \u03a3I , respectively. From Figs. 18 and 20, \u03c11 = \u03c61 (7) The position vectors for the joints are expressed as Bqj = [xj yj]T (j = 1 \u2212 3). When \u03a3 and \u03a3 parallel \u03a3 , Bq = [0 R ] T , I II B 1 B 1q2 = [WBf \u2212 RB + rBf ]T and 2q3 = [0 rBf ]T (Figs. 17 and 20). 12 = \u03c1123 = \u03c11 + \u03c12 because \u03c13 = 0. Thus, the homogeneous ransformation matrices Bt2 and Bt3 are as given below. t2 : [cos \u03c112 \u2212 sin \u03c112 x2 sin \u03c112 cos \u03c112 y2 0 0 1 ] (8) Bt3 : [cos \u03c112 \u2212 sin \u03c112 x3 sin \u03c112 cos \u03c112 y3 0 0 1 ] (9) Here, Bq2 : [ x2 y2 ] = [ cos \u03c11 \u2212 sin \u03c11 sin \u03c11 cos \u03c11 ][ WBf \u2212(RB \u2212 rBf ) ] + [ 0 RB ] (10) Bq3 : [ x3 y3 ] = [ cos \u03c11 \u2212 sin \u03c11 sin \u03c11 cos \u03c11 ][ WBf \u2212(RB \u2212 rBf ) ] +rBf [ cos \u03c112 sin \u03c112 ] + [ 0 RB ] (11) When the tread of the robot\u2019s front wheels, q3, is at the bottom of the wheel, \u03c112 = \u221290 deg (Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001576_s11012-020-01259-2-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001576_s11012-020-01259-2-Figure1-1.png", "caption": "Fig. 1 a Schematic of porous journal bearing system. b Another view of porous journal bearing system", "texts": [ " Keywords Wear defect Non-Newtonian lubricant Finite element method Double layer porous journal bearing system List of symbols c Radial clearance (mm) Cij Fluid film damping coefficients (i; j \u00bc x; z) (N s mm-1) D Diameter of journal (mm) e Journal eccentricity (mm) Fxo;Fzo Components of fluid film reaction (N) Fo Fluid film reaction (N) g Acceleration due to gravity (mm s-2) h Nominal fluid film thickness (mm) Dh Change in bearing geometry due to wear (mm) H Wall thickness of porous bearing, (mm), (Fig. 1b) H1 Thickness of inner layer of porous bearing (mm), (Fig. 1b) L Length of bearing (mm) k1 Permeability of inner layer of porous material (mm2) k2 Permeability of outer layer porous material (mm2) Mj;Mc Journal mass, critical mass (kg) n Power-law index Q Lubricant flow of oil (mm3 s-1) p Pressure (N mm -2) p1 Pressure in the inner layer of porous matrix (N mm -2) p2 Pressure in the outer layer of porous matrix (N mm -2) ps Supply pressure (N mm -2) Rj Radius of journal (mm) A. Singh (&) S. C. Sharma Department of Mechanical and Industrial Engineering, Tribology Laboratory, Indian Institute of Technology, Roorkee, Roorkee 247667, India e-mail: anilsingh0191@gmail", " Thus, for the smooth operation of bearings it is desirable to have an improved stability threshold speed margin xth\u00f0 \u00de for the given geometric and operating conditions. Therefore, the author(s) gets motivated to explore the combined influence of non-Newtonian behaviour of lubricant and influence of wear on the performance of hybrid DLPJBs. In this article, the performance of hybrid DLPJBs considering the effects of shear thinning/shear thickening and influence of wear has been studied. The results presented in this paper are expected to be quite useful to the bearing designers. Figure 1a, b illustrates the schematic of hybrid DLPJBS configuration. In this study, the material of porous bearing is assumed to be isotropic and homogeneous. The modified Reynolds equation governing the flow of non-Newtonian lubricant in clearance space of a hybrid DLPJBS considering steady state, incompressible, and laminar flow is given as [8, 10, 13, 30, 31]: o ox h3 12l op ox \u00fe o oy h3 12l op oy \u00bc U 2 oh ox \u00fe oh ot \u00fe k1 l op1 oz Z\u00bc0 \u00f01\u00de where z \u00bc 0 indicates the porous surface and fluid film region interface. The pressure distribution p1\u00f0 \u00de and p2\u00f0 \u00de in the inner and outer layer of porous surface (Fig. 1a, b) is represented by the Laplace equation of the form [8, 10, 13]: o2p1 ox2 \u00fe o2p1 oy2 \u00fe o2p1 oz2 \u00bc 0 \u00f02a\u00de o2p2 ox2 \u00fe o2p2 oy2 \u00fe o2p2 oz2 \u00bc 0 \u00f02b\u00de Integrating Eq. (2a) with respect to \u2018z\u2019 within the limits H1 to 0, yields op1 oz Z\u00bc0 \u00bc Z0 H1 o2p1 ox2 \u00fe o2p1 oy2 dz\u00fe op1 oz Z\u00bc H1 \u00f03\u00de Since, the normal flow between the boundary of two porous layer is equal, i.e. k1 op1 oz Z\u00bc H1 \u00bc k2 op2 oz Z\u00bc H1 \u00f04\u00de Therefore, op1 oz Z\u00bc0 \u00bc Z 0 H1 o2p1 ox2 \u00fe o2p1 oy2 dz \u00fe k2 k1 op2 oz Z\u00bc H1 : \u00f05\u00de Further Integrating Eq", " (13) reduces to the following form: a \u00bc x Rj ; b \u00bc y Rj ; p \u00bc p ps ; X \u00bc xJ lrR 2 j c2Ps ! ; U \u00bc xjRj; h \u00bc h c ; t \u00bc t c2Ps lrR 2 j ! ; l \u00bc l lr ; W \u00bc k1H c3 ; F0 \u00bc F0 lr h ; F1 \u00bc F1 lr h2 ; F2 \u00bc F2 lr h3 o oa F2 h 3 \u00fe F0W Kr 1 c\u00f0 \u00de \u00fe c\u00f0 \u00de n o o p oa \u00fe o ob F2 h 3 \u00fe F0W Kr 1 c\u00f0 \u00de \u00fe c\u00f0 \u00de n o o p ob \u00bc X o oa 1 F1 F0 h \u00fe o h o t \u00f014\u00de where F0; F1and F2 are known as cross-film viscosity integrals and is given as: F0 \u00bc Z 1 0 1 l d z; F1 \u00bc Z 1 0 z l d z and F2 \u00bc Z 1 0 z2 l z l F1 F0 d z: 2.1 Fluid film thickness Figure 1a shows the worn out bearing geometry. Dufrane et al. [15] based on their study of worn out bearings in industries, found that the footprint created by the shaft is almost exactly symmetrical at the bottom of the bearing. The model given by Dufrane et al. [15] for the change in bush geometry (D h) due to wear defect is expressed as [14, 20, 21]: D h \u00bc dw 1 sina; for ab a ae \u00f015a\u00de D h \u00bc 0; for a\\ab or a[ ae \u00f015b\u00de The ab and ae (as shown in Fig. 1a) angles indicate the start and end of footprint for the bearing, respectively. Where dw is wear depth parameter of worn zone. The start angle ab and end angle ae of the worn zone are calculated by considering D h \u00bc 0 at that location and is obtained as; sina \u00bc dw 1; a \u00bc sin 1 dw 1 \u00f015c\u00de For a given value of dw, Eq. (15c) provides a negative value of sina, which may be considered to occur in both the 3rd and 4th quadrants as represented by ab and ae, respectively. The expression for film thickness ( h) in hybrid DLPJBS including the change in bush geometry (D h) due to wear defect, in dimensionless form is given as [14, 20]: h \u00bc 1 Xjcosa Zjsina\u00fe D h \u00f015d\u00de 2" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002086_tia.2020.3015693-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002086_tia.2020.3015693-Figure4-1.png", "caption": "Fig. 4. (a) 2-D sliced rotor of a 3-step skewed VF IPMSM; (b) 2-D sliced rotor of a 5-step skewed VF IPMSM.", "texts": [ " Cogging torque does not contribute to the net effective torque. Different cogging torque minimization techniques were already mentioned in [10]-[14]. The cogging torque of the 27 slot 6 pole VF IPMSM at 100% magnetization is simulated using FEA and minimized by skewing the PM-pole for several steps. The optimum skewing angle to minimize the cogging torque is found analytically. The VF IPMSM has a magnetic symmetry of mechanical 120\u00ba. Therefore a mechanical 120\u2070 sliced rotor of the VF IPMSM with 3-step and 5-step skewed PM-pole is shown in Fig. 4(a) and 4(b), respectively. Cogging torque was derived in [11], [13] by using the Fourier series expansion based on the relative air-gap permeance function and the flux density in an equivalent slotless PMSM. The optimum skewing angles to eliminate the cogging torque were mentioned in [11], [13] and given by equation (1). 2 [ 1,2,3........] skew L k k N (1) \u03b8skew is optimum skewing angle to eliminate the cogging torque theoretically. NL is the least common multiple of the number of slots and the number of poles and also the fundamental period of the cogging torque" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002484_lra.2021.3061361-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002484_lra.2021.3061361-Figure1-1.png", "caption": "Fig. 1. The 3D model of the 4-DOF stationary exoskeleton with uncertain inertia and loads from the user is shown in (a). Body 1 and Body 2 are attached to link 3 and link 4, respectively; (b) shows the 3D model in the MATLAB environment, where the load tetrahedrons are used to identify the COM of the unknown bodies.", "texts": [ " (14) through simulations of a stationary upper limb exoskeleton and a wearable wrist exoskeleton [9] carried out in MATLAB1 [27]. For all simulations, the reference r and disturbance w are selected as bounded periodic and quasiperiodic trajectories with multiple harmonic components. The simulation sampling rate is 500 Hz, and the control input update rate is 250 Hz. The gravitational acceleration cg is along the \u2212z axis of the global frame. The 4-DOF stationary upper limb exoskeleton dynamical system is presented in Fig. 1. The conceptual design adopts a structure similar to the EXO-UL8 exoskeleton [5] without the forearm and wrist actuations. The joint of the ith link is labeled as \u03b8i, which is directly actuated by a motor torque input. When the user equips the exoskeleton, inertia and load uncertainties are introduced from the upper arm (Body 1) and forearm (Body 2) to Link 3 and Link 4, respectively. For the ith load, the unknown parameter pi \u2208 R7 can be written as pi = [ mi \u03c6i cm ]T (20) where m is the mass, \u03c6 = [\u03c6xx, \u03c6yy, \u03c6zz] T includes the moment of inertia elements in the local frame, and cm = [cm,1, cm,2, cm,3] T contains gravitational force parameters", " To conveniently observe the convergence of parameter estimations, we simplify the inertia tensor to a diagonal matrix. With the 1The MATLAB codes for both simulations are available online at: https: //github.com/VibRoLab-Group/IORAC_Exo center of mass (COM) position of the uncertain body defined as di = [di,x, di,y, di,z] T, cm,i is introduced so that along with cm,i,4 = mi \u2212 (cm,i,1 + cm,i,2 + cm,i,3) (21) each cm,i,j (j from 1 to 4) is a point mass that introduces a gravitational force component at a vertex of a load tetrahedron shown in Fig. 1(b). The load tetrahedron is located in the local frame of an uncertain body. Since gravitational forces are conservative in the global frame, the sum of all cm,i,jcg will be the total gravitational force micg , and the weighted average of the translational positions of vertices by cm,i,j is the COM of the uncertain body. The true values of uncertain parameters and default controller parameters are listed in Table I. We first test the controller by assuming no disturbance and only Body 2 is unknown" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001312_s00170-020-05434-3-Figure8-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001312_s00170-020-05434-3-Figure8-1.png", "caption": "Fig. 8 Cross section at different overlapping rates of typical variable width thin-walled structures. a \u03b71, v1 . b \u03b72, v2. c \u03b73, v3. d \u03b74, v4", "texts": [ " 9 Geometric model of the cross-sectional profile of an VWTW part The forming of the first layer is carried out using calculated process parameters. The variable width thin-walled part with the flat top surface was fabricated layer by layer by changing overlapping rate and scanning speed. In order to illustrate the implications of the above definitions, typical variable width thin-walled structures are shown in Fig. 7. The typical variable width thin-walled structure was cut along the four cross-sectional positions shown in Fig. 7 to study its cross-sectional structure (\u03b71 < \u03b72 < \u03b73 < \u03b74). As shown in Fig. 8, it can be observed from the cross section at different overlapping rates that the designed width of the VWTW parts can be achieved according to the above forming process. However, the determination of the Z axis increment (\u0394h) directly affects the accuracy of the height of the VWTW parts, and the unreasonable Z axis increment is liable to cause the forming failure. In the forming process of the VWTW parts, except that the first layer is deposited on the flat substrate, the subsequent cladding passes are performed based on the previous cladding track", " Therefore, after the completion of the subsequent cladding, the total height formed must be lower than the simple single-layer height Hd \u00d7 the number of cladding layers N, so the Z axis increment cannot be set as the height of the single-layer cladding track during the forming process; otherwise, the distance between the powder feeding nozzle and the cladding track will constantly deviate. Therefore, the Z axis increment used to deposit the VWTW parts should be purposefully investigated. In order to obtain a suitable amount of the Z axis increment, the two representative cross-sections in Fig. 8 a and c were selected to investigate the Z axis increment, as shown in Fig. 9. For the cross section with the overlapping rate \u03b71, the middle portion is ignored, and the case where the single-track filler material shortage region of the adjacent layers on both sides is mainly studied. For the cross section with the overlapping rate \u03b72, (\u03b71<\u03b73), the overlapping rate is large and the shape of the cladding layer after multi-track overlapping is approximately parabolic. This part mainly studies the filling of the material shortage area of the whole cladding layer" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002505_s00170-021-06711-5-Figure16-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002505_s00170-021-06711-5-Figure16-1.png", "caption": "Fig. 16 Top four corner a predicted deformation results at 150 W\u2013750 mm/s, b experimental deformation behavior at 150 W\u2013750 mm/s", "texts": [ " For the temperature measurement, the predicting errors between the simulation and experimental results are in a range of 22\u201338 K. These errors can be expected because of the scattering in the experimental data as well as possible variation in the thermophysical properties as utilized for simulation and actual properties during the experiments. Overall, the developed model is appropriate to employ in modeling and simulation of the SLM of Ti6Al4V powder. In order to validate the numerical investigation, experimental investigation has been carried out. Figure 16b presents the experimental 3D-printed part to measure the deformation of top four corners. Table 4 has been presented to show the comparison of predicted deformation results and experimental results. The overall shape distortion comparison shows a good agreement. As shown in Fig. 16b, the distortion is measured at the bottom right of the support material. The predicted top four corner deformations from the model are 248.7, 257.7, 332.1, and 287.5 \u03bcm, which is Point Simulation result (\u03bcm) Experimental result (\u03bcm) Deviation (\u03bcm) (*) 1 248.7 197.2 51.5 in good agreement with the experimental measurement with the highest 74.2 \u03bcm deviation. The difference between the prediction and experiment can be attributed to model simplification, including deformation relaxation in the experiment" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000976_s11044-019-09704-1-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000976_s11044-019-09704-1-Figure1-1.png", "caption": "Fig. 1 Thin-walled four-point contact ball bearing", "texts": [ " Modeling and simulation of thin-walled four-point contact ball bearing are suffering from difficulties such as identification of internal parameters, mechanical and mathematical modeling, and implementation of numerical algorithms. The multibody system software ADAMS is a general-purpose software in the field of multibody dynamics and is being used in the automotive, railway, aerospace, and other general mechanical industries. The following chapters propose an approach for multibody contact dynamics modeling of a thin-walled four-point contact ball bearing, which is incorporated into the commercial MBS code ADAMS. Figure 1 shows a thin-walled four-point contact ball bearing with crown-type cage applied to harmonic gear drive in industrial robots. A thin-walled four-point contact ball bearing has four contact points of ball-to-ring raceway surface contact as theoretically shown in Fig. 2. There\u2019re four ring raceways for four-contact points of ball-to-ring raceway surface contact: two outer-ring raceways and two inner-ring raceways, respectively. The section consisting of four ring raceways is similar to an ellipse; d , Di and Do are the diameters of balls and inner and outer raceway, respectively; ri and ro are radii of curvature of the inner and outer ring raceway, respectively; \u03b1i and \u03b1o are nominal contact angles of the inner and outer ring raceway, respectively; gi and go are the distances to the center of curvature of the inner and outer ring raceway, respectively; hi and ho are the distances of ball to vertex of the inner and outer ring raceway, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001403_j.ymssp.2020.107075-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001403_j.ymssp.2020.107075-Figure2-1.png", "caption": "Fig. 2. Interaction between the ball and the ring of ordinary ball bearings.", "texts": [ " The shim size and radial clearance of the rolling ball bearing with double half-inner rings are denoted by gi and Pd, respectively. The contact condition and operating state of this type of bearing are different from that of ordinary ball bearings. The interaction between the ball and the ring of ordinary ball bearings is first introduced. Considering the special structure of the two halves of the inner ring, the interaction of the rolling ball bearing with double half-inner rings is obtained from ordinary ball bearings. Fig. 2 shows the geometry relationship between the ball and the ring of ordinary ball bearings. Define the inertial frame Oixiyizi, the ball azimuth frame Oaxayaza, the ring fixed frame Orxryrzr , and the contact frame Ocxcyczc . In the inertial frame, the origin Oi is at the center of the circle where the curvature center of the outer raceway is located and the axis xi is collinear with the central axis of the bearing. In the ball azimuth frame, the origin Oa is at the geometric center of the ball, the axis xa is parallel to the axis xi and the axis ya is perpendicular to the central axis of the bearing" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000033_s11665-019-04206-9-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000033_s11665-019-04206-9-Figure1-1.png", "caption": "Fig. 1 Depiction of a wind tunnel balance which acts as a structural link between the model and sting", "texts": [ " These results are viewed with respect to the intended application of this material as a structural element for wind tunnel testing. Keywords additive manufacturing, hot isostatic pressing, heat treatment, precipitation hardened steel, 17-4 stainless steel 17-4 precipitation hardened stainless steel is commonly used to produce force transducers. When these force transducers are designed for wind tunnel measurements, they are referred to as wind tunnel balances or simply balances. A depiction of a balance in a typical wind tunnel application is shown in Fig. 1 where it can be seen that the balance acts as a structural ink between the model and sting. Balances commonly require a year to manufacture (Ref 1) using conventional means because they involve complex, intricate designs and are made from a single piece of metal. Thus, additive manufacturing (AM) of balances is very appealing from both a manufacturing time and cost perspective. The challenge is that balances are routinely designed with safety factors as low as 1.5. This combined with the dynamic forces present in a wind tunnel requires high confidence in the structural integrity of the material used for the balance since failure can result in costly damage to high-value facilities and test article assets" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001717_jmes_jour_1959_001_016_02-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001717_jmes_jour_1959_001_016_02-Figure1-1.png", "caption": "Fig. 1. An element of shell wall", "texts": [ " Internal pressure loading of a complete torus has been examined by Dean (4). The membrane solution for internal pressure by Foppl (5) is of course well known and is discussed, for example, by Timoshenko (13). Nakamura (9) has recently studied internal pressure in vessels with torispherical heads. DERIVATION OF GOVERNING EQUATIONS The governing equations may be found in terms of various quantities. They are briefly derived here in terms of the rotation 4 of the shell meridian, Fig. 2, and the horizontal force F per unit length, Fig. 1, since these are convenient variables in some engineering applications. From the initial geometry, Fig. la, it is seen that, in the limit and d r = d S i n e . . . . (1) * A numerical list of references is given in Appendix 111. t The topic is erroneously stated to be corrugated pipes by Timo- shenko .( 13). J O U R N A L M E C H A N I C A L E N G I N E E R I N G S C I E N C E EQUILIBRIUM Considering only small elastic deformations which do not substantially change the six and shape of an element of shell wall, Fig. 1, horizontal equilibrium gives (to a first order)$ d dr Vertical equilibrium gives (3) sin e - pr)-~,--pr cos e+2hrx = o d dl - (wr)+pr sin e-2ht.y = o whence either or, s i n e = o . . . - (4b) where V is a constant of integration to be determined from the applied vertical load. Equation (4a) relates to a general shell and equation (4b) to the special case of a cylindrical shell. From moment balance d dr - (mor)-Fr cot O+ Wr-m# = o . (5) The horizontal and vertical components F and IV can be related to the normal and shear forces at any section by N,= wcose+Fsine ~ = w s i n e - F c o s e " ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002614_j.compositesa.2021.106449-Figure21-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002614_j.compositesa.2021.106449-Figure21-1.png", "caption": "Fig. 21. Tow-level FE model representation of a pressure vessel, (a) helix wind, (b) hoop wind, (c) indication of slight penetration issue, (d) magnified view of helix wind. Note: the artificial gaps introduced in the helix wind are only used for clarification purpose and the actual tow width needs to be adjusted based on the ascured width.", "texts": [ " A more efficient technique, such as improved integration methods [40] or augmented cohesive element method [41] which allows using comparable or relatively coarser mesh size needs to be integrated into the TWM to predict the failure events. It is worth mentioning that the developed algorithm can record the sets of tow-tow contact pairs which greatly ease the implementation difficulties for subsequent interlaminar failure modelling. The second case study presents the layup of a filament-wound pressure vessel. A 45-degree helical and a hoop wind layer were modelled using the TWM algorithm (Fig. 21a, b) to capture the semiwoven nature of the vessel wall. Although the vessel is manufactured from a single continuous tape, the tow path has been pre-processed to segment the tape at each polar approach. The preprocessing step is facilitated to avoid computational limitations of placing one very long tow, improve the robustness of the z-offset and help with visualization. The separate tows are reconnected at their adjoining ends before any FE modelling is conducted. Several slight penetrations are unavoidable due to the limitations of X. Li et al. Composites Part A 147 (2021) 106449 intersection detection using inner tow points, as shown in Fig. 21c. The penetration can be mitigated by reducing the detection tolerance but comes at a computational cost. As the contact area between the penetrated tows is very small, it can be accounted for during the contactdetection phase of any subsequent simulation. Alternatively in most FE codes such as Abaqus, the penetration issue can be tackled by invoking the option \u201cadjust to remove overclosure\u201d before applying any mechanical loads. The AP-Ply laminate is chosen as the third case study due to the interweaving architecture which generates a network of distributed defects in the laminate" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002847_j.mechmachtheory.2021.104521-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002847_j.mechmachtheory.2021.104521-Figure7-1.png", "caption": "Fig. 7. Photograph of experimental setup.", "texts": [ " 6, were 3D printed from plastic material (a variety of Stratasys RDG131 Digital ABS). The dimensions defining the ND configuration of Fig. 6(b) and skeleton representation of Fig. 4(b) are: d = 0.003 m, l12 = l4 = l67 = 0.02 m, r3 = r5 = 0.004 m. The FD design pictured in Fig. 6(a) and whose geometry is shown in Fig. 4(a) is defined by d = 0.003 m, l2 = l4 = l6 = 0.02 m, r3 = r5 = 0.004 m, r1 = r7 = 0.006 m. For both designs the hinge root radius (see the mechanism of Fig. 5) is R = 0.03 m. The experimental setup, which is pictured in Fig. 7, was designed to measure the vertical (z-axis) displacement uz at the rigid platform center when a downward z-axis force fz is applied at the same point. With the aid of uz and fz, the corresponding experimental stiffness is calculated asKfz \u2212 uz = fz/uz. The force was generated by an electromagnetic actuator (Thorlabs VC625M) and the resulting translation was measured by a laser displacement sensor (\u03bcEpsilon NCDT ILR 1320-25). Several test runs were conducted for both mechanism specimens. Each test run consisted of incrementally applying the force fz in the range 1 N \u2013 6" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002210_j.robot.2020.103670-Figure14-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002210_j.robot.2020.103670-Figure14-1.png", "caption": "Fig. 14. Level sensor.", "texts": [ " The motion controllers are connected to the PC that controls the robot\u2019s motors. The wheelchair is equipped with the same accelerometer and ultrasonic sensor systems (Fig. 13(b)). These wheelchair sensor systems are connected to the robot PC using a ZigBee wireless communications module. Two motion controllers and two batteries are added to the electric drive unit of the wheelchair. These motion controllers are also connected to the robot PC via a ZigBee module. Level sensor systems are set on both the robot and the wheelchair (Fig. 14). When the wheelchair or robot is level with the floor, the installed weights will push against the level sensors. The robot and wheelchair level sensors are connected to the motion controller installed on each vehicle. These sensor systems are able to detect the level of each vehicle and use that information to detect the end of the wheelchair or robot when climbing the step (see Section 3). 3. Step-climbing process When the wheelchair and the robot climb a step, the vehicles are connected one after the other", " [Stage 4] <13> The wheelchair and the robot continue to move forard. <14> Next, the robot\u2019s middle wheels come into contact ith the step as the wheelchair pulls the robot forward. At this oint, the robot\u2019s rear wheels are lowered and provide additional upport as the robot\u2019s middle wheels begin to climb the step. <15> The wheelchair and the robot continue to move forard. The robot\u2019s rear wheels remain lowered. <16> When the robot\u2019s middle wheels have reached the pper level of the step, the robot\u2019s level sensor detects the end of he step-climbing process (Fig. 14). At this point, both vehicles are topped, and the robot\u2019s rear wheels are folded upward (Fig. 4). . Geometrical and static analysis of the step-climbing process In this system, the wheelchair and assistive robot are deployed n a forward-and-aft configuration when the vehicles climb a tep. The system lifts the front wheels of the wheelchair (Stage ) using the velocity differences between the connected vehicles. owever, if the wheelchair is too close to the step in Stage 1, the ront wheels will collide with the step, and a step-climbing failure w b t w t t r s a i t T a w t f 4 s w ( a r p v a t ill occur [23]" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001800_978-3-030-40410-9-Figure4.9-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001800_978-3-030-40410-9-Figure4.9-1.png", "caption": "Fig. 4.9 Operative ranges and specifications of LabVolt 5150", "texts": [ "7) shows that the position of the central point of the end-effector (P) relative to the base frame is p02 = \u23a1 \u23a3 x y z \u23a4 \u23a6 = \u23a1 \u23a3a1.C1 + a2.C12 a1.S1 + a2.S12 0 \u23a4 \u23a6. (4.8) Also, the end-effector orientation relative to the base frame is R0 2 = \u23a1 \u23a3C12 \u2212S12 0 S12 C12 0 0 0 1 \u23a4 \u23a6. (4.9) Example 4.2 (LabVolt 5150 manipulator) Figure 4.8 shows LabVolt 5150 manipulator (Ghafil and J\u00e1rmai 2019), which is a 5DOF, four-link educational robot. The joint limits of the robot actuators, as well as link specifications, are given in Fig. 4.9. To start modelling the forward kinematics equations of this robot, zero-configuration was chosen to be as in Fig. 4.10. The same way as Example 4.1 has been followed to assign six-coordinate frames starting from the base point to the end-effector central point. Note that frame 3 and frame 4 coincide at the same point. Table 4.2 reveals the spatial parameters of the four links of the manipulator. Referring to Fig. 4.10, we have substituted the unknowns in Table 4.2 as follows: d1 = 255.55mm a2 = a3 = 190mm 78 4 Manipulator Kinematics d5 = 115mm The fourHTMs can be determined by substituting each link parameter in Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000093_j.engfailanal.2019.104180-Figure17-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000093_j.engfailanal.2019.104180-Figure17-1.png", "caption": "Fig. 17. Location and symbols of the points of support of the loading bridge.", "texts": [ " For each of the measurement series, the readings were done during lifting and lowering of the loading bridge to eliminate the influence of friction on the values of pressure in hydraulic cylinders. Such a procedure is important to eliminate measurement errors. The results of measurements are reactions in the points of support for chosen placements of the elements of the loading bridge and the location of the center of gravity calculated using said results. The values are determined based on the information on the location of three points of support: P1, P2, and P3 \u2013 P4 (points P3 and P4 have been hydraulically connected creating one point of support). Fig. 17 presents a diagram and symbols of the points of support of the loading bridge\u2019s undercarriage. Fig. 18 presents the P3 and P4 hydraulic cylinders with a pressure gauge used for measurements. Four series of lifting and lowering have been conducted during the measurements for consecutive, individual placement configurations of the elements of the excavator. It was aimed at identifying the distribution of masses of individual, movable elements of the loading bridge. Each series was characterized by different placements of the loading bridge\u2019s caterpillar undercarriage relative to its loadcarrying structure and the last belt conveyor. Table 4 presents a summary of measurement results in the form of locations of the center of gravity (coordinates x and y), eccentricity (e) and the mass of the machine being weighed (G). The coordinate system assumed for calculations is in accordance with the one presented in Fig. 17. The x and y coordinates in the table determine the location of the center of gravity. Based on these results it was stated that the location of the center of gravity is independent of the placement of the elements of the loading bridge and is located in its stability field for the cases of standard loads and relative placements of individual elements of the machine. It was also determined that loading bridge is heavier for 5% in comparison to theoretical design data. What is more, the results of weighing were used for numerical calculations instead of a theoretical mass of the machine" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001197_s42405-020-00259-6-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001197_s42405-020-00259-6-Figure5-1.png", "caption": "Fig. 5 Precise models of the honeycomb panel with different element size", "texts": [ "46 As meshing configurations, such as element size and mesh quality, will affect the simulation results of finite element models to a great extent, a convergence test was carried out in order to clarify the influence of element size on the analysis accuracy of the precise model used in Sect. 2.3. In the convergence test, three typical precisemodels of the honeycombpanelwith element size of 3mm, 6mmand 9mm (as 1 time, 2 times, 3 times the length of the honeycomb rib) were built, respectively, as shown in Fig. 5, static analysis and free mode analysis with the same loadcase settings as the example in former section were carried out. The analysis results are shown inTables 5 and 6, and the relative deviations of the 6mmand 9mmmodelswith respect to the 3mmmodel were also calculated. According to the statistics, the maximum deviation of the 6 mm and 9 mm models with respect to the 3 mm model is no more than 0.94% in static analysis, while in the free mode analysis, the maximum deviation is no more than 1.09%" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002163_is48319.2020.9199967-Figure17-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002163_is48319.2020.9199967-Figure17-1.png", "caption": "Figure 17. Experimental environment (aerial ladder)", "texts": [], "surrounding_texts": [ "We placed the proposed robot on an aerial ladder and a bumpy wall. In these experiments, the robot was manually controlled. Fig. 18 and 20 to 22 show the results. From these results, we confirmed that the complex behavior to complete these tasks is realized by pulling and loosening the strings by the two motors. We conclude that we can control many degrees of freedom of the soft body and produce intelligent behavior by the robot\u2019s dynamics No. 1 No. 2 No. 3 No. 4 Figure 20. Realized behavior (Bumpy walls Pattern 1). No. 1 No. 2 No. 3 No. 4 468 V. CONCLUSION In this paper, we focused on the physical properties of the soft arm and developed a ladder-climbing robot inspired by an octopus. We conducted experiments to demonstrate the effectiveness of the developed robot and confirmed that adaptive intelligent behavior could be realized by a twodimensional control input. Complex movements with many degrees of freedom were produced by the interaction between the soft arm and the environment. In our future work, we will apply machine learning to the soft robot to acquire the control input and demonstrate the effectiveness of the soft body with machine learning. ACKNOWLEDGEMENT This research was partially supported by the Japan Society for the Promotion of Science through the Grant-in-Aid for Scientific Research (18K11445). REFERENCES [1] K. Ito, S. Hagimori, \u201cFlexible manipulator inspired by octopus: development of softarms using sponge and experiment for grasping various objects\u201d, \u201c Artificial Life and Robotics, Vol. 22, issue 3 pp. 283\u2013 288, 2017 [2] K. Ito, R. Aoyagi, and Y. Homma, \u201cTAOYAKA-III: A Six-Legged Robot Capable of Climbing Various Columnar Objects,\u201d J. Robot. Mechatron., Vol.31, No.1, pp. 78-87, 2019. [3] T. Mukai, K. Ito,\u201dFlexible manipulator inspired by Octopi: Comparative study of pushing and pulling mechanisms in realizing intelligent behavior\u201d, Proc. of the Twenty-fourth International Symposium on Artificial Life and Robotics (AROB 24th), pp. 393\u2013396, 2019 [4] K. Ito, Y. Aso and K. Aihara, \u201cMulti-legged robot for rough terrain: SHINAYAKA-L VI\u201d, Proc. of the International Conference on Advanced Mechatronic systems (ICAMech 2019)\u201d, pp. 32, 2019 [5] K. Ito, H. Maruyama, \u201cSemi-autonomous serially connected multicrawler robot for search and rescue\u201d, \u201cAdvanced Robotics\u201d, Vol. 30, issue 7, pp. 489-503, 2016. [6] A. Saito, K. Nagayama, K. Ito, T. Oomichi, S. Ashizawa, and F. Matsuno, \u201cSemi-Autonomous Multi-Legged Robot with Suckers to Climb a Wall\u201d, \u201cJournal of Robotics and Mechatronics\u201d, Vol. 30, No.1, pp.24-32, 2018. [7] H. Yoneda, K. Sekiyama, Y. Hasegawa and T. Fukuda, \u201cVertical Ladder Climbing Motion with Posture Control Considering Gravitation Momentum for MultiLocomotion Robot\u201d, \u201cThe Japan Society of Mechanical Engineers\u201d, Vol. 75, No. 751, pp. 12-19, 2009. [8] M. Kanazawa, S. Nozawa, Y. Kakiuchi, Y. Kanemoto, M. Kuroda, K. Okada, M. Inada and T. Yoshiike, \u201cRobust Vertical Ladder Climbing and Transitioning between Ladder and Catwalk for Humanoid Robots\u201d, Proc. of the 2015 IEEE International Conference on Intelligent Robots and Systems (IROS), pp. 2202-2209, 2015. [9] T. Takemori, M. Tanaka and F. Matuno, \u201cLadder Climbing with a Snake Robot\u201d, Proc. of the 2018 IEEE International Conference on Intelligent Robots and Systems (IROS), pp. 8140-8145, 2018. [10] S. Fujii, K. Inoue, T. Takubo, Y. Mae and T. Arai, \u201cLadder Climbing Control for Limb Mechanism Robot \u201cASTERISK\u201d\u201d, Proc. of the 2008 IEEE International Conference on Robotics and Automation (ICRA), pp. 3052-3057, 2008. [11] D. Rus, M. Tolley, \u201cDesign, fabrication and control of soft robots\u201d, Nature, Vol.521, pp.467-475, 2015 [12] G. Sumbre, Y. Gutfreund, G. Fiorito, T. Flash and B. Hochner, \u201cControl of Octopus Arm Extension by a Peripheral Motor Program\u201d, Science, Vol. 293, No. 5536, pp. 1845-1848, 2001. [13] Y. Gutfreund, T. Flash, G. Fiorito and B. Hochner, \u201cPatterns of Arm Muscle Activation Involved in Octopus Reaching Movements\u201d, \u201cThe Journal of Neuroscience\u201d, Vol. 18, No. 15, pp. 5976-5987, 1998. Authorized licensed use limited to: Carleton University. Downloaded on November 04,2020 at 13:12:36 UTC from IEEE Xplore. Restrictions apply." ] }, { "image_filename": "designv11_5_0002297_khpiweek51551.2020.9250138-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002297_khpiweek51551.2020.9250138-Figure5-1.png", "caption": "Fig. 5. Schematization of magnetic flux paths in AMB1 and AMB3.", "texts": [ " (7) The full energy of the magnetic circuits is calculated by the sum of circuit section ones. Then W1 for AMB1 (W2 for AMB2 similarly) and W3 for AMB3 are: ). )( )(( 2 1 , 2 1 2 8917 2 8818 2 8717 2 8620 2 8519 2 8418 2 8317 2 82182019 2 811718173 8 1 2 32 2 24 2 16 2 8 2 1 \u03a6+\u03a6+\u03a6+\u03a6+\u03a6+ +\u03a6+\u03a6+\u03a6+++ +\u03a6++= \u03a6+\u03a6+ +\u03a6+\u03a6+\u03a6 = = ++ ++ allgg ggspp spp k kkakkl kkskkgkkp RRRRR RRRRR RRRW RR RRR W (8) Magnetic resistances Rk included in the magnetic energy expressions of each AMB type can be found on the basis of formula (5) and magnetic flux path diagrams shown in Fig. 5. Magnetic resistances of some AMB circuit sections are constant and depend only on geometric dimensions of these sections and environment properties (Rp, Rs, Rl, Ra), and also depend on the generalized mechanical coordinates q (for example, displacements of supporting sections x1, y1, x2, y2, z3) for the active magnetic bearing clearances or disk Rg. Some resistances are given in (9) and (10) for the radial AMB1 and the axial AMB3 respectively: + = \u2212 += \u2212=\u2212= \u2212=\u2212= ==\u2212= + + . ),( dd 2 )( )(; 2 1 ;2/)(),2/()( ;2/)(,)2/()( );8," ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000950_tmech.2019.2941279-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000950_tmech.2019.2941279-Figure6-1.png", "caption": "Fig. 6: (a) Completed CAD model of the bio-inspired wrist for simulation with OpenSim (Human dynamics simulator). (b) Energy consumption comparison between the proposed wrist and the other wrist of [5] using the OpenSim.", "texts": [ " By using the simulation framework, we compared our bioinspired wrist with another model of [5] in terms of muscle energy consumption in the whole upper limb, even though Verleg\u2019s wrist is lighter. By replacing the wrist part in the original human model with each prosthetic wrist part, we compared their energy consumptions for the same tasks. The task of throwing a bottle was selected for the simulation as it requires rotations in both FE and RU directions. During the task, the metabolic energy consumptions of both wrist models were estimated using a Houdijk\u2019s energy model [23]. As shown in Fig. 6b, our bio-inspired wrist can largely reduce the energy consumption by 56 % (from 5.05J to 2.84J) in the muscles of the whole upper limb. Energy consumption of our proposed wrist in the whole upper limb muscles is much lower than that of the other model, while particular actuators of the proposed wrist consumed more energy. Approximately 24 % larger energy consumption in those actuators is considered to be caused by the large wrist RoM, which, however, reduce extra motions, and consequently energy consumption, in the upper limb" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001812_1077546320912109-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001812_1077546320912109-Figure2-1.png", "caption": "Figure 2. NU 205 cylindrical roller bearing.", "texts": [ " Rotor mass and its unbalanced effect have been added and no other external load interactionwith the system has been considered. The \u2018+\u2019 sign in the subscript indicates that the deformation is only applicable when there is compression between the roller and raceways. The equations are solved using the fourth-order Runge\u2013Kutta iterative method. As the outer race is fixed to the casing, deflection due to rotor load (W) is compensated by the inner race deflection, i.e. xir \u00bc yir \u00bc 10 6 m and initial velocities, i.e. _xir \u00bc _yir \u00bc 0m=s. The dimensions and the sketch of NU 205 cylindrical roller bearing (Figure 2), considered in the present study for the dynamic behavior, are given in Table.1. In this section, mathematical simulation of local defects (spalls) in races and rollers is given. For a given size (s) of the localized defect, the angular width (\u03a6r) will vary with the location of the defect, viz., inner race, outer race, or roller. As the rollers start repetitive hitting of the races in the load zone, stresses are developed, which causes the initiation of deformation at the subsurface level, and gradually spread to surfaces and then grow" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002693_s40192-021-00209-4-Figure14-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002693_s40192-021-00209-4-Figure14-1.png", "caption": "Fig. 14 Lack-of-fusion pattern for case B35: a predicted density field after laser scanning presenting the lack-of-fusion pattern. The color legend indicates the density at spatial points. The blue represents the effective powder density ( 4300 kg\u2215m3 ) and the red represents the solid material density ( 8400 kg\u2215m3 ). b Experimentally observed lack of fusion from the top view reported in reference [45]", "texts": [ " The quantitative comparisons of melted track geometries for those cases are provided in Supplementary Figs.\u00a01 to 5. More interestingly, the developed model can predict lackof-fusion patterns in the L-PBF process. Lack of fusion, where there is no complete adherence of the current melt to the surrounding part, is a type of AM process-induced defect that has a negative influence on as-built mechanical properties. It is caused by insufficient energy input that cannot create a large enough overlapping region between two adjacent tracks [44]. Figure\u00a014a shows the predicted lack-of-fusion patterns for case B35. The toolpath of this case is provided in Fig.\u00a06. As shown in Fig.\u00a014a, the color legend indicates the density field after the L-PBF process. The minimum value (illustrated as blue) is the effective powder density ( 4300kg\u2215m3 ) and the maximum value (illustrated as red) is the solid material density ( 8400kg\u2215m3 ). The top layer is originally assigned as the effective powder density. During the L-PBF process, if the peak temperature of a spatial point is higher than the solidus temperature, the density of this point will be updated that of the liquid. After the temperature drops down to solidus temperature, the density is converted to that of the bulk solid, which is much higher than the effective powder density. The values of those densities are provided in Table\u00a01. During this process, a few powders between two laser scanned tracks might remain unmelted if the energy input from the laser is insufficient to create a large enough melt pool. Those small regions become lack of fusion after the toolpath is complete. For example, Fig.\u00a014b shows a few unmelted powders and open pores between the laser scans reported in reference [45]. Our model can predict the lackof-fusion regions (small blue dots in Fig.\u00a014a) between laser 1 3 scanned tracks and toolpath blocks. The pattern of the lack of fusion depends on toolpath strategy and process parameters. The developed model is a potential effective tool to optimize the process parameters for mitigating or eliminating the lack-of-fusion defect. For example, Fig.\u00a015 shows predicted lack-of-fusion patterns for three cases: B27, B31, and B34. These cases have the same toolpath pattern and process parameters except for the hatch spacing. The hatch spacing of the three cases is 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002431_s11694-020-00780-y-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002431_s11694-020-00780-y-Figure7-1.png", "caption": "Fig. 7 Peak currents for aflatoxins of AFM1, AFB1 and AFB2", "texts": [ " The detection limit and linear range obtained from different analytical methods were summarized in Table\u00a01. As can be seen in this table, the detection limit and linear range for AFM1 determination at the proposed biosensor are comparable to those published in other literatures, indicating that the ss-HSDNA/AuNPs/ECNF mat electrode is promising for AFM1 detection. The selectivity is significant factor to investigate the performance of aptamer. To judge the selectivity of aptamer, the sensing results of AFB1, AFB2 and AFM1 were (1)LOD = 3\u03c3 |s| 1 3 compared. As shown in Fig.\u00a07, the response signal to the AFB1 and AFB2 is approximately negligible whereas that to the AFM1 is reasonable. The obtained result demonstrates that the selectivity of the aptasensor, due to its high affinity towards AFM1 molecules, is acceptable for the practical application. This is in good agreement with result reported by Sharma et\u00a0al. [24]. The reproducibility of the ss-HSDNA/AuNPs/ECNF mat aptasensor was carried out by measuring the current response of the four electrodes (Fig.\u00a08a). The aptasensor showed that the relative standard deviation (RSD) was less than 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001800_978-3-030-40410-9-Figure9.2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001800_978-3-030-40410-9-Figure9.2-1.png", "caption": "Fig. 9.2 Four-bar mechanism", "texts": [ " In crossover a block of genes (0,1) from one parent can be replaced with another block of genes from another parent; in our example, we could replace block b by d and vice versa. The result is two different binary strings that can be exposed to the mutation operation which is replacing a specific gene with different value, for instance, the gene highlighted in grey is replaced by 0 instead of 1. Crossover and mutation are random operations, and there are a lot of methods on how to do that. In this section, we will explain how to represent a mechanism by genetic operators with an example given by (Oliva and Goodman 2010) and shown in Fig. 9.2. Genetically, this mechanism may be represented by a chromosome of seven genes each one corresponds to a physical component. Two genes for x and y coordinates of the pivot point a, Two genes for x and y coordinates of the pivot point d, gene for link length ab, a gene for link length bc, and a gene for link length cd that are shown in Table 9.1. In the beginning, these seven genes initiated randomly with random numbers, by assuming that the objective is to meet a given motion, GA keeps iterated until it finds the proper values of the genes that enable the 4-bar mechanism to perform the given motion", " Graph theory (West 2001; Bollob\u00e1s 2013) has some terminologies that can be described briefly in this text to give background for subsequent sections. A collection of sets of vertices V and edges E is called graph and can be described by the triple (V, E, g). V is always nonempty set while E could be an empty set. g is mapping called incidence mapping, and for each edge, e \u2208 E there is subset g(e) = {u, v} where u and v are vertices. Figure 9.4 explains an example of a graph with four vertices v1, v2, v3, and v4 as well as four edges e1, e2, e3, and e4. Consider Fig. 9.2, every edge is said to be incident with the vertices that it connects, so edge e2 is incident with vertices v1 and v2. Also, the vertex is said to be incident vertex if it belongs to one or more edges, so v1 is incident with edges e1 and e2. Vertices are said to be adjacent if the same edge connects them, for example, v1 is adjacent to v2 and v4. The edge is said to be adjacent to another if they have a common vertex, for example, e3 and e4 are adjacent edges. An edge has the same end vertices, for example, e6 in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002414_012043-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002414_012043-Figure3-1.png", "caption": "Figure 3. Mises stresses distribution in rotor system elements.", "texts": [ " Let us set some nominal parameters of rotor system under study (see Figure 1): shaft length 0L = 0,24 m, shaft diameter 0d = 0,019 m, disk diameter 0D = 0,18 m, length l = 0,12 m, distance between bearing supports 0l = 0,1 m, shaft material is steel (elasticity modulus \u0415 = 2\u221910 11 N/m 2 , Poisson's ratio =0,3); disk with blades made of aluminum alloy (elasticity modulus \u0415 = 7,1\u221910 10 N/\u043c 2 , Poisson's ratio = 0,33); nominal rigidity of bearing supports \u0441 = 1,96\u221910 7 N/m, nominal operating speed 0 = 20 thousands rpm. First, a quasi-static analysis of this rotor loaded by centrifugal forces has been performed with a commercial finite element package ANSYS Workbench. Figure 2 shows the obtained distribution of the total displacements of rotor system elements by nominal parameter values. Similarly, von Mises stresses distribution in this system are presented in Figure 3. It can be seen that the stress-strain state of the rotor system is characterized by a substantially heterogeneous distribution of displacements u and stresses \u03c3 . In particular, the largest displacement values are at the periphery of the blade disc. However, the dominant component cannot be distinguished: the axial and radial displacements are proportional to each other. Therefore, it is necessary to control the change in both the radial and axial clearance between the rotor and stator parts of the construction" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001241_j.ijfatigue.2020.105632-Figure8-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001241_j.ijfatigue.2020.105632-Figure8-1.png", "caption": "Fig. 8. FEM model, (a) overall model, (b) the model of planetary gear.", "texts": [ " In this way, the process of fretting contact will cause fretting wear and fretting fatigue. 2.2. \u2a7eT TjGB When the instantaneous torque TGB is greater than the limit torque Tj, the conditions of Amonton's law are not satisfied. According to the previous force analysis, it is clear that slipping will occur between the planetary gears and bearings. To further verify the situation of the planetary gears during operation, this study uses finite element software to model the planetary gear trains. The established finite element model is shown in Fig. 8. The modulus of the gears in the model is 10 mm, and the mesh sizes of the gears (sun gear, planetary gears, and Table 2 Chemical compositions of gear and bearing. Chemical compositions and contents (%) C Si Mn Cr Mo Ni P S Gear Gear 0.19 0.30 0.66 1.61 0.25 1.40 \uff1c0.005 0.0016 18CrNiMo7-6 0.15\u20130.21 \u22640.40 0.50\u20130.90 1.5\u20131.8 0.25\u20130.35 1.4\u20131.7 \u22640.035 \u22640.035 Bearing Inner ring 1.00 0.30 0.40 1.31 0.020 \u2013 0.0094 0.0040 Outer ring 0.98 0.30 0.38 1.30 0.019 \u2013 0.0088 0.0044 SUJ2 0.95\u20131.10 0.15\u20130.35 \u22640.50 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000654_smc.2016.7844562-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000654_smc.2016.7844562-Figure1-1.png", "caption": "Fig. 1. Octorotor geometry", "texts": [ " The control approach presented in this paper yields similar results as the state of the art, which suggests that the model obtained by this paper is correct. The remainder of the paper is organized as follows. In Section II the mathematical model of the octorotor system is derived. Section III describes the synthesis of a simple stabilization controller around hover configuration, while control 978-1-5090-1897-0/16/$31.00 \u00a92016 IEEE SMC_2016 002182 simulation results are given in Section IV. Concluding remarks and future work are outlined in Section V. The octorotor system shown in Fig. 1 consists of eight arms of the same length that are fixed to a support plate, where each arm is equipped with a DC motor driving a fixed pitch rotor. The spacing between two adjacent arms is 45 degrees. Rotors depicted with blue color (2, 4, 6 and 8) rotate clockwise, while the ones depicted with red color (1, 3, 5 and 7) rotate counterclockwise. Therefore, the octorotor is inherently balanced with regards to the drag moment generated by the rotors. All other system components (on a real system) should be placed in a protective case which is mounted to the support plate" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000888_j.promfg.2019.06.223-Figure17-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000888_j.promfg.2019.06.223-Figure17-1.png", "caption": "Fig. 17. An AM system with two robotic arm actuators for fabrication of a multi-material fixture", "texts": [ " This shows that the proposed approach, with more realistic spatial constraint modelling and fewer independent actuators, compares favorably with the current methods, which virtually over-simplify or consider very few operational constraints. In manufacturing, fixtures are often used for securing workpieces. A possible application of AM is to fabricate Yi Cai et al. / Procedia Manufacturing 34 (2019) 584\u2013593 591 8 Author name / Procedia Manufacturing 00 (2019) 000\u2013000 customer-specific fixtures. Fig. 17 shows an AM system consisting of two Cartesian robotic arm actuators for fabrication of a fixture of eight materials. The color STL model of the fixture model is sliced into 250 layers, and the hatch width for internal filling of the CFs is 1mm. According to their distribution, the eight materials are assigned to the two robotic arms. To model the region-based constraint, R7 and R3 are set as the work regions for these two actuators respectively. Moreover, a 20mm safety distance is assigned to each robotic arm according to the dimensions of their end-effectors" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000187_s10846-019-01113-y-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000187_s10846-019-01113-y-Figure4-1.png", "caption": "Fig. 4 Overview of working environment", "texts": [ " The other one is cascade connection. In practice, two manipulators which are the Core Module Manipulator (CMM) with larger size and the Experimental Module Manipulator (EMM) with smaller size, can end-to-end connect with each other, thus presenting 14DOFs. Figure 2 and Fig. 3 demonstrate these innovative functions. In this paper, mathematic solutions are provided. The working environment of space manipulators is quite complex, involving four space stations and other surrounding obstacles, as is illustrated in Fig. 4. In most cases, manual operation of manipulators is high-demanding and inaccurate. Hence, automatic trajectory planning with collision-avoidance for space manipulators is of crucial importance. Potential collision should include the collision against the environment and against the manipulator itself [6\u20139]. Here, an improved discrete-collision prediction method is proposed to eliminate the tunneling problem and reduce the time consumption significantly. The disturbance of base is not discussed. Due to the redundant DOF, path planning in joint space is more suitable than in Cartesian space, yet also more challenging" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002726_s12206-021-0639-4-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002726_s12206-021-0639-4-Figure6-1.png", "caption": "Fig. 6. The one-DOF rigid link model.", "texts": [ " The data are obtained from a set of experiments on the proposed PAM, and the relative experimental setup has been introduced from the Refs. [47, 48]. For safety, inner air pressure ranges from 0 to 4 bar at a step of 0.2 bar, and the maxi- mum contraction force is over 100 N with 7.5 kg external loads. Since the elbow exoskeleton is firmly anchored to the human arm and shoulder, the upper arm is assumed not involved in the elbow extension and flexion movement [35]. Hence, the system is simplified as a 1-DOF rigid link model (see Fig. 6). It can be described in the following form: + + + =dJq G \u03b4 \u03c4 \u03c4 (1) where J denotes the inertia of the link; q represents the angle for the revolute joint; q is the angular acceleration for the revolute joint; \u03c4 is the generalized input torques of the actuator; G is the gravitational torques; \u03b4 is the sum of the external disturbances and modeling uncertainties; d\u03c4 is the joint friction torques. At time instant k , the state form of the model is defined as follows: ( ) ( ) ( )\u2212 = k D k q k J \u03c4 , (2) ( ) ( ) ( ) ( )= + + dD k G k k k\u03b4 \u03c4 " ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001197_s42405-020-00259-6-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001197_s42405-020-00259-6-Figure4-1.png", "caption": "Fig. 4 Finite element models of the honeycomb panel", "texts": [ " Deformation under typical static loadcases and freemodes of the honeycomb panel described above are analyzed using the equivalent modeling method proposed in this paper, and then these statistics are compared with the results of the precise model, in this way, the accuracy of the modeling method proposed in this paper can be evaluated. In the precise modeling method, two-dimensional shell elements are used to simulate both the thin-walled cellular structure and the two panels of honeycomb sandwich structure. As shown in Fig. 4a, the precise model contains 3814 nodes and 5814 shell elements of type CTRIA (triangular three-node shell element) or CQUAD4 (quadrilateral fournode shell element). According to the method in this paper, the finite elementmodel of the honeycombpanel in the example is obtained as shown in Fig. 4b, there are 1880 nodes, 836 shell elements of type CQUAD4 and 1254 solid elements of type CHEXA (six-sided solid element with eight nodes) in the equivalent model. A uniform element size of 6 mm was used in the meshing of both modeling methods. In this example, the main direction of honeycomb structure is the X direction of the coordinate system in the Fig. 4. In the analysis, six degrees of freedom of all nodes at the end of + X side in the honeycomb panel are constrained, uniform tensile load Px (force that along X axis), in-plane shear load Py (force that along Y axis), normal shear load Pz (force that along Z axis), pure bending load M (torque that around Y axis) are applied at the other end, respectively. The maximum displacements of the honeycomb panel under different loadcases are calculated using different models, as shown in Table 3. In the loadcases of Px , Py, Pz, the displacements are the components along the direction of loading forces, respectively, while in the loadcase ofM, the displacement is the magnitude" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001329_icac49085.2019.9103346-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001329_icac49085.2019.9103346-Figure6-1.png", "caption": "Fig. 6. Boundary area where 1 is near, 2 is nearer and 3 is nearest", "texts": [ " After classifying the words accurately then only it would capable of recognizing the uncertainties of the command. Especially if there is no commanding verb in the user\u2019s input the system will not identify it as a command. Example results of the uncertainty detection process can be denoted as follows, Command: \u201cGo closer to the bag quickly\u201d Output: Pointed object: bag Uncertain distance: closer >> position (comparative) Uncertain velocity: quickly >> position (normal) We consider the command, \u2018go near the chair\u2019 the uncertain term \u2018near\u2019 is defined as shown in Fig. 6. When defining the value for the uncertain terms instead of giving a constant value, a percentage value is given because the value can be changed depending on the situation. Initially, the value for the uncertain term (near) is gathered using the website and stored in CSV files locally. Then the percentage value for the term \u2018near\u2019 was taken using an ontology knowledge expert. PREFIX rdf: PREFIX owl: PREFIX rdfs: 0, and the steering angle is \u03b4 (t). The angle from the X-axis to the line O\u2039 \u2212 Oh is \u03b8hb, from the X-axis to the line Oh \u2212 Ot is \u03b8tb. The configuration of the trucktrailer system for backward motion is represented by qb =[ xh yh \u03b8hb xt yt \u03b8tb ]T . The variables in the vector qb is measured from the localization system of the truck-trailer. The kinematic model of the truck-trailer system for backward motion is derived as q\u0307b = \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 x\u0307h y\u0307h \u03b8\u0307hb x\u0307t y\u0307t \u03b8\u0307tb \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 = \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 vh+ cos \u03b8hb vh+ sin \u03b8hb \u2212 (vh+/lh) tan \u03b4 vh+ cos (\u03b8tb \u2212 \u03b8hb) cos \u03b8tb vh+ cos (\u03b8tb \u2212 \u03b8hb) sin \u03b8tb \u2212 (vh+/lt) sin (\u03b8tb \u2212 \u03b8hb) \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 " ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002502_j.tws.2021.107512-Figure16-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002502_j.tws.2021.107512-Figure16-1.png", "caption": "Fig. 16. Comparison of the bifurcation curves of different buckling modes of the ircular plates at the crests predicted by the reduced model.", "texts": [ " As mentioned above, two manners can be adopted to predict the wrinkling patterns of the structures. In this section, in order to analyze the buckling modes that we are interested in without introducing any additional initial defects to the geometry, only the zero and first harmonics are adopted, while the wave number is preassigned. The material properties and geometric dimensions are given in Table 4. i 4 c b d b d c c O i B e T n n i d c o a n 4.1. Three buckling modes of close critical loads Fig. 16 illustrates the comparison of the bifurcation curves of different buckling modes of the circular plates at the crests predicted by the reduced model. The bifurcation points of the modes are close, especially the ones of \ud835\udc451\ud835\udc366 and \ud835\udc452\ud835\udc363. Thus, when simulating such wrinkling patterns by the full shell model, the initial perturbation should be carefully applied to detect the buckling mode, which leads to significant difficulties in nonlinear calculation. However, it is easier to adopt the reduced model for the analysis, since an uniform transverse perturbation force is given while only the wavenumber needs to be adjusted" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000555_j.mechmachtheory.2016.09.009-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000555_j.mechmachtheory.2016.09.009-Figure1-1.png", "caption": "Fig. 1. A planar 3-DOF 3-RRR parallel manipulator.", "texts": [ " Given that N N> *, the constraint forces tend to zero as the singularity is approached if and only if the following conditions are also satisfied in addition to the removability conditions of the second form: \u03bc t i m n( ) = 0, = 1,\u22ef, \u2212i N s ( +1) Once that is the case, the singularity-removed constraint force equations of the second form are given by \u23a7\u23a8\u23a9\u03bb t \u03bb t t t i m n t t i m n \u02dc ( ) = ( ), \u2260 , = 1,\u22ef, \u2212 0, = , = 1,\u22ef, \u2212i i s s Proof. Under the removability conditions of the second form, the proof follows simply by substituting \u03bc t( ) = 0i N s ( +1) , i m n= 1,\u22ef, \u2212 into Eqs. (32) and (33), respectively. Q.E.D. Consider a three-degree-of-freedom (3-DOF) 3-RRR planar parallel manipulator (i.e. n = 3), as shown in Fig. 1. Links 1 to 6 are homogenous uniform slender bars with masses m1 to m6 and lengths a1 to a6, respectively. The location of the mass centerG7 of link 7, which has mass m7 and centroidal moment of inertia I7, is given byC G c=1 7 7 and G C C \u03b2\u2220 =7 1 2 . The base joints A1, A2 and A3 are to be actuated by motor torques \u03c41, \u03c42 and \u03c43, respectively. Gravity is assumed to be acting in the negative z-direction. The following numerical values are selected for the robot parameters: a0=1.2 m, b0=0.74 m, c0=1" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002142_b978-0-08-102965-7.00006-0-Figure6.15-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002142_b978-0-08-102965-7.00006-0-Figure6.15-1.png", "caption": "Figure 6.15 (A) Computed-aided design model illustrating the form and components of the right foot passive dynamic response ankle-foot orthose (PD-AFO). (B) Schematics of the build orientation and position of PD-AFO within the selective laser sintering build [99].", "texts": [ " [98] used the SLS system to fabricate patient-specific prototype wax implant of the partial nose. Their experiments exhibited such promising results that they followed their research by manufacturing a PS based nasal prosthetic and tested that on a patient as shown in Fig. 6.14. The team reported a satisfactory nasal defect restoration and cited several benefits for the affected patients from the prescribed treatment mode. Schrank et al. [99] discussed the physical and dimensional features of passive dynamic response ankle-foot orthoses (PD-AFOs) fabricated from nylon via SLS. Fig. 6.15A illustrates the structure and various components of the orthosis while in Fig. 6.15B volume buildup of the orthosis in a SLS system is illustrated in schematic form. AM technologies like SLS have been used to create complex anatomical models that are a source of valuable assistance during clinical treatment and medical education. These models can be easily customized and exhibit the desirable features for comparative assessment prior to a complex surgery or to teach the medical students about the complications of human anatomy and the comprehension of pathology [100]. Negi et al" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000561_978-3-319-42975-5_14-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000561_978-3-319-42975-5_14-Figure5-1.png", "caption": "Fig. 5 Real Duckiebot and its simulation model", "texts": [], "surrounding_texts": [ "Simulations are a powerful tool for teaching robotics, since allow the creation of cheap robotics environments, and access to the latest (simulated) robotics technologies at a very low cost. Simulations can be simplified for teachers and students by making use of web based simulators. Web based simulators universalize the access to such software because allow anybody use them with any device with a fraction of the cost of the desktop simulator. Finally, simulations can be integrated in a complete learning system that includes real robots in the path, minimizing costs and maximizing usage and students experience (ORTE method). Acknowledgments We would like to thank all the Duckietown team for making it possible and specially Andrea Censi for making us part of the project." ] }, { "image_filename": "designv11_5_0000082_s11665-019-04280-z-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000082_s11665-019-04280-z-Figure2-1.png", "caption": "Fig. 2 Process of SLM-fabricated 24CrNiMoY alloy steel sample", "texts": [ " Orientation imaging microscopy (OIM) analysis of the fabricated alloy steel was performed by electron backscattered diffraction (EBSD) system. (The scanning step was 0.2.) The microhardness along the building direction of the asSLM samples was measured using a Wilson Wolpert 401 MVD microhardness tester with a 200 g load and a dwelling time of 10 s. The tensile property was performed by an AG-X100KNEl electronic universal material testing machine at a loading rate of 0.5 mm/min. The tensile sample size and shape are shown in Fig. 2. The tensile fracture morphology is observed by FESEM. The relative density of SLM-fabricated alloy steel was measured by the Archimedean drainage principle. The samples were weighed in the air-marked dry weight M1. Then the samples were put into the wide-mouthed bottle filled with boiling water. Subsequently, connect the bottle to vacuum pump. After taking out the samples, put them in a tray (immersed in water) to measure floating weight M2; finally, dry them with a filter paper to measure the wet weight M3" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001800_978-3-030-40410-9-Figure6.10-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001800_978-3-030-40410-9-Figure6.10-1.png", "caption": "Fig. 6.10 2DOF planar revolute-joint manipulator", "texts": [ " For instance, suppose that we have the initial configuration at time t0, one via point at time t1, and the final configuration at time t f and then the constraints on the polynomial will be q(t0) = q0 q\u0307(t0) = v0 q\u0308(t0) = \u03b10 q(t1) = q1 q(t f ) = q f q\u0307(t f ) = v f q\u0308(t f ) = \u03b1 f If we would like to use all seven constraints on a polynomial, then this polynomial should be of the 6th order: q(t) = a0 + a1t + a2t 2 + a3t 3 + a4t 4 + a5t 5 + a6t 6. (6.20) As a general rule, we can say that the order of the polynomial should be polynomial order = Number of the used constraints\u22121. (6.21) Example 6.4 Figure 6.9 shows a path planned using particle swarm optimization given an initial point and target point. Cubic spline has been used for this problem, where three control points give the shape of the cubic spline function and they are the via points for the trajectory planning. The planar 2DOF revolute-joint manipulator shown in Fig. 6.10 is required to follow this path with its end-effector, as shown in Fig. 6.11. Table 6.1 shows the start, target, and via points in the Cartesian space and mapped to the joint space using inverse kinematics. For this example, wewill assume that the robot begins tomove from the start point at time equal to zero t0 = 0 and reaches the final configuration at time t f = 1 s. The time interval between 0 and 1 s is equally spaced so that robot reaches configuration 1 at time 0.25 s, configuration 2 at time 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000801_j.jpowsour.2019.05.001-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000801_j.jpowsour.2019.05.001-Figure1-1.png", "caption": "Fig. 1. The design of flow channel plates for (a) typical EBC (with pump) and (b) self-pumping EBC (without pump); (c) schematic diagram of top and bottom end plates in the self-pumping EBC stack and (d) photograph of self-pumping EBC assembly.", "texts": [ " All these serial measurements were conducted without replacing the new membrane inbetween. The cell right after each testing was stored in a refrigerator at 4 \ufffdC, as suggested by the enzyme manufacturer, to prevent the degradation of GOx from ambient temperature. The typical EBC and self-pumping EBC rely on the same electrochemical mechanism to convert the fuel into the electricity. However, the different concept of fuel supply between the typical EBC and selfpumping EBC makes the design of flow channel plate distinct, as shown in Fig. 1(a) and Fig. 1(b). The flow plate of parallel type pattern composed of 27 channels for self-pumping EBC is detailed as shown in Table S1. It can be found that the fuel supply of self-pumping EBC based on the capillary effect is allowed to miniaturize its dimension of flow channel plate and thereby provide the advantages of the lightweight and portable applications. In addition, there are extra three reservoirs designed on top and bottom end plates of self-pumping EBC stack, as shown in Fig. 1(c). The photograph of self-pumping EBC assembly is shown in Fig. 1(d). The function of reservoir 1 is the fuel supply while that of reservoirs 2 and 3 loaded with cottons can be offered with a selfdriven source to create the capillary effect. During the operation, the fuel is firstly added into the reservoir 1, transported to reservoir 2 by capillary force, and arrived at the inlet of flow channel plate. The fuel then flows through the anode/cathode flow channel plate where the oxidation/reduction reaction occurs. Once the fuel reaches the outlet of flow channel plate, the wetted cotton in reservoir 3 can be observed" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001158_s00170-020-04957-z-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001158_s00170-020-04957-z-Figure1-1.png", "caption": "Fig. 1 Schematic diagram of thermal error of gear hobbing machine", "texts": [ " In the process of gear hobbing, the heat sources, such as the built-in motor, high-speed bearings, and cutting contact position, generate a lot of heat. Through radiation, convection, and conduction, some of the heat diffuses into the air, while the rest is absorbed by the major components such as bed, coolant liquid, and cutter. The temperature gradient could lead to the tilting of the upright and the elongation of the bed. So, the relative displacement between hob and the center of workpiece axis is changed, as shown in Fig. 1. The change of the distance between the hob and the workpiece spindle is usually defined as the thermal error in X-direction, which is used to reflect the error in the center distance. As shown in Fig. 2, when the thermal error \u03b4x occurs, the cutting point moves from K toK\u2032. The lineKK\u2019 is perpendicular to the gear reaming, and it is defined as the total error of tooth profile. The calculation formula of KK\u2019 can be expressed as F\u03b1 \u00bc KK 0 \u00bc \u03b4xsin\u03b1 \u00f01\u00de where F\u03b1 represents the total error of tooth profile; \u03b1 is the angle between KK' and the horizontal direction" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002356_978-3-030-50460-1-Figure27.2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002356_978-3-030-50460-1-Figure27.2-1.png", "caption": "Fig. 27.2 (a) Construction of a masonry vault by means of centrings; (b) Example of geometrical description of a centring as a polycentric arc.", "texts": [ " Then, by means of Monte Carlo simulation, the collapse condition was studied, whose statistical moments up to second order and probability density functions was rated for different levels of uncertainties. The comparison between the results for the nominal geometry and those obtained studying the reconfigured random arches outlined the influence of that shape uncertainties towards the bearing capacity of those masonry structures. The arched structures are one of the most ancient load-bearing systems used already in the IV millennium b.C. This constructive technique foresees the use of a temporary shoring: the centrings (Fig. 27.2(a)). Those elements are essential to prop up the voussoirs until they became a self-supporting masonry structure. The design of 458 Cavalagli, Gusella, Liberotti such wooden frames envisages the union of different arc sectors with a view to outline the intrados corresponding with the designed arch (Veihelmann and Stefan, 2015). It is conceivable that such a construction process may result in geometrical imperfections affecting the final arch shape, so that, in this paper, this topic is ascribed as a-priori source of uncertainties and therefore investigated. A first attempt useful to describe the centrings design, conceived by composition of several arc sectors, was the geometrical descriptive approach, i.e. the graphic construction of the polycentric arch, which is chosen as the reference shape for the following modelling implementation (Figure 27.2(b)). In literature Veihelmann and Stefan (2015) reported that the number of the arc sectors which constitute a centring is related to the vernacular constructive customs and to the project-size of the pursued arch. Without loss of generality in the method presentation, in this work polycentric arches composed by five arc sectors were studied, since it results a recurring circumstance for mildly sized arches. In order to achieve a simulation of the constructive method closer to the real one, the modelling procedure of the polycentric arches was addressed with a probabilistic approach by means of input random variables, which allow to obtain asymmetrical and irregular 2D random arches, starting from a nominal deterministic geometry" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001969_j.adhoc.2020.102236-Figure12-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001969_j.adhoc.2020.102236-Figure12-1.png", "caption": "Fig. 12. YAW Rotation.", "texts": [ " B: Direction change process The UAVs\u2019 direction change, in our model, simulates the concept f the direction change of the rotating magnetic field, where the d d m i i c t t t t t g O t r N T N Fig. 11. Rotating Magnetic Field. s o s a r t N T R irection of its poles changes over time. To model this process, we efine a constant value named \u201dCommonDirection,\u201d which is com- on to all the UAVs, by which each UAV changes direction periodcally. The objective of defining \u201dCommonDirection\u201d as a constant s to model the case of an ideal rotating magnetic field ( Fig. 11 ) hanging direction at a constant angular rate. Indeed, in our model, we consider two rotation forms, namely he YAW (Z-Axis) ( Fig. 12 ) and the ROLL (Y-Axis) ( Fig. 13 ). We use he YAW rotation to change the direction to the desired one; on he other hand, the ROLL is used to go through that chosen direcion. The YAW angle, in our case, is the defined \u201dCommonDirecion,\u201d which is assumed to be the value of the optimal YAW anle ensuring direction change with minimum energy consumption. nce the desired direction is achieved, the UAV will move through he positive X-Axis, for a predefined slot of time, using the ROLL otation. ewDirection = Y AW angle herefore: ewDirection = CommonDirection g The change of direction after each a predefined constant period imulates the alternating electric current reversing direction peridically, with a regular time interval" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000390_j.protcy.2016.03.068-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000390_j.protcy.2016.03.068-Figure3-1.png", "caption": "Fig. 3. Profile of the master leaf spring in its free state and its major dimensions.", "texts": [ " The photographs of experimental set-up taken from front and right sides are shown in Fig. 1 and some measurement contrivances are marked as items A-D. The corresponding schematic diagrams shown in Fig. 2 (a-b) indicate the major sub-assemblies and some major dimensions of the set-up. Individual sub-assembly drawings are not furnished to maintain brevity. The kernel of the set-up is the master leaf of a LS bundle (item 8) which is mounted on a machined cast iron bed (item 10) by using roller supports at the ends (item 9). Profile of the master LS in its free state is shown in Fig. 3. In Fig. 3, cross-section and major dimensions (span, camber, eccentricity and arc length) of the master leaf are also shown and enlarged details of mounting drill hole is shown separately. It is apparent that the centre of the drill hole is not passing through the midpoint of span and the eccentricity between them is 1.6 mm. LS is loaded by placing slotted discs (item 4) over load connector (item 7). Horizontal movement of the load is restricted by the vertical guide rod (item 1). The rod is kept fixed in its position by the bush (item 2), which is fitted on the guiding disc (item 3) for centre line alignment" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002529_1350650121998519-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002529_1350650121998519-Figure1-1.png", "caption": "Figure 1. Schematic diagram of half of the groove-textured water-lubricated journal bearing (GTWJB).", "texts": [ " Third, the numerical results are discussed with respect to the rotary speed, eccentricity ratio, groove depth, width, and length ratio to evaluate the effects of groove textures on the loadcarrying capacity and friction force of a water-lubricated hydrodynamic journal bearing; furthermore, quantity comparisons between GTWJB and smooth bearing are carried out. The shapes of surface texture investigated in the present research include groove and rectangular. The schematic diagram of the GTWJB is illustrated in Figure 1. As the bearing is axial symmetry relative to the y\u2013O\u2013x plane along the z-axis, half of the bearing is modeled and analyzed in this study. Water is supplied into the bearing through the inlet and flow out from the axial sides of the bearing. The CFD simulation model of the GTWJB incorporates the following assumptions: 1. The journal and bearing surface are rigid. 2. The bearing is aligned with the journal. 3. A steady-state operation is considered. 4. The isothermal condition is assumed. In this study, the full Navier\u2013Stokes equations for 3D flow are solved for numerical studies", " The other parameter named friction ratio \u03b4f is similarly defined as the ratio of the difference of friction force between the textured and smooth bearings to the values of a smooth one. It is should be noted that the calculated load capacity and friction force below refer to half of the values of the whole bearing due to only half of the bearing being modeled and studied. The initial geometrical parameters of the GTWJB are illustrated in Table 2. The surface of the bearing is considered to be rigid. As shown in Figure 1, the groove circumferential spans are 30\u00b0 and distributed in the range of 210\u2013240\u00b0. There are 10 grooves arranged along the circumferential direction with an interval of 3\u00b0. The initial width of a groove is 0.1\u00b0. Figure 6 presents the pressure distributions of the GTWJB with respect to the rotary speed. As shown in the figures, with the increasing of the rotary speed, the negative zones increase. The Reynolds number ranges from about 654 to 3142 when the rotary speed increases from 5000 to 20,000 r/min for smooth and groove water film" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001827_0954406220908616-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001827_0954406220908616-Figure4-1.png", "caption": "Figure 4. Coordinate system mounted on point A0.", "texts": [ " There is a little difference between the analytical solution process for path generation with and without prescribed timing, that is in this solution, we let the point A0 (that must be chosen) to be the center of the coordinate system. The parameters and vectors used are the same as in the previous section. The algorithm of the presented article for path generation with prescribed timing is as follows: 1. Select the place of point A0, and calculate the loca- tion of P1, P2, and P3 relative to A0 by using A0P1 ! , A0P2 ! and A0P3 ! vectors (Figure 4) A0P1 ! \u00bc A0P1j jei l\u00f0 \u00de A0P2 ! \u00bc A0P1 ! \u00fe P1P2j jei \u00bc A0P2j jei 0 A0P3 ! \u00bc A0P1 ! \u00fe P1P3j jei \u00bc A0P3j jei 0 8>>< >>: \u00f07\u00de where A0P1j j is the distance between A0 and P1 and l is the angle of A0P1 ! with the global x-axis direction. Moreover, 0 and 0 are the angles of the A0P2 ! and A0P3 ! with the global x-axis direction, respectively. 2. For locating joint A1, one should solve a system of linear equations with two equations and two unknowns x and y (equation (8)) (based on steps 3 and 4 of section \u2018Graphical solution for path generation with prescribed timing\u2019)" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000093_j.engfailanal.2019.104180-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000093_j.engfailanal.2019.104180-Figure7-1.png", "caption": "Fig. 7. Location of overturning edges.", "texts": [ " It is in this area where the mining product is transferred between the excavator\u2019s and loading bridge\u2019s belt conveyors. The failure took place while the machine was changing location. At the moment of the incident, the crawler-mounted loading bridge was at higher level than the excavator (Fig. 5). The placement schematic of the loading bridge\u2019s elements is shown in Fig. 6. The loss of stability took place due to crossing the first overturning edge located between the axle of six a wheel bogie and the axle of a swing girder (Fig. 7). It means that the location of the center of gravity moved outside the area of the stability field. Overturning edges define the boundaries of the field of stability. Due to the loss of stability, the loading bridge leaned onto the second overturning edge located between the drive wheel of the fixed girder\u2019s track and the axle of swing girder. Fig. 8 present the crawler-mounted loading bridge directly after the incident. As a result of the loss of stability, the structural node of the loading bridge\u2019s support located on the undercarriage of the excavator was damaged, as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002870_s00500-021-06159-5-Figure11-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002870_s00500-021-06159-5-Figure11-1.png", "caption": "Fig. 11 Rubber bushing loading nomenclature (Goossens et al. 2017)", "texts": [ " Designs E and F are calibrated during the manufacturing process and adjusted if needed by expanding the inner sleeve. The preload on design G is added during assembly. Design H is an example of a rubber sleeve joint developed to withstand large cardanic (axis angular) loads and displacements while maintaining a high stiffness in the radial direction. In the example shown, an extruded hollow ball center sleeve is used to reduce weight (Hei\u00dfing and Ersoy 2015). Rubber bushings have different loading directions for each translational and rotational direction, as shown in Fig. 11. Stiffnesses of each direction are naming with their loading direction nomenclatures, radial, axial, torsional and conical/cardanic stiffnesses. The proposed CKH algorithm has been used in the shape optimization of the rubber bushing. Desired stiffness values of the rubber bushing are targets of the design, and they are listed in Table 12 The optimization problem of the rubber bushing is a constrained problem due to it has four targets. However, every target could be written as a constraint function" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002073_lra.2020.3013862-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002073_lra.2020.3013862-Figure1-1.png", "caption": "Fig. 1. The fail-safe flight of the T 3-Multirotor in a single-motor failure condition (red arrow indicating the faulty motor).", "texts": [ " To solve the yaw rotation problem, [7] adopted a tilt-rotor-type quadrotor platform with eight c-DoFs, where both translational and yaw motion were recovered after the thruster failure by utilizing the tilting function of the thrusters. From this research, we can see that the flight performance can be restored even with a quadrotor configuration if the platform has an internal mechanism that can increase the c-DoF of flight. In this paper, we introduce a new quadrotor fail-safe flight solution for the fully-actuated quadrotor platform called the T 3-Multirotor (the platform in Fig. 1). First introduced in [12] and [13], the T 3-Multirotor platform is a new 6-c-DoF aerial platform developed to overcome the dependency of fuselage attitude on translational motion. By utilizing the unique mechanical features of the T 3-Multirotor, in this paper, we introduce a new strategy to restore the controllable degrees of freedom by actively changing the platform\u2019s center of gravity position in a single motor failure scenario. The remaining of the paper is organized as follows. In Section II, the mechanism and dynamics of the T 3-Multirotor are introduced" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000076_j.colsurfa.2019.123843-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000076_j.colsurfa.2019.123843-Figure1-1.png", "caption": "Fig. 1. Schematic representation for formation of NiO nanoparticles and NiO-PANI nanocomposite.", "texts": [ " 98 \u03bcL of distilled aniline was added to the NiO nanoparticles dispersed in HCl solution and stirred for half an hour in an ice bath. APS solution (0.245 g APS in 50mL 1M HCl) was added dropwise to the above solution with continuous stirring and the solution was kept in stirring condition for 12 h. The resultant product was filtered followed by several times washing with distilled water and then dried at room temperature for characterization and adsorption studies. The yield product was 75% measured after drying (Fig. 1). Stock solutions of concentration 1000mg/L were prepared for both the pollutants (resorcinol and p-nitrophenol). 0.1 g of resorcinol and pnitrophenol was dissolved in 100mL distilled water respectively. These stock solutions were further diluted to prepare the test solutions of required concentration. The surface functional groups present in the nanoparticles were analyzed by Fourier-transform infrared (FTIR) spectorscopy. FTIR spectra were recorded using IR Affinity-1 spectrophotometer (SHIMADZU) at a scanning range over 4000\u2013500 cm\u22121" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001800_978-3-030-40410-9-Figure7.5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001800_978-3-030-40410-9-Figure7.5-1.png", "caption": "Fig. 7.5 RRP manipulator", "texts": [ " This ideal formulation is based on the assumption that we deal with rigid bodies and bearing. To give the reader significant insight, we have presented a problem, and all its parameters are calculated from the previous calculation which are path planning, trajectory planning, Jacobian, and inverse kinematics. Also, we have introduced MATLAB symbolic toolbox to drive the dynamic equations of the robot. This toolbox has significant benefits, especially when the designer would like to drive dynamic equations in a systematic and automatic way. Exercises Fig. 7.5, find the dynamic equations at the moment that the robot is at point (136,100). Assume any missing information that you may need. 7.2 For the robot in Problem 7.1, find the force history of the actuators assuming that the robot moves from (197,15) to (27,198) in 2 s by writing an appropriate program. Assume any missing information that you may need. 7.3 Assume that three DC motors are actuating the robot in Problem 7.1 and each one has a maximum speed of 500 rpm. What is the appropriate time to pass the trajectory from the start point (197,15) to the goal point (27,198), and is it more than 2 s or less than 2 s" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001312_s00170-020-05434-3-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001312_s00170-020-05434-3-Figure3-1.png", "caption": "Fig. 3 Depositing a single-layer variable width structure on the substrate. a The overall view. b The rudimentary model. c Longitudinal section", "texts": [ " The method proposed in this paper does not require complicated additional equipment. Variable width forming can be achieved using commonly used laser direct metal deposition equipment. It can not only save costs but also ensure forming accuracy. Although the method proposed by Shi et al. [16] can also achieve the purpose of variable width forming, it needs closed-loop control and additional equipment, which increases the cost to a certain extent. In variable width structures, the schematic diagram of a single layer is shown in Fig. 3. When the variable width forming is performed, the number of the tracks in a layer is same. The width of the cladding layer is changed by changing the overlapping rate. Scanning speed is simultaneously changed to ensure the uniformity of the height of the cladding layer. This section takes a variable width structure containing three tracks in a single layer as an example to analyze, but these models are not limited to three tracks. As shown in Fig. 3b, c, the scanning speed is gradually increased from the wide side to the narrow side, so the height of the single clad track is gradually decreased. But the overlapping rate is gradually increased, so the height of the narrow side cladding layer is gradually increased. Ultimately, the entire cladding layer is highly uniform. The smaller the width of the cladding layer, the greater the overlapping rate, and the corresponding cladding layer scanning speed is also increasing. However, to achieve the consistency of cladding layer height, it is necessary to find the optimal combination of the number of tracks, overlapping rate, and scanning speed" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001885_tvt.2020.2986395-Figure16-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001885_tvt.2020.2986395-Figure16-1.png", "caption": "Fig. 16. Iron core and PM eddy current loss distribution. (a) CP PM machine. (b) MSCP PM machine. (c) MSCP-AW PM machine. (d) MSCPAT PM machine.", "texts": [ " In order to analysis the influence of different MS structures on the iron core and PM eddy current loss, the loss Authorized licensed use limited to: University of Exeter. Downloaded on June 19,2020 at 02:47:19 UTC from IEEE Xplore. Restrictions apply. 0018-9545 (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. distributions of the analyzed machines are shown in Fig. 16, whilst the corresponding loss are listed in Table. IV. The loss distribution is corresponding to the flux density distribution on no-load, as shown in Fig. 9. The iron loss of the MSCP-AT PM machine is the largest due to the saturate in the AT, while its PM eddy current loss is the smallest. The MSCP-AW1 PM machine possesses the smallest efficiency due to its longer end-windings. The efficiency of the MSCPAT PM machine is the largest in the MSCP PM machines, which is only 0.49% lower than that of the traditional CP PM machine" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0003110_520237-Figure13-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0003110_520237-Figure13-1.png", "caption": "Fig. 13 is a typical tooth layout in the mean normal section. Since most gear engineers are familiar with the general method of construction, I will not review the process. The important thing to note is the method for locating the point of load application. Measure off along the line of action from its point of intersection with the addendum circle of the mating gear the distance pa. This gives point K'. Then swing an arc about the gear center passing through point K' and intersecting the gear profile at point K\". The remaining construction is carried out in the usual way using the tangent to the parabola and determining the distance XLN.", "texts": [ " and: Z r 2 [ 211 1~ Concave 5m0 J Convex (34b) The corresponding distance in the normal section (see Fig. 11) is p3. Then: _ V\\ cos \\pi, When mo is less than 2: p \\ I .. \" ' 0 \" ' 0 ZirlF \u00bb'o / L 2 2?/i,, 5mp When m0 is greater than 2: / m , \\ ! r ?\u00ab0 2 m.F \u00bbic P3 = ?2 I I \u00b1 \\ \u00bb'o / L 2??!p 5mp (35) ]Conc?ve (35 1Convex a Concave Convex (35b) where: cos 4 (cos2$ + tan2q5) Equations (35a) and (35b) are the same as those given in equation (2). APPENDIX F Layout Procedure The following data are necessary for the layout (Fig. 13) : tp, tG = Mean normal circular thickness of pinion and gear (tp + tG = pn) ap aG = Mean addendum of pinion and gear bp, bG = Mean dedendum of pinion and gear TTP, TTG = Tool edge radius of pinion and gear cutters RNP, RNG = Equivalent pitch radius in the normal plane 2 cos 7 cos21/- /. = Normal pressure angle Generally the layout is made to an enlarged scale, in which case the above linear values must be multiplied by a scale factor. It is quite common practice to use a scale factor of P. (P=diametral pitch) The layout is made in the usual way using RNp and RNG for the pitch radii" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000647_s0081543816080186-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000647_s0081543816080186-Figure2-1.png", "caption": "Fig. 2. (a) Full-scale model of a floating platform with two internal cams and (b) a trajectory of the model.", "texts": [ " In addition to the proof itself, the studies contain specific algorithms for controlling the motion of a body by means of internal mechanisms. In [59, 60], the authors also consider the issues of optimal control of the motion of bodies in a fluid by genetic algorithms. Note that experimental investigations qualitatively confirm the applicability of the model of an ideal fluid at least in the first approximation. At the Laboratory of Nonlinear Analysis and the Design of New Types of Vehicles, Udmurt State University, an experimental model of a watersurface platform with two internal cams was designed (Fig. 2). The experiments have confirmed the possibility of its self-propulsion and shown the qualitative agreement between the experimental data and theoretical results obtained within the model of an ideal fluid. However, for a quantitative description of self-propulsion, it would be interesting to evaluate the contribution of the vortex component, the effect of viscosity, and surface phenomena. Taking account of circulation of a fluid around the body. One of the methods of taking into account the viscosity of a fluid is the addition of vorticity to the model of an ideal fluid" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001061_mees.2019.8896485-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001061_mees.2019.8896485-Figure2-1.png", "caption": "Fig. 2. The basic kinematic scheme of the uneven rotation drive automatic moderator.", "texts": [ " 100 Nm at a variable injection advance angle ( 2 1). Furthermore, for all sections of plungers active stroke, this moment has one definite sign (here it seeks to reduce the and angles). Consequently, designed moderator operating conditions are similar to the operating conditions of the automatic injection advance angle change coupling, and therefore the design of this coupling was taken as the basis for this mechanism design. In the chosen variant of uneven rotation drive automatic moderator, the basic kinematic scheme of which is shown in Fig. 2, the centrifugal coupling 1 is rotated through a toothed rim of the drive plate 2 of a double-reduction camand-lever arrangement. On drive coupling 3 and driven half-coupling 4, satellites axles 5, 6 are fixed. Tooth gears 7, 8 are rigidly interconnected by shaft 9, which is simultaneously supporting for half couplings 3, 4. Differential gear 10 is fixedly attached to the framing and fixed against rotation, gear 11 is connected through gear wheels system (not shown in Fig. 2) to the gear sectors, cut at the ends of the cam plates through the system of gears (not shown in Fig. 2) connected with the gear sectors, cut at the cam plates ends. In this mechanism, the half coupling shift by a certain angle causes a proportional gear 11 rotation and a corresponding cam plates turn on the 1, 2 angle. By combining the stiffnesses and initiation of spacer springs installed in the centrifugal coupling, one can obtain the required (n) and (n) characteristics of moderator operating. The method for determining the and angles, providing the specified values of the j speed rotation unevenity rating and the injection advance angle, was described in [8]" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000113_j.finel.2019.103319-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000113_j.finel.2019.103319-Figure1-1.png", "caption": "Fig. 1. Straight cantilever beam test (L\u00bc 6.0; b\u00bc 0.2; thickness\u00bc 0.1; E\u00bc 1.0E7; \u03bd\u00bc 0.3; F1\u00bcF2\u00bcF3\u00bc 0.5).", "texts": [ " Both QUAD\u00fe and TRIA\u00fe shell element passed the zero energy mode test. In order to pass the patch test, calculated stress fields should be constant [41\u201343]. Both QUAD\u00fe and TRIA\u00fe shell elements passed the patch tests. Table 2 shows all the benchmark tests performed in the present research to compare the displacement results. Numerical solutions were normalised with respect to the benchmark solutions. The following subsections describe the test cases with related results. Although the straight cantilever beam (Fig. 1) was a simple test, it showed all the main issues encountered in structural finite elements analysis of shells (bending, shear, and twist). Loads were applied at the two vertices of the free edge. The analysed QUADmeshes (regular, trapezoidal, and parallelogram) are depicted in Fig. 1. TRIA mesh was obtained by splitting the related QUAD element along the diagonal side (dotted line in Fig. 1). To understand the sensitivity to mesh distortion, three different load conditions are adopted: The out-of-plane (F1 load) condition caused the bending of the beam, and therefore, the results showed that the out-of-plane behaviour of the shells for different meshes. For the in-plane (F2 load) condition, a unit load was directed along the transverse axis so that in plane bending and shear deformations aroused. This problem allowed to estimate the sensitivity to shear locking phenomenon and the effects of the distorted mesh" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002614_j.compositesa.2021.106449-Figure24-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002614_j.compositesa.2021.106449-Figure24-1.png", "caption": "Fig. 24. (a) The FE model of AP-Ply laminate (mesh not shown for clarification), (b) The FE model of magnified sample view cut from AP-Ply, (c) Cross-sectional view of the FE model.", "texts": [ " In this case study, a layup configuration of [0/90]13s was manufactured and used for demonstration purposes, as shown in Fig. 23a. A small sample was cut through the CNC Router and the crosssection was scanned under an optical microscope. The corresponding tow waviness and resin-rich pocket resulted from the wavy pattern are shown in Fig. 23c. The as-cured tow thickness was obtained by averaging the total laminate thickness. The crimp angle was measured from the micrographs (~50). These parameters were applied to the TWM algorithm for the generation of the FE model, as indicated in Fig. 24. The triangular resin-rich pockets and tow waviness with a user-defined crimp angle are explicitly captured. X. Li et al. Composites Part A 147 (2021) 106449 To demonstrate the efficiency of the TWM, a sensitivity test was performed in terms of the effect of the tow number, element size and geometric complexity on the computational cost. In this study, two different geometries were studied including a flat panel with dimension 200x200 mm and a pressure vessel with dimension 500x250mm (length \u00d7 diameter)" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000196_s13204-019-01219-7-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000196_s13204-019-01219-7-Figure1-1.png", "caption": "Fig. 1 Friction pair used in the friction test", "texts": [ " Infrared and Raman spectroscopy were measured via Nicolet iS10 infrared spectrometric analyzer (Thermo Fisher, USA) and LabRAM HR Evolution Raman spectrometer (HORIBA, France), respectively. A level-type four ball friction and wear test machine (MRS10P, Yihua Tribology Testing Technology Co., Ltd.) was applied to carry out friction test on grease samples. Four Gcr 15 steel balls (diameter of 12.7\u00a0mm, Rockwell hardness of 64 ~ 66) were used in the test. Three of them were fixed in the bottom, one was rotated on its axis in the upper side and contact to three bottom balls in point contact form, the test setup was shown in Fig.\u00a01. The friction test was conducted according to a standard GB3142-82 with conditions as follows: The balls rotated at a speed of 1200r/min under a fixed load of 392\u00a0N at a temperature of 75\u2103 for 60\u00a0min. Before the test, all steel balls were ultrasonically cleaned with alcohol for 15\u00a0min. All tests were repeated for three times at same conditions to ensure a reliable test data. The wear surface of steel balls was observed and investigated via metallographic microscope, 3D profilometer (NANOVEA ST400) and SEM (FEI Quanta 650 FEG), Raman spectrometer was used to identify the chemical component on wear surface" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001800_978-3-030-40410-9-Figure3.10-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001800_978-3-030-40410-9-Figure3.10-1.png", "caption": "Fig. 3.10 Rotating frame (2) \u221230\u00b0 relative to frame (1)", "texts": [ "30) In the same way, we can conclude the HTM for the short link as H2 = \u23a1 \u23a2\u23a2\u23a3 1 0 0 14.52 0 1 0 0 0 0 1 0 0 0 0 1 \u23a4 \u23a5\u23a5\u23a6. (3.31) The overall HTM of both of the two links is the serial multiplication of the two matrices: H = H1 \u00b7 H2 = \u23a1 \u23a2\u23a2\u23a3 1 0 0 34.74 0 1 0 0 0 0 1 0 0 0 0 1 \u23a4 \u23a5\u23a5\u23a6. (3.32) Indeed, if you measure the distance between points o3 and o1, it will be 20.22 + 14.52 = 34.74 with zero displacement in the direction of the y1-axis and z1-axis. Now, consider rotating frame o2x2y2z2 with (\u221230\u00b0) relative to frame o1x1y1z1, as shown in Fig. 3.10. The rotation in a counterclockwise direction is assumed to be negative while the rotation in a clockwise direction is positive; however, the reader can assume the reverse. The position vector of point o2 did not change while the rotation matrix changes to be like Eq. (3.10). Thus, HTM of the first link will be H1 = \u23a1 \u23a2\u23a2\u23a3 cos(\u221230) sin(\u221230) 0 20.22 \u2212 sin(\u221230) cos(\u221230) 0 0 0 0 1 0 0 0 0 1 \u23a4 \u23a5\u23a5\u23a6. (3.33) The origin point o3 does not change its relative position or its orientation with respect to frame o2x2y2z2. Thus, the rotation matrix between frame (3) and frame (2) is still the identity matrix with the same position vector as in Eq. (3.31). The overall HTM of both is H = H1 \u00b7 H2 = \u23a1 \u23a2\u23a2\u23a3 1 0 0 32.7947 0 1 0 7.26 0 0 1 0 0 0 0 1 \u23a4 \u23a5\u23a5\u23a6. (3.34) Figure 3.10 was drawn, and the distance between o3 and o1 was measured. It is 32.79 in x1-direction and 7.26 in y1-direction, the same as the results in Eq. (3.34). As shown in Fig. 3.10, when the coordinate frame (2) has rotated by 30\u00b0, the whole link with all of its points rotates with the same orientation that is the meaning of coordinate frame attached rigidly to a body. Translational and rotational movement of a point in space can be represented by a single 4 \u00d7 4 homogenous transformation matrix which holds both the position and rotation of a point. This matrix usually used to represent the spatial pose of a rigid body. Exercises R = \u23a1 \u23a3 0.2962 \u22121.0680 0.5335 0.3379 0.0517 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002356_978-3-030-50460-1-Figure20.5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002356_978-3-030-50460-1-Figure20.5-1.png", "caption": "Fig. 20.5 8 Degree of freedom dynamic model", "texts": [ "1 is evaluated for different spalling fault sizes as given in Table 20.2. The gear-mesh stiffness was evaluated for healthy and tooth having spall faults for the gear parameters tabulated in Table 20.1 and for different spalling size as given in Table 20.2. The gearmesh stiffness for the one revolution of pinion is represented for healthy and faulty gear are represented in Figure 20.3 and 20.4 respectively. For obtaining a vibration response a dynamic model developed by Bartelmus (2001) having both torsional and lateral motions have been utilized (Figure 20.5). It is a 8 degree of freedom mass-spring-damper dynamic model with M1 input torque by 368 Handikherkar, Phalle motor and M2 output torque. Two flexible coupling is used one to connect the motor with the input shaft carrying pinion and second to connect shaft carrying gear and load. The pinion and gear shafts are mounted on the rolling element bearings and these bearings are then mounted on the gearbox casing. The equations of motion are given as follows (Equations (20.10)\u2013(20.17)): m1y\u03081 + c1y\u03071 + k1y1 = \u2212Fk \u2212 Fc (20" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001034_j.mechmachtheory.2019.103681-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001034_j.mechmachtheory.2019.103681-Figure4-1.png", "caption": "Fig. 4. Coordinate systems for processing conical worm wheel.", "texts": [ " (5) , (27) , (28) , and (29) , it is easy to attain the normal curva- tures, k ( l\u22122 ) \u03be and k ( l\u22122 ) \u03b7 , of the workpiece l \u2212 2 helicoid (S) l\u22122 along its tangent vectors ( g (l) 1 ) o ,l\u22122 and ( g (l) 2 ) o ,l\u22122 respectively, and its geodesic torsion, \u03c4 ( l\u22122 ) \u03be , along ( g (l) 1 ) o ,l\u22122 as follows k ( l\u22122 ) = \u2212 g 2 lz , k ( l\u22122 ) \u03b7 = k ( l ) 2 \u2212 \u03bc2 l ,l \u22122 , \u03c4 ( l\u22122 ) = \u03bcl ,l \u22122 g lz . (30) \u03be l ,l \u22122 l ,l \u22122 \u03be l ,l \u22122 2.3. Generation of conical worm wheel and its cutting mesh parameters Fig. 4 shows all the coordinate systems related to the conical hob 4 and the conical worm wheel 2 during processing the conical worm wheel. A static coordinate system \u03c3o2 { O 2 ; i o2 , j o2 , k o2 } and a rotating coordinate system \u03c32 { O 2 ; i 2 , j 2 , k 2 } related to the worm wheel are used to depict its original position and current position, respectively. Both the unit vectors k o2 and k 2 are coincident with the centerline of the worm wheel and point to its big end. The vector i o2 coincides with the common perpendicular of k o4 and k o2 " ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000539_s00170-016-9337-2-Figure16-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000539_s00170-016-9337-2-Figure16-1.png", "caption": "Fig. 16 Adjust two regions using \u03bb", "texts": [ " We introduce the parameter \u03bb , (0\u2264\u03bb\u2264 1) to adjust the ratio of the Zig Area. When\u03bb= 0, the entire tool path is morphing type; when \u03bb= 1, the dividing line is generated using the algorithm above and the ratio of Zig Area reach the maximum; when 0 <\u03bb< 1, the dividing line is an offset line from PdivQdiv toward the turning corner. His the distance between the turning corner on the outer boundary and the original dividing line PdivQdiv; the new dividing line Pdiv 0 Qdiv 0 is the offset line of PdivQdiv with a distance of (1\u2212\u03bb)H, as shown in Fig. 16. In each Zig Area, the projection of the dividing line on the platform surface is also the first zig/zigzag tool path, the other zig/zigzag tool paths are geodesic offset lines (with a distance of step-over)of the dividing line. In this research, we use the iso-curves of the optimized virtual potential field to compute the geodesic offset lines. The energy value of the vertices on the dividing line is fixed to be zero, and the potential values of the other vertices are computed through the optimization process" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002145_00405000.2020.1819007-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002145_00405000.2020.1819007-Figure6-1.png", "caption": "Figure 6. Intersection points in Sample S4 and S6.", "texts": [ " Sample SP6 showed minimum auxeticity because there is a significant difference in the float of warp and binder yarns; there are maximum intersection points in one repeat of sample SP4. When the tensile load was applied, binder yarns tries to increase the thickness. Still, due to maximum warp intersection points, binder yarn faces the opposition force and, as a result, hinders an increase in thickness. This force cancels out phenomena, and the number of intersection points in both samples is expressed in Figure 6. The values of the initial thickness of reinforcement (without load) and final thickness (with load) are expressed in Figure 6. Figure 7 and Figure 8 show the Poisson\u2019s ratio of all reinforcement. Figure 9 shows the plot of the axial and transversal strain of 3D reinforcements. It can be seen in Figure 9 that the axial strain does not vary much and is almost constant. It can be concluded from the above discussion that as the float length of binder yarn and warp yarn increases with the equal and same ratio, the Poisson\u2019s ratio also increases. As the float length of binder yarn decreases, the auxetic behavior also decreases" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002478_s40194-020-01061-4-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002478_s40194-020-01061-4-Figure2-1.png", "caption": "Fig. 2 a Schematic diagram of the TIG welding V-type groove configuration and b weld pad", "texts": [ " The machined coupons were buffed and cleaned with acetone to eliminate stains, surface impurities and clamped on to 25- mm-thick copper backing plates. The activated flux was centrifuged by a multicomponent flux with acetone. A few microns level thick activated flux was coated on the superficial surface of the weld joint prior to welding. The coating was left to dry for a few seconds. Then, autogenous TIG welding was undertaken on the surface. Edge preparation with a 70\u00b0 angle, 2-mm root gap and 1-mm root face for multipass TIG weldment configuration is shown in Fig. 2 (a) schematic diagram of the TIG welding V-type groove configuration and (b) weld pad. ERNiCrMo-3 was used as the filler material and its chemical composition was determined with an optical emission spectroscope (OES). The results are presented in Table 3. Welding process parameters used for the TIG weldment are given in Table 4. The welds\u2019 transverse sections measuring 30 mm \u00d7 10 mm \u00d7 6.5 mm were metallographically characterized to cover the base metal, HAZ and the fusion zone. The sample\u2019s surface was prepared to ensure a good surface finish using procedures like disc polishing with several emery papers of 80 to 2000 \u03bcm thickness" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002285_j.jhydrol.2020.125812-Figure9-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002285_j.jhydrol.2020.125812-Figure9-1.png", "caption": "Fig. 9. Variation of D with \u03b6 and \u03c4 for Pe = 1.0, \u03b1 = 1.0, a1 = 10, Tr = 1.0 and \u03d6 = 1.0. The colour bar represents the values of D .", "texts": [ " According to Eqs. (37)\u2013(39), the second moment is related to a1, \u03b1, Pe, and Tr. Since the complexity of the zeroth and first moments, it is quite difficult to get the explicit expression of m2(\u03b6, \u03c4). In this work, the finite difference method with time marching of Euler scheme is used to solve Eqs. (37)\u2013(39) and then to explore the effects of dimensionless parameters on the longitudinal dispersivity which can be calculated as (Aris, 1956; Yang et al., 2020) D ( \u03b6, \u03c4 ) = 1 2 d d\u03c4 ( m2 m0 \u2212 m2 1 m2 0 ) . (40) Fig. 9 presents the variation of D with \u03b6 and \u03c4 for Pe = 1.0, \u03b1 = 1.0, a1 = 10, Tr = 1.0, and \u03d6 = 1.0. It is shown that the longitudinal dispersivity cannot reach an asymptotic value as it does in the fully developed wetland flow. Instead, D would reach a stable oscillatory status in the tidal flow through a wetland. The effect of initial M.Y. Guan et al. Journal of Hydrology 593 (2021) 125812 distribution of active particles on the longitudinal dispersion is mainly limited to the first oscillatory period" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000718_978-1-4471-7278-9-Figure9.2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000718_978-1-4471-7278-9-Figure9.2-1.png", "caption": "Fig. 9.2 The boost-discharge circuit", "texts": [ " Both are voltage- and time-dependent, which creates a rich repertoire of computational primitives. Each of these primitives has the ability to implement a variety of linear and non-linear, high-gain operations. 2Koch C (1999) Biophysics of Computation: Information Processing in Single Neurons. Oxford University Press. 1999 3Cole KS (1972) Membranes, Ions and Impulses. University of California Press: Berkley, CA, 1972 168 9 Euphoria Dynamics Assume the membrane resistance (R) can be ignored. We can configure a boostdischarge circuit as shown in Fig. 9.2. This is similar to widely-used circuits for power voltage regulation, and camera flash charging and discharge. The underlying principle of the camera flash is a charge pump circuit built using a capacitor and a boost circuit discharging into a bulb filled with xenon gas. For a very bright flash, xenon atoms must be excited to high energy levels. This requires the electrons to be imparted with very high energies. To accomplish this, the terminal voltage across the tube must be raised to high values", " Thus, the discharge equation is: dvC dt D i C D vC R2C (9.7) The equations from 9.6 through 9.7 can be solved by physical circuit simulation or numerical computation. 170 9 Euphoria Dynamics We can simulate boost firing with an electronic circuit, but a few adjustments are necessary. The physical Switch 1 can be replaced with a transistor for fast pulse generation. When the transistor is in cutoff mode and no current flows through it, then all the components are in series, which is the situation when Switch 1 is at the B position in Fig. 9.2. When the transistor is in saturation mode, then it becomes a short circuit. After flowing through the resistor and inductor, the current is shorted straight to ground, bypassing the components to the right of the transistor. This is the same situation when Switch 1 is at the A position. This experimental physical circuit can be built with parts commonly available at electronic hobby stores such as RadioShack in the US, including an 82 mH inductor, an 82uF capacitor, a 0.3 V Schottky diode, and a standard NPN transistor" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0003120_ac50132a017-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0003120_ac50132a017-Figure2-1.png", "caption": "FIGURE 2. CARBON DRILLING AND POINTING TOOL", "texts": [ " This ensures a rapid cleanup of traces of the ash residue when the wall has burned down to the floor. In spectrographic determination of mercury in plant tissue the authors found that a more intense mercury line is produced a t 2536.7 A. by using an unusually deep electrode cavity (about 15 mm.). With the use of 8-mm. electrodes, pointing the wall a t the top serves to prevent arc wandering during the short exposure needed to volatilize the mercury present. A tool designed to perform the operations of drilling and pointing, simultaneously, is shown in Figure 2. It ANALYTICAL EDITION 219 APRIL 15,1939 consists of a standard 6-mm. twist drill and a steel collar with a set screw on which projects a cutting edge a t a suitable angle (about 20\"). The depth of the cavity here again can be varied by merely sliding the collar along the drill and tightening the screw. Also, a good drill press combined with an electrode-holding device (chuck, collet, etc.), attached firmly to the bed or base plate, can be substituted where a lathe is not available. For successful production of this thin-walled cavity, both tool and carbon electrode must be held immovable as in a chuck or collet" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002450_j.compstruct.2021.113608-Figure9-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002450_j.compstruct.2021.113608-Figure9-1.png", "caption": "Fig. 9. The contact models between", "texts": [ " The element type used was C3D8R and the total number of elements of any contact model was about 0.1 million. The deformation process of the compaction roller is simulated by applying a fixed constraint on the mould surface and a concentrated force on the mandrel. The simulation process starts when the compaction roller has just touched the curved surface, and applies the specified force until the deformation of the compaction roller ends. According to the deformation modelling method of the compaction roller in the previous section, the contact models as shown in Fig. 9 were established for the three main contact situations to analyze the maximum deformation of the roller under different laying forces. The dimensional parameters of the compaction roller and the mould surface are shown in Fig. 10. There is no deformation of the compaction roller mandrel and the mould, only the silicone rubber layer of the compaction roller is deformed, and the maximum deformation is equal to the displacement of the mandrel. Therefore, the maximum deformation of the compaction roller can be obtained by looking at the mandrel displacement under different forces" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001846_j.mechmachtheory.2020.103877-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001846_j.mechmachtheory.2020.103877-Figure6-1.png", "caption": "Fig. 6. Free-body diagrams of the transmission parts. Single arrows are forces, double arrows are torques.", "texts": [ " (7) and (10) gives the four spin coefficients are derived as functions of the tilt angle and the creep coefficients: \u03c3S0 = \u03c9 S0 \u00b7 z 0 | \u03c9 0 | = \u2212\u02dc r0 tan (\u03b80 + \u03b3 )(1 \u2212 C R 0 ) + sin \u03b80 (11) \u03c3Sa = \u03c9 Sa \u00b7 z a | \u03c9 a | = \u2212 sin \u03b1 \u2212 \u02dc ra tan (\u03b1 \u2212 \u03b3 ) (1 \u2212 C Ra ) (12) \u03c3Sb = \u03c9 Sb \u00b7 z b | \u03c9 b | = \u02dc rb tan (\u03b1 + \u03b3 ) (1 \u2212 C Rb ) + sin \u03b1 (13) \u03c3S2 = \u03c9 S2 \u00b7 z 2 | \u03c9 2 | = \u02dc r2 tan (\u03b82 \u2212 \u03b3 ) (1 \u2212 C R 2 ) \u2212 sin \u03b82 (14) where the unit vectors of z-axes are defined as follows: z 0 = \u2212 sin \u03b80 j + cos \u03b80 k z 1a = \u2212 sin \u03b1j \u2212 cos \u03b1k z 1b = sin \u03b1j \u2212 cos \u03b1k z 2 = sin \u03b82 j + cos \u03b82 k Spin coefficients are purely kinematic quantities. Spin coefficients can be positive or negative, and they are here defined in such a way to be coherently related with the spin torques defined in the next section and shown in Fig. 6 . Fig. 5 shows the spin coefficients as functions of the tilt angle. Calculations are performed assuming \u03b8 = \u03b80 = \u03b82 , k = 1 and no creep. Three values of \u03b80 are considered which correspond to different speed ratio ranges. It can be observed that larger speed ratio ranges correspond to larger maximum values of the spin coefficients. Forces and torques equilibrium equations of each part of the variator are derived from the free-body diagrams shown in Fig. 6 . A number n of equal rollers is considered. F D 0 and F D 2 are the clamping forces applied to input and output discs; in the free-body diagram, clamping forces are divided by n , considering the amount of them applied to each single roller. A similar argument can be discussed for the axial component of the reaction force on the inner rollers F Da and F Db . These are the axial component of the reaction forces of bearings. Rollers are distributed in such a way to make the outer and inner discs and the carrier radially self-equilibrated, thus radial reaction forces are zero" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000200_s00202-019-00882-x-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000200_s00202-019-00882-x-Figure3-1.png", "caption": "Fig. 3 General arrangement drawing of the coaxial magnetic gear design", "texts": [ " Permanent magnets\u2019 permeability is important for reluctance estimation. Materials such as AlNiCo, ferrite, SmCo and NdFeB are utilized in order to estimate the outer torque transmission reduction during MG performance. Even though reluctance is determined under static conditions, the magnetic permeance in the air gaps depends on rotors\u2019 relative positions and must be calculated according to different rotors\u2019 alignment configurations. A depiction of the general arrangement of the CMG under consideration is shown in Fig.\u00a03. At present, it is one of the higher-specific torque\u2013density ratio designs [14]. It has 4 inner permanent magnet pole pairs, 22 outer permanent magnet pole pairs and 26 modulating steel segments. The rated rotational speeds of the inner rotor and the outer rotor 1 3 are 150\u00a0rpm and 27.3\u00a0rpm, respectively, modulating steel segments are fixed to the housing, and the gear ratio is 5.5. The dimensional drawing of the CMG design is shown in Fig.\u00a04, and dimensions are presented in Table\u00a01. Magnetic circuit is optimized for NdFeB 35 permanent magnets [20]" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002444_j.measurement.2021.109021-Figure12-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002444_j.measurement.2021.109021-Figure12-1.png", "caption": "Fig. 12. Health status of the intermediate speed stage components after 345 h of run time (a), (b) & (c).", "texts": [ " Also, the A4 channel had yielded second best favourable results in the HSS stage and could distinguish between the three health conditions zones of the gearbox as seen in Fig. 10. Fig. 11 shows the \u2018R2\u2019 value for the above mentioned cases of A2 and A4. It can be noted that, the A2 channel has yielded highest \u2018R2\u2019 values for all the three LSTM network models. Also, the A2 channel run-to-failure vibration data has shown better precision among all the different vibration axis data. The ISS (planetary 2) consisted of six 6004 2RSH bearing, three planet gears, one sun gear, one ring gear and two SYJ25TF bearings. Fig. 12 (a) illustrates the ISS stage at the end of 345 h of continuous runtime. It was noticed that, the ISS had taken moderate damage when compared to the other stages of the wind turbine gearbox. The ISS was in a partial operational condition. Pitting damage was observed on all the gears and oil seal of the multiple 6004 2RSH bearings failed. Besides H.M. Praveen et al. Measurement 174 (2021) 109021 these damages, there were no symptoms of any major damage in the bearings (SYJ25TF) on the transition region between ISS and LSS, refer Fig. 12 (b) & (c). However, excessive play could be noticed on the planet gear axis and metal fragments were also observed between the gears which indicated the occurrence of severe abrasion. Fig. 13 describes the predicted vs actual health condition for the ISS stage using A3 at scale 1 (A3-1-1, A3-1-2 & A3-1-3) and A4 scale 7 (A4-7-1, A4-7-2 & A4-7-3). It was observed that, the prediction capability for channel A3 is better than A4 during this stage. This can be due to the fact that, the A3 sensor was located closer to the ISS stage and the A4 sensor was located a little farther from the stage" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000338_j.1538-7305.1960.tb01602.x-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000338_j.1538-7305.1960.tb01602.x-Figure6-1.png", "caption": "Fig. 6 - Transfluxor, its mirror-symbol equivalent, and flux patterns.", "texts": [ " In a sense, the laddie can be considered as consisting of a number of toroidal cores interconnected in series by magnetic linkages, while in the transfluxor the toroidal cores are interconnected in parallel. The magnetic linkages make possible a transfer of flux from one core to another without the use of a coupling winding. The latter, because of added resistive losses, reduces the effi ciency of circuits. A modified mirror symbol representation will be used to describe circuits made with multiapertured cores. The procedure is illustrated in Fig. 6 for a two-hole transfluxor. Each leg of the core shown in Fig. 6(a) is represented by a heavy vertical line in Fig. 6(b), the notation at the bottom giving the flux capacity set by the minimum cross section INTEGRATED MAGNETIC CIRCUITS FOR LOGIC MACHINES 301 of the corresponding leg. The horizontal striated lines are included simply to draw attention to the fact that the legs are connected mag netically, so that a flux reversal in one must be balanced by an opposite reversal in one or both of the other two in order to maintain flux con tinuity. The preferred flux return will be through the closest available path", " the result of setting flux up in leg 2 will be to induce an emf in the wind ing ab in a direction to drive current from terminal a through the winding to terminal b. This notation simplifies the explanation of the more complicated circuits. It is often helpful in describing the operation of these structures to represent the state of magnetization graphically by means of closed flux patterns. As an example, three of the possible states of magnetization of the transfluxor are shown in Figs. 6(c), (d) and (e). In this case, each of the flux arrows represents the saturation remanent flux capacity of each of the smaller legs. Thus, in Fig. 6(c) all legs are fully saturated; whereas in Figs. 6(d) and (e) legs 2 and 3 only are saturated, leg 1 hav ing zero resultant magnetization-which, it should be remarked, is not arrived at by a sinuosidal demagnetization, but by a partial reversal from a saturated state. A further subdivision of the flux patterns may be introduced if necessary, provided that the condition of flux continuity is maintained. It should be kept in mind that such flux patterns are merely a convenient fiction, and that the actual domain structure will be considerably more complicated. Nevertheless, the model is helpful and adequate for most purposes. It is usually sufficient to represent the state of magnetization of the individual legs by upwards or downwards pointing arrows, each representing one unit of flux, as in Fig. 6(b). Frequently, a multiapertured core can be replaced by an equivalent toroidal core circuit.\" For example, the transfluxor shown in Fig. 6 can be replaced by the circuit shown in Fig. 7. The three magnetically linked legs of the transfluxor are replaced by three electrically coupled toroidal cores having appropriate flux capacities and thresholds. If the coupling INTEGRATED MAGNETIC CIRCUITS FOR LOGIC MACHINES 30::1 winding has zero resistance, then flux may be transferred from core 1 to core 2 by II , and then from core 2 to core 3 by a current-limited drive /2 ,as hetween the corresponding legs of the transfluxor. In general, however, the operation of the multiapertured core equivalent is simpler to explain", "\" CLEAR \"0\" CLEAR tt ~ ~ tt ~ ~ U t ~ tt ~ ~ tt ~ ~ tt ~ ~ (b) '0\u00b7 CLEAR \"I-PRIMEO\u00b7 CLEAR '0' CLEAR tt ~ ~ tt ~ ~ n ,t tt ~ ~ tt ~ ~ tt ~ ~ (c) CLEAR \"0' CLEAR \\1,\" CLEAR \"0\" (d)tt ~ ~ tt ~ ~ tt ~ \u2022 tJ t ~ tt ~ ~ tt ~ ~ (e) 6 2 1 43 <1 2 2 I-- - ~ -- - ~ -- - ~ -- - ~ -- - ~ -- - 2 -- - -- - -- - -- - -- - -- .- P AI A2 E, E2 CORE NO, Fig. 10 - (a) Skeleton shift register circuit; (b) initial flux setting; (c) primed flux setting; (d) advanced flux setting; (e) complete circuit. fluxor structure shown in Fig. 6. The framework of the shift register is shown in Fig. 10(a). Leg 3 of each core is coupled by a closed winding to leg 1 of the following one. The resistance shown in Fig. 10(a) is the resistance of the winding itself, normally a few tenths of one ohm. A turns ratio is used to give the flux gain necessary to make up for transfer losses.* In practice, it is convenient to make n = 2. The flux setting of each leg will be represented by arrows in Figs. lO(b), (c) and (d), each arrow representing one unit of flux (cfI)", " 10(b) the odd-num bered cores 1 and 5 are shown set in the \"zero\" state, core 3 being set \u2022 It should be possible to use a unity turns ratio in the coupling loop by taking advantage of the flux gain mechanism produced by a re-entrant B-H characteris tic as mentioned in Section II. However, the materials property necessary is somewhat critical, and as a result the alternative complication of a turns ratio is preferred at present. Furthermore, the flux gain is then sufficiently large that the allowable tolerances are improved. INTEGRATED MAGNETIC CIRCUITS FOR LOGIC MACHINES 309 in the \"one\" state. The corresponding flux patterns are shown in Fig. 6(c) and Fig. 6(d) respectively. These two states correspond to the \"blocked\" and \"unblocked\" states of a transfluxor.? The \"clear\" state also corresponds to the \"blocked\" condition. Observe that it has been possible to change the flux pattern to insert a one into core 3 without producing a flux change in leg 3, which would induce an emf in the winding coupling core 3 to core 4. Therefore, it was possible to effect a full flux transfer from core 2 to core 3 without including the diode, which was required in the conventional toroidal-core circuit discussed in con nection with Fig", " The bit information may be transferred from the odd-numbered cores to the following even-numbered cores, leaving the former in the \"clear\" state, as follows: The information cores are first \"primed\" by applying a drive to leg 3 of every core in a direction to switch flux upwards. This is the P (prime) drive shown in the complete circuit, Fig. lO(e). It is limited in current amplitude so that although it can switch flux between legs 2 and 3, that is, around the small hole of the transfluxor, it cannot produce a flux reversal around the longer path including legs 3 and 1. The P drive will not affect the zero or \"clear\" cores. However, in a core which contains a one it will set leg 3 up and leg 2 down. In other words, the flux pattern is changed from that shown in Fig. 6(d) to that shown in Fig. 6(e). As leg 3 is being set up, a current will be induced in the winding coupling cores 3 and 4. However, the direction of this current is such that it does not change the flux setting of core 4. The final setting of all cores follow ing the P phase is shown in Fig. 1O(c). Note that it was at this point that advantage was taken of the threshold characteristic of the core, in this case to prevent the improper conversion of a zero into a one. The closest corresponding situation in the conventional toroidal-core circuit was the adverse backwards propagation of a one" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000667_iris.2016.8066074-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000667_iris.2016.8066074-Figure1-1.png", "caption": "Fig. 1. Kinematic chain of ABB IRB 6640", "texts": [ " Robot calibration is the process of finding and modifying these error parameters to compensate corresponding errors. The Denavit-Hartenberg Model (DH model) [15] is the most used representation in related works where Cartesian coordinate frame is attached to each joint of robot according to such rules: the Z-axis is along the axis of joint; the X-axis is along the common perpendicular between two adjacent Z-axes. In accordance with these rules, the kinematic chain of the ABB IRB 6640 robot is built and shown in Fig. 1. Homogeneous transformation matrix can be used to represent the relationship of two adjacent joints, and this matrix is determined by four independent parameters named as id , ia , i , i . For robots composed of rotary joints, id , ia , i are fixed and i represents the rotation angle of joint i. The homogeneous transformation matrix of joint i-1 and i can be written as: 1000 0 A 1 iii iiiiiii iiiiiii i i dcs sascccs casscsc (1) where 1A i i is the homogeneous transformation matrix; ic and ic represent icos and icos ; is and is represent isin and isin " ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000976_s11044-019-09704-1-Figure15-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000976_s11044-019-09704-1-Figure15-1.png", "caption": "Fig. 15 Radial displacement of ball bearing", "texts": [ " Because it is relatively simple and provides a fairly accurate load distribution in the rolling elements, bearing stiffness, and fatigue life, the quasi-static bearing model is often adopted in the design of rolling-element bearings. Considering zero clearance and a simple radial load, Stribeck [43] conducted one of the first studies for the static ball bearings. This study became the basis of (International Organization for Standardization) ISO 76:1978 [44] which specifies the methods for calculating the basic static load. For a rigidly supported bearing subjected to a simple radial load showed in Fig. 15, the radial deflection at any ball angular position is given by \u03b4\u03c6 = ur cos\u03c6 \u2212 Pd/2, (26) where ur = \u03b4max + Pd/2 is the ring radial displacement, occurring at \u03c6 = 0, and Pd is the diametral clearance; \u03b4max is the maximal radial deflection, occurring at \u03c6 = 0. Equation (26) may be rearranged in terms of maximum deformation as follows: \u03b4\u03c6 = \u03b4max [ 1 \u2212 1 2\u03b5 (1 \u2212 cos\u03c6) ] , (27) where \u03b5 = 1 2 ( 1 \u2212 Pd 2urc ) . (28) According to Eq. (12), Q\u03c6 Qmax = ( \u03b4\u03c6 \u03b4max )1.5 . (29) Therefore, from Eqs. (27) and (29), Q\u03c6 = Qmax(\u03b4\u03c6/\u03b4max) 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0003120_ac50132a017-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0003120_ac50132a017-Figure1-1.png", "caption": "FIGURE 1. CARBON DRILLING AND CUTTING TOOL DRILL BIT", "texts": [ "cavity floor, which is essential if the last trace of fused ash is to be burned off. Too thick a wall also unduly increases background on the plate. Since in the course of analytical routine, a large number of electrodes are prepared, it is desirable that facing the end of the electrode, drilling the hole, and cutting down the outside wall be accomplished in one operation. A tool was made that answers the above specifications. It was designed to produce craters of uniform wall thickness but variable depths by an adjustment of the set screw as is shown in Figure 1. By machining the outside of the cavity wall, uniformity of thickness is ensured, so that irregularities in graphite electrodes as purchased become unimportant; a long outside cut of 9.5 mm. when the cavity depth is only 3.5 mm. has the advantage of the smaller diameter 6-mm. (0.25-inch) electrode. It will be noted from Figure 1 that the angle on the cutting edge of the bit has been reduced, so that the depression in the floor of the cavity is comparatively slight. This ensures a rapid cleanup of traces of the ash residue when the wall has burned down to the floor. In spectrographic determination of mercury in plant tissue the authors found that a more intense mercury line is produced a t 2536.7 A. by using an unusually deep electrode cavity (about 15 mm.). With the use of 8-mm. electrodes, pointing the wall a t the top serves to prevent arc wandering during the short exposure needed to volatilize the mercury present" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002393_jrproc.1954.274823-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002393_jrproc.1954.274823-Figure7-1.png", "caption": "Fig. 7-Graphical solution of (19); upper curve for 4 = - (7r/8) and H=K/2j.", "texts": [ " (17) 7 -r/4 7r a. When a field H is applied externally in the y direction the domains forming an angle q5 with the z axis will E = K sin2 0 COS2 6 - HJ sin (6 + 4) (18) since the angle which the domain forms with the applied field is 90- (6 +k). Equating to zero the derivative of E with respect to 0, we obtain 2HJ sin 40= cos (6+4). K (19) Equation (19) determines 0 as a function of 4 for a given H. The solution can be found graphically from the intersection of the two curves on each side of this equation (Fig. 7). The smaller root 01 determines the minimum of E and hence the angle by which J rotates when H is applied. It can be seen that for K H < 2J 01 is smaller than ir/8, for every 0; thus the domains remain closer to the crystal axis along which they were directed before H was applied. When H is removed, they return to their original position and the magnetization is fully recovered. For 1954 1285 4 - '- 4 PROCEEDINGS OF THE I-R-E K 2J the curves sin 40 and 2HJ/K cos [6- (r/8)] are tangent. For K 4 K - = R\n(67)\nr\n\u03c32 \u03c30 =\n{\n0 r < R\n2 C 1 a 2 / b 2 + 1 \u2212\n\u221a\nC 2 1 + M\n2\nr 2 r > = R\n(68)\n\u00af =\n3 M\u0304\n8 ( 1 \u2212 a 2\nb 2\n)[ 1 R\u0304 2 \u2212 1\nB\n+ ln\n( B\nA\n) + 1\nR\u0304\n2 \u2212 8\n3\n( a\nb\n)2\n+\n5\n3\n] (69)\nhere\n\u00af =\n3 E\u03d5\n4 \u03c30\n( 1 \u2212 a 2\nb 2\n) (70)\n.2.1. Verification of accuracy\nIn this sub-section, the annular membrane is analyzed with he new wrinkling model and the results are compared with he theoretical solutions obtained from Eqs. (63) to ( 65 ), ( 67 ) o (69). The finite element model is shown in Fig. 20 . Its outer adius b and inner radius a (shown in Fig. 19 ) are 1250 mm nd 500 mm, respectively. The thickness t is 1 mm. The Young\u2019s\nodulus E = 10 0 0 MPa and the Poisson\u2019s ratio \u03c5 =1/3. The inner dge is restrained in the radial direction and the circumferential OFs of nodes on the edge are coupled to have the same torsional otation. The outer edge is firstly pretensioned to introduce a", "u c t\nr r a m p M i a s c b p r o a p a\nw fl v t\n6\nm ( 0\n8 As for the numerical examples in Section 6.2 , their geometrical and physical parameters are the same as given in Section 6.2.1 except the quantities assigned specifically.\nniform radial stress \u03c3 0 and then fixed along the radial and ircumferential directions. With a twisting moment M applied to he inner edge, wrinkles arise and their radii increase gradually.\nThe relation between the twisting moment and the torsional otation is given in Fig. 21 . In the figure, the horizontal axis repesents the dimensionless torsional rotation (please see Eq. (70) ) nd the vertical axis denotes the dimensionless twisting mo-\nent (please see Eq. (66) ). Contours of the major and minor\nrincipal stresses and the wrinkling strain corresponding to\n/ (2 \u03c0\u03c30 a 2 t) = 2 . 9 in Fig. 21 are illustrated in Figs. 22\u201324 , where t can be found that the principal stresses and the wrinkling strain re all uniformly distributed along the circumferential direction. It ignifies that the variation of stresses and strains along the radial oordinate may be typical of their distribution over the mem-\nrane. Fig. 25 illustrates the distributions of the major and minor rincipal stresses along the radial direction with R\u0304 = 1 . 2 and 1.6 espectively. The horizontal axis in the figure represents the ratio f the radial coordinate r to the inner radius a and the vertical xis denotes the major or minor principal stresse ( \u03c3 1 or \u03c3 2 ) to the restress \u03c3 0 . Figs. 21 and 25 indicate that the wrinkling model is ccurate and the results agree well with the theoretical solutions.\nFig. 21 also manifests that the major principal stress in the inkled region is much larger and its distribution curve becomes at in the taut area. Conversely, the minor principal stress almost anishes in the wrinkled region while it increases gradually with he radial coordinate in the taut zone.\n.2.2. Influence of the prestress\nThe annular membrane shown in Fig. 19 subjected to a twisting\noment M = 10,0 0 0 N \u2022mm is analyzed 8 with different prestresses \u03c3 = 0 . 1 \u03c4, 0 . 2 \u03c4, 0 . 3 \u03c4, 0 . 4 \u03c4 and 0.5 \u03c4 , where \u03c4 is the uniformly", "o u m\n\u223c s i r\ndistributed shear stress along the circumferential direction at r = a and obtained from the boundary condition with \u03c4 = M\n2 \u03c0a 2 t ).\nThe influences of the prestress on the major and minor principal stresses and the wrinkling strain are illustrated in Fig. 26 (a) and (b), respectively. The vertical axis in Fig. 26 (a) represents the ratio of the principal stress \u03c3 1 or \u03c3 2 to the shear stress \u03c4 and the vertical axis in Fig. 26 (b) denotes the ratio of the wrinkling strain \u025b w to the shear strain \u03b3 ( \u03b3 = \u03c4 G , G is the shear modulus). The solid\nr dashed lines in Fig. 26 (a) represent the major principal stresses nder different prestresses while the discrete symbols denote the\ninor stresses.\nFig. 26 (a) shows that when the prestress increases from 0.1 \u03c4 0.5 \u03c4 , the major and minor principal stresses are augmented lightly in the taut area. Fig. 26 (b) indicates that the prestress s able to decrease the wrinkling strain notably in the wrinkled egion." ] }, { "image_filename": "designv11_5_0000020_978-3-030-12391-8_5-Figure5.1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000020_978-3-030-12391-8_5-Figure5.1-1.png", "caption": "Fig. 5.1 (left) Photo of a test on the S4 Beam. (right) S4 Beam finite element model", "texts": [ "7) The energy dissipated for each vibration cycle is the area under the hysteresis curve, Dr (\u03b1) = \u222b qr (\u03b1) \u2212qr (\u03b1) ( f\u03021 (qr ) \u2212 f\u03022 (qr ) ) dqr = 2 \u222b qr (\u03b1) \u2212qr (\u03b1) ( fr ( qr+qr (\u03b1) 2 ) + fr ( qr (\u03b1)\u2212qr 2 ) \u2212 \u03b1 ) dqr (5.8) The integral above is readily evaluated using the trapezoid rule and then the effective damping ratio \u03b6 r(\u03b1) may be determined by analogy with a linear system [9, 12]. \u03b6r (\u03b1) = D (\u03b1) 2\u03c0(qr (\u03b1) \u03c9r (\u03b1))2 (5.9) The structure of interest was named the S4 Beam in [19] and is shown in Fig. 5.1. The structure consists of two C shaped beams, bolted together at each end. Impact testing was used to determine the amplitude dependent natural frequency and damping for the first several modes of the structure. These experimental measurements will be compared with the amplitude dependent natural frequency and damping predicted from various finite element models. A detailed explanation of experimental setup for the S4 Beam can be found in [19]. The finite element model used in this work borrowed from the 3D model created in [20], using the parameters, mesh density, etc", " The joint interfaces on the beams were machined to be nominally flat, but measurements revealed that there was some curvature, and so an effort was made to quantify the curvature of the contact patches and apply this to the FEM. In later sections of this paper, the measured surface contours of the beams were applied to the nodal positions on the joint interface to model a more realistic contact. In order to do this, surface contour plots were constructed which showed that the nominally flat surfaces had variation of up to 150 \u03bcm in height. Two measurements were used to extrapolate a surface contour. Referring to the coordinate system in Fig. 5.1, one measurement was parallel with the X axis and one with the Y, both measurements intersecting at the center of the bolt hole. Each measurement was fit with a sixth degree polynomial. All fits had R2 values of greater than 0.97. The curvatures in x- and y- were then added together to approximate the surface of the real machined components. For brevity, the measurements and the fits to them were not included here, but plots of the surface profiles are shown in Fig. 5.2. The titles above each surface indicate which interface they belong to, referring to the coordinate system in Fig. 5.1. For example, the +Y, +X surface is on the top beam on the right joint in Fig. 5.1, and \u2212Y, \u2212X would be on the bottom beam on the left. It should be clearly stated that the surface contours for this model are gross approximations meant only to roughly represent the variations in the surfaces. This method is likely a poor approximation of the edges of the interfaces, but as is shown later, the edges of the interface likely play no role in the contact of the joint. In seeking to obtain agreement between the simulations and the experimental measurements, three main variables were studied" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000958_13621718.2019.1666222-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000958_13621718.2019.1666222-Figure3-1.png", "caption": "Figure 3. Details of the evaluation parameters.", "texts": [ " The hot wire was in the middle of the two electrode, and therefore the distance to each electrode was equal to 12.5mm. Several welding experiments are performed with the range of the process parameters listed in Table 2 based on the actual manufacturing field conditions. During the experiments, the trailing electrode condition was fixed through the preliminary experiments. In the case of the fillet welding, the desired penetration geometry is a rounded penetration rather than a finger-like shape, as shown in Figure 3. The finger-like penetration is formed due to the high current, which generates high pressure in the centre of the welding pool. To predict the penetration profile from the process parameters is difficult without destructive tests of welded specimen. The range of the major parameters in the hybrid tandem MAG is roughly selected through the initial experiments. The selected welding parameters are the current and the voltage of the leading electrode, and the current of the hot-wire, all of which are known for being deciding factors for the shape of the fusion zone. Each parameter has five levels. After the welding experiments, the specimen is cut. The cut samples are mounted with epoxy resin and polished for the measurement of the bead geometry. The prepared macro specimens are processed by 5%nital etchant. After etching, the bead geometry of the specimen is measured by an opticalmicroscope, then thewelding performance of each experiment is evaluated by the criteria indicated in Figure 3. Generally, in the tandem welding, the electromagnetic force between two DC-EP electrodes produces arc Table 2. Parameter values for the hybrid welding experiments. Current Voltage CTWD Travel angle Work angle Travel speed Gas flow rate [A] [V] [mm] [deg] [deg] [mmin\u22121] [Lmin\u22121] Leading electrodes 300\u2013380 30\u201338 18 10 42 1.5 18 Trailing electrodes 300 30 21 10 42 1.5 18 Hot-wire 0\u2013160 \u2013 25 0 42 1.5 18 interference and weld pool disturbances because of the attractive forces induced between the electrodes", " In fillet welding, the desired shape is a rounded shape, as mentioned above. The experiment with optimal values shows a good rounded shape and equal leg lengths. This confirms the GPR model validity. In this paper, a hybrid tandem metal active gas welding process, which consists of tandem electrodes and the hot-wire, is modeled based on Gaussian process regression (GPR). The fillet welding shape is defined using seven performance characteristics based on the International Organization for Standardization (ISO) shown in Figure 3. The performance characteristics are predicted with the GPR model to evaluate the welding quality. When this GPR is modelled, the ISO standard and thumb rules of the field are considered on the weighting factor and the cost function definition. The comparison shows that GPR can accurately model the welding process in the presence of big uncertainties and noise. These uncertainties came from the metal transfer when the electrodes are engagedwith a high current. The proposed approach is used to optimise the process parameters, such as the voltage and current of the leading electrodes, and to improve the system performance in other manufacturing processes that are based on welding technology" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002202_j.jmapro.2020.10.004-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002202_j.jmapro.2020.10.004-Figure2-1.png", "caption": "Fig. 2. System elements of laser-PTA processing platform.", "texts": [ " Powder feeding system continuously provides powder flow, which is fed coaxially by plasma arc torch. The powder is 304 austenitic stainless steel powder and the granularity is 125~300 mesh. The base metal is 10 mm thick 304 stainless steel substrate. Heat source melts the feeding powder and substrate to form molten pool and deposition layer on the surface of the substrate. The metal substrate is clamped on the operating platform. A programmable motion control mechanism helps to achieve precise control of the operating platform in three dimensions. Fig. 2 shows the system elements of Laser-PTA processing platform. This paper proposes to use a pulsed laser beam to excite high-frequency oscillation in the molten pool, and to regulate the internal structure of deposition layer through the oscillation effect. The plasma arc torch and the laser lens have a certain inclination angle in Fig. 2. A suitable working position helps the pulsed laser to perform the thermal dynamics intervention on the liquid molten pool. Table 1 lists the main arc, laser and powder feeding parameters used in the experiment. The arc current is kept at 50A. Under this arc current value, a circular arc spot about 10 mm in diameter will be produced on the substrate. The defocusing amount of laser beam is \u2212 1 mm, which produces a circular spot with a diameter of 0.8 mm on the substrate. In order to ensure the comparability of the experimental results, other parameters are kept constant except for the peak power of laser" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001312_s00170-020-05434-3-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001312_s00170-020-05434-3-Figure2-1.png", "caption": "Fig. 2 Method of depositing variable width structures using multi-track overlapping. a Parallel track fill mode. b Variable overlapping rate", "texts": [ " A transfer function for prediction of the bead width was established based on a large number of single-pass deposition tests. A thin wall ranged from 1.60 to 0.83 mm was deposited. However, these open-loop control methods proposed in do not guarantee the stability and repeatability of the deposition, and their ranges of width variation were small. Therefore, the LDMD technology based on coaxial powder feeding mainly uses multi-track overlapping to realize the change of width. One method is parallel track fill mode, and the other is variable overlapping rate as shown in Fig. 2 [18]. J. Michael Wilson et al. [19] repaired turbine blade defects by LDMD using parallel line laser scanning mode. But the parallel track fill mode will result in a \u201cstaircase effect\u201d with low forming accuracy as shown in Fig. 2a. And for a variable width thin wall structure with a small width, it cannot be formed by parallel line laser scanning mode. However, if only changing the overlapping rate to change the width will result in deteriorating cladding layers due to different overlapping rates, it will affect the forming result as shown in Fig. 2b, so there is still no better way to form variable width thin-walled parts using LDMD technology based on coaxial powder feeding. In addition, in the field of metal wire AM technology, there are also a lot of research on the forming method of thin-walled parts with variable width. Ding et al. [20, 21] deposited two variable-width parts with multi-bead overlapping. However, this method does not guarantee the geometric accuracy of the deposition. Furthermore, the surface of the shaped part is not uniform" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002569_j.polymertesting.2021.107183-Figure13-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002569_j.polymertesting.2021.107183-Figure13-1.png", "caption": "Fig. 13. Finite element model of specimens and deformed profiles: (a) UT, (b) PT, (c) ET, (d) CT.", "texts": [ " Abaqus is driven by script commands, and the finite element simulation results are imported into the platform as the basis for comparison with test data. The validity of the proposed constitutive model and prediction method are verified by experiment results. Finite element models are generated to simulate the experiments described in Section 3. C3D8HT element with 8-node linear brick, hybrid algorithm is utilized for three-dimensional model and CAX4HT element with 4-node is applied for two-dimensional model. Finite element models of rubber specimens for UT, PT, ET and UC deformation modes and their deformed shapes are shown in Fig. 13. For the hyper-elastic analysis, the load-deflection curves of two rubber materials obtained from both the simulation and the experiment are respectively compared in Figs. 14 and 15. It is shown that the simulated response agrees well with those observed in the experiment, which implies the suitability of the material parameters. Fig. 16 presents the numerical and experimental comparison of stress relaxation with the stress against time of the studied Type A and Type B rubber at different strain levels" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000564_tia.2016.2616397-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000564_tia.2016.2616397-Figure3-1.png", "caption": "Fig. 3. (a) Stator of the simulated machine using FEM and its meshing and (b) Parameters expressing stator slot geometry.", "texts": [ " The stator may be wound with parallel branches (known as split-phase winding), which is common for hydro- generators [18]. Where such a winding is employed in the generator and their connections cannot easily be un-welded, some of the proposed approaches, particularly the threeammeter method, cannot be utilized to exactly determine the faulty conductor. This is not scrutinized in this paper and can be a topic of future work. To evaluate the proposed methods, a power station generator with specifications given in Table I is simulated using FEM. Fig. 3a shows meshing arrangement of the studied generator using the 2-D FEM-based simulation software including the stator core and its three-phase conductors. Fig. 3b presents the geometry of a stator slot with the corresponding parameters. To study the three-ammeter and the two-ammeter approaches, the rotor is extracted and one of the stator phase winding is energized, as illustrated in Figs. 1 and 2. The energized phase winding consisting of an internal PG fault is studied employing the Circuit Editor module of the FEM-software as shown in Fig. 4 where La1 to La10 are respectively the inductance of conductors 1 to 10 belonging to the stator phase under study" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000451_j.robot.2016.05.012-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000451_j.robot.2016.05.012-Figure3-1.png", "caption": "Fig. 3. A spatial multi-legged walking robot on irregular terrain.", "texts": [ " Accordingly, if a vertical force, force perpendicular to the ground, is applied to the robot, the stability will be more sensitive to top-heaviness as compared to other parameters. This implies that when modifying the FFSM, the scaling factor should be taken into consideration along with the geometrical configuration and mass of the robot. The following sections provide a general MFFSM for spatial robots, discuss planar robots as a special case, and consider the application of MFFSM to wheeled systems. Consider a spatial multi-legged walking robot traversing an irregular terrain as shown in Fig. 3, where the support polygon is not restricted to be planar. In Fig. 3, Mt is the net moment vector acting at the CG and Ft is the net force vector acting at the CG where Ft = \u2212 n i=1 Fi and Fi is ith foot force vector whose normal component is fi with amagnitude of fi = \u2225fi\u2225. Hence, f\u0304 = 1 n n i=1 fi directly correlates to the net force acting on the system and can be used for top-heaviness sensitivity with the FFSM [6]. On the other hand, the stability enhancement due to height reduction and foot placement spread could be directly taken into consideration by scaling the FFSM by Pi hi , where Pi = \u2225Pi\u2225, Pi is the tip-over axis normal vector created by perpendicularly connecting the CG to the tip-over axis, and hi is the height of the CG with respect to the tip-over axis aligned with the gravity vector which is the vertical portion of the tip-over axis normal vector, Pi" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002011_j.jmapro.2020.05.026-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002011_j.jmapro.2020.05.026-Figure3-1.png", "caption": "Fig. 3. The uneven top surface in a variable parameter deposition process.", "texts": [ " constructed a feedback control system to control the deposition height, and two kinds of sine-shaped variable-height parts were deposited using a designed PID controller [28] and a PID with sliding mode controller [29] with excellent control performances. The aforementioned variable height deposition strategies using open-loop and closed-loop control methods remain incapable to well achieve a geometrical stable and accurate deposition process, and the sample parts were deposited with only a small number of layers. Compared with the uniform deposition, the continuously varying process parameters in different deposition heights is more likely to produce an uneven top surface, as illustrated in Fig. 3. A small unevenness can increase after the next several layers of deposition. Therefore, the control task is not only to realize a variable height track in every single layer, but also to measure and compensate the produced convex and concave positions. The PID controller for the deposition height presented in reference [28] demonstrated remarkable dynamic performances, and the controller had compensated two slots in the substrate. However, the molten pool moved about 2mm during the settling time, and this 2mm area in the forming part was not smoothed and would lead to a new unevenness. Zeinali and Khajepour [30,31] built an adaptive fuzzy inverse dynamic model for the LMD process and devised a sliding mode controller, which is capable to compensate unknown and uncertain disturbances like slots with excellent control performance. However, the settling time and overshoot remained. Besides the unevenness caused by the settling time and overshoot during the closed-loop control process, the consumed time in the feedback process is worthy of consideration. As shown in Fig. 3, a convex point is measured at the time point t1. After the time delay of the feedback process in the hardware, the cladding nozzle has moved with a scanning speed v to another height position at the time point t2, and then the execution of the actuators ceases to have effect on the previous height point. This controlled value, as well as the measured value shifts with the movement of the molten pool. Furthermore, the high frequency change in the control input value requires a high real-time performance of the control system and is possible to cause a geometrical unstable deposition process and non-uniform heat absorption", " The substrate material used was 304 stainless steel. The powder material was the Fe-based alloy Fe313 (composition in wt,\u2212%: C\u22480.1; Si= 2.5\u20133.5; Cr= 13\u201317; B=0.5\u20131.5; Fe= bal.) with the particle size of 75\u2212150 \u03bcm. Both shielding gas and powder carrier gas were nitrogen. The \u201cself-healing\u201d effect in the LMD process was validated in multiple studies [36,37]. In a certain range of the defocus distance d (Fig. 4a), the deposition height of the actual layer varies with the working distance w (Fig. 4a) inversely. That means in Fig. 3, the deposition height of a new layer is lower in convex points and higher in concave points with a varying working distance. Therefore, the uneven surface can be automatically smoothed after several layers under an open-loop control. The \u201cself-healing\u201d effect of an IBPF nozzle with a range of \u22123mm \u223c \u22125mm of the defocus distance d was verified in [37]. The model of the IBPF nozzle used in this study has a larger \u201cselfhealing\u201d effect range of \u22123mm \u223c \u22129mm. In a layer or a segment of an equal-height deposition, the \u201cselfhealing\u201d effect can be taken advantage of when the process parameters are invariant" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001800_978-3-030-40410-9-Figure4.14-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001800_978-3-030-40410-9-Figure4.14-1.png", "caption": "Fig. 4.14 Cylindrical robot with spherical wrist", "texts": [ "4 Denavit\u2013 Hartenberg (DH) Convention 87 cord_3 =H_1*H_2*H_3; cord_4 =H_1*H_2*H_3*H_4; cord_5 =H_1*H_2*H_3*H_4*H_5; cord_6 =H_1*H_2*H_3*H_4*H_5*H_6; cord_7 =H_1*H_2*H_3*H_4*H_5*H_6*H_7; drw_line(cord_0,cord_1); hold on grid on drw_line(cord_1,cord_2); drw_line(cord_2,cord_3); drw_line(cord_3,cord_4); drw_line(cord_4,cord_5); drw_line(cord_5,cord_6); drw_line(cord_6,cord_7); axis([-2000 2500 -1000 1000 0 1500]) The code of the function HTM is the same as in Example 4.3, while the code for the function drw_line is: function [] = drw_line(H_1,H_2) plot3([H_1(1,4),H_2(1,4)],[H_1(2,4),H_2(2,4)],[H_1(3,4) ,H_2(3,4)],'k.-','MarkerSize',15); end Example 4.5 (Cylindrical robot with spherical wrist) This manipulator has 6DOF with two prismatic joints and three revolute joints. Figure 4.14 shows the robot, where the three last joints bounded by a dotted ellipse are called the spherical wrist. The coordinate frames have been assigned using the convention as usual, where zaxis for each link is the direction of actuation. To simplify the calculations, we have chosen the origins of the last three joints of the wrist to coincide at the same point at the centre of the wrist. The simplification comes from dropping any additional offsets between coordinate frames. Equation (4.4) should be used for the revolute joint while Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002267_tpel.2020.3038741-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002267_tpel.2020.3038741-Figure1-1.png", "caption": "Fig. 1. Winding configuration and flux paths of three-phase 12/8 MCSRM.", "texts": [ " Section III analyzes the effect of the current excitation angle and saturation level on power factor. Section IV presents the proposed optimized control method for MCSRMs. Section V validates the proposed method with experiments. Finally, Section VI concludes this article. In order to conduct torque ripple and power factor analysis, a three-phase 12/8 SRM has been used in this article. The motor used in this investigation is designed to operate as a CSRM. The winding configuration has been changed to operate it as an MCSRM. Fig. 1 shows the winding configuration and the flux directions for the three-phase 12/8 MCSRM. The motor specifications are shown in Table I. Finite-element analysis (FEA) has been conducted to analyze the torque ripple behavior of this motor with sinusoidal current excitation. The harmonic content of the generated torque waveform is shown in Fig. 2 when the d-axis current (id) and q-axis current (iq) are equal to 4 and 20 A, respectively. It can be noticed from Fig. 2(b) that the generated torque waveform has a dc component in addition to the 6th-, 12th-, 18th-, and 24th-order harmonics" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000590_j.triboint.2016.10.045-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000590_j.triboint.2016.10.045-Figure3-1.png", "caption": "Fig. 3. The loading condition of H half-spaces for overlapping with twice mirroring.", "texts": [ " As MA, NA, MB and NB are only related to the mesh structure and Poisson's ratio, these four matrices can be calculated in advance and stored, and then employed in different loading conditions or friction coefficients by Eqs. (9)\u2013(11). As an advantage of matrix formulation, it provides a convenient way to avoid the interactive calculation each time [7]. For thin finite-length space, the shear stresses on free end surface much increases and they can\u2019t be neglected during the overlapping process in Fig. 2, and in this case an improved model can be adopted by extending the loads on H half-space, as shown in Fig. 3. Base on the mirrored loads in Fig. 2, second mirroring for Ph2, T2, Ph3 and T3 are applied, so that the two shear stresses of surface I decrease because they are introduced by Ph5 and T5, which are farther away from surface I than Ph3 and T3. Although the loading condition in twice mirroring model is more complicated, the matrix formulation can be modified easily by adding terms of the loads on surface 4 and 5 to Eqs. (2) and (3). The loads can be further extended for the third time or more if necessary" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002726_s12206-021-0639-4-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002726_s12206-021-0639-4-Figure1-1.png", "caption": "Fig. 1. Structure of the PAM elbow exoskeleton.", "texts": [ " Meanwhile, the terminal sliding mode control combined with a twisting algorithm is applied in the proposed method to reduce the tracking error of the PAM exoskeleton under unknown nonlinear functions and external disturbances. The study is organized as follows. In Sec. 2, the mechanical structure of the exoskeleton system is presented. In Sec. 3, the proposed control method is illustrated. Sec. 4 gives several experiments with results. Lastly, the conclusion is drawn in Sec. 5. The PAM elbow exoskeleton refers to a type of robot assistance device designed to help weak people enhance their arm force. The exoskeleton is designed as a humanoid arm struc- ture (see Fig. 1). The exoskeleton has an overall weight of 1.02 kilograms without a gas source. The rigid link structure was adopted by complying with several design schemes to ensure the loading capacity of the exoskeleton [44, 45]. The entire mechanical system primarily consists of a handle, two brackets, two forearm links, and two upper arm links. The links present some positioning holes to regulate the length of the links in accordance with the human arm size. At the end of the upper arm links, a soft shoulder protector support is utilized to anchor the exoskeleton to the human shoulder" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001644_j.conbuildmat.2020.121961-Figure15-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001644_j.conbuildmat.2020.121961-Figure15-1.png", "caption": "Fig. 15. Elongation of a strand wire.", "texts": [ " Therefore, as the lay angle increases, the axial strain of the locked coil wire rope increases. In the process of wire twisting, if the cable lay length is large, steel wire skipping will occur. Therefore, the cable lay length should be increased within a certain range. Benndorf [16] was the first to work out a partition of the wire rope tensile force in wire tensile force. The calculation of tensile stress of the spiral strand in the current study is based on the research of Feyrer [17]. The geometric relationship in the process of steel wire stretching is shown in Fig. 14. and Fig. 15.. A partition of the tensile force in wire tensile forces must be calculated [17]. The wire tensile force component in the strand axe direction (neglecting the small shear force) is Si \u00bc Ficosai Ui \u00bc Fisinai : \u00f02\u00de The tensile force of the wire layer i is Fi \u00bc Dli li EiAi; \u00f03\u00de where li is the wire length, Dli is the wire elongation, Ei is the elasticity modulus, and Ai is the cross-section of a wire in wire layer i. The extension of the wire is ei \u00bc Dli li : \u00f04\u00de With lS respect to the length of the strand, the length of the wire is li \u00bc lS cosai ; \u00f05\u00de Therefore, when the failures of higher classification are neglected, the wire elongation is Dli \u00bc \u00f0DlS Dui tanai\u00decosai: \u00f06\u00de The contraction of the winding radius, that is, the circumference in relation to the wire extension that transverses the contraction ratio, can also be designated as \u2018Poisson\u2019s ratio\u2019 of the wire helix as follows: t \u00bc Dui=ui Dli=li ; \u00f07\u00de Dui \u00bc t ui Dli=li; \u00f08\u00de Dui \u00bc t ui Dli=li; \u00f09\u00de Dui \u00bc t li sinai Dli=li: \u00f010\u00de On this basis, the elongation of a wire in the wire layer i is: Dli \u00bc DlS cosa 1\u00fe tsin2a : \u00f011\u00de This equation, together with Eqs" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000798_j.neucom.2019.04.056-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000798_j.neucom.2019.04.056-Figure5-1.png", "caption": "Fig. 5. Transverse swing p ( z, t ) of the suspension cable with proposed control scheme.", "texts": [ " l The initial conditions are set as s (z, 0) = 0 , s \u2032 (0 , t) = 0 , \u02c6 D\u0304 (0) = 10 , and \u02c6 m (0) = 1 . 0 . All the design parameters are \u03b21 = 1 , \u03b22 = 1 , \u03bb1 = 1 . 0 \u00d7 10 \u22123 , \u03bb2 = 0 . 5 \u00d7 10 \u22123 , \u03b3 = 0 . 01 , k 1 = 1 . 5 \u00d7 10 6 , and k 2 = 1 . 0 \u00d7 10 3 . Figs. 4 \u20138 are the responses of the helicopter suspension cable system which is described by (10) under boundary conditions (16) and (17) with cable swing constraint and unknown dead-zone. The transverse oscillation amplitude without control and with the proposed control scheme are illustrated by Fig. 4 and Fig. 5 , respectively. The slop of the suspension cable with the proposed control scheme is illustrated by Fig. 6 . It shows the closed-loop system can achieve a good performance by using the proposed unilateral boundary adaptive control scheme even considering cable swing constraint and unknown dead-zone. Fig. 7 illustrates that cable swing constraint can not be violated under the designed unilat- 0 50 100 150 \u22120.03 \u22120.02 \u22120.01 0 0.01 Time (s) s (L ,t ) Fig. 7. The boundary slope s \u2032 ( L, t ) with proposed control scheme" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000685_1077546318818694-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000685_1077546318818694-Figure1-1.png", "caption": "Figure 1. Faulty generator: (a) faulty bearing; and (b) rubbed rotor and stator.", "texts": [ " Some failed subassemblies, for example, gearbox and generator, have both expensive production costs and transportation fees while being replaced. In this case, long shut down and a large amount of economic loss will not be avoided. In wind turbine generators, bearings are the critical parts which support the rotors rotating at a high revolution speed. Regardless of the health condition of these parts, the incipient faults on bearings will deteriorate and lead to results such as the eccentricity of air gap and even the rub between the rotor and stator of the generator, as shown in Figure 1. In a wind farm, tens or even hundreds of wind turbines operate together and several vibration transducers are mounted on the drive train of each wind turbine, producing massive monitoring vibration signals contaminated by intensive background noise. The processing and analyzing of these signals by operators at wind farms to detect faults on bearings is time-consuming and of low accuracy. In the light of this, the need to develop an adaptive fault detection method is urgent for improving the intelligent fault diagnosis accuracy of wind turbines" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000649_icinfa.2016.7832081-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000649_icinfa.2016.7832081-Figure1-1.png", "caption": "Fig. 1. Quadrotor aircraft with hanged payload", "texts": [ " In this paper, we applied the ADRC strategy to the payload transporting issue. This paper is organized as follows. Section II presents the model of a hanged payload and the control problems. In section III we describe the control strategy using ADRC. Section IV provides simulation results. Finally, we draw conclusions in Section VI. In this paper, we choose to study the system combining the quadrotor with a hanged payload, which is connected with the quadrotor by using a rigid pole, as shown in Fig. 1. To simplify the model, we assume that the quadrotor move within the longitudinal plane. First we select the coordinate frame, including inertial coordinate frame {I}, body coordinate frame {B}. Where {ex,ey} is the ground fixed frame of {I}, {e1,e2} denote the body fixed frame of {B} as shown in Fig. 1. The configuration variables are q = [ x z \u03b8 \u03b1 ]T , q\u2208 R4, where x 978-1-5090-4102-2/16/$31.00 \u00a92016 IEEE 1641 is the quadrotor position along ex, z is the quadrotor position along ez, \u03b8 is the pitch angle, \u03b1 is the payload angle with respect to ez, while M is the mass of quadrotor, m is mass of the load, m0 is the mass of the pole between the payload and the center of quadrotor and l is the length of the pole, d is the distance from motors to the center of quadrotor, f1 and f2 represent the thrust force provided by the motors" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001037_tec.2019.2951659-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001037_tec.2019.2951659-Figure4-1.png", "caption": "Fig. 4: Analytical calculations of RAS in AFM", "texts": [ " See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. To model the AFM under RAS, the magnetic flux (\u03b2r) in (1) needs to be modified around the circumstance of each computational plane. In which, the area where the rotor and the stator are not aligned, \u03b2r is assumed to be zero (the leakage flux is neglected), otherwise, there is no change in \u03b2r. In SRAS, this area is fixed in space, however, in DRAS, this area will be rotating with the rotor position. Based on Fig. 4, the representation for \u03b2r is derived in (22) - (26). The change in \u03b2r depends on the amount of the shift between the rotor and the stator, and on the average radius of each computational plane. Based on amount of the shift between the rotor and stator geometry (i.e.\u03b5XY ), the motor geometry can be divided into five regions (shown in Fig. 4). For each computational plane, Ravg,i is calculated and used to define which region this computational plane intersects with. The section of the computational plane in region 1 or 3 will have zero magnetic flux density, while the section in region 2 will assume to have a nonzero magnetic flux density (Regions 4 and 5 will not be considered because the computational planes will never intersect with them). Therefore, the magnetic flux density around the circumstance for each computational plane are: \u2022 For Rout-Ravg,i < \u03b5XY and Ravg,i-Rin > \u03b5XY , the computational plane intersects with region 1 and region 2" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001443_rpj-11-2019-0287-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001443_rpj-11-2019-0287-Figure2-1.png", "caption": "Figure 2 Dimensions and building angle of the 60\u00b0sample", "texts": [ " Two groups of samples, one group with support structures and the other group without support structures, were designed with the same specifications and printed at the same time by the SLM process. The dimensions of all samples were 20 20 5 mm (length width height). To get a more realistic result and further understand the effect of support structures, the two groups of samples were printed at 45, 60 and 75\u00b0 building angles. The building angle refers to the angle with respect to the horizontal line, as illustrated in Figure 2. During the sample printing process, all printing parameters remained the same, including laser power, scanning speed, hatch distance, layer thickness, scanning strategy, etc. Therefore, the only difference was the support structure. For each design with different building angles, three samples were printed with and without support structures, which were mainly used to take the average value in the subsequent measurements. Therefore, a total number of 18 samples were printed. All the support structures adopted the same default design of block support structures generated by the Materialise Magics software" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000471_j.ijsolstr.2016.07.004-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000471_j.ijsolstr.2016.07.004-Figure5-1.png", "caption": "Fig. 5. Three states of a membrane ( Jarasjarungkiat et al., 2009 ).", "texts": [ " 2008 and 2009 ) nd U 2 ] T [ C][ U 1 ] [ U 1 ] T [ C][ U 2 ] [ U 3 ] T [ C][ E] \u2212[ U 1 ] T [ C][ U 1 ] [ U 2 ] T [ C][ U 2 ] [ U 3 ] T [ C][ E] + [ U 3 ] T [ C][ U 1 ] [ U 2 ] T [ C][ U 2 ] [ U 1 ] T [ C][ E] \u2212[ U 3 ] T [ C][ U 1 ] [ U 1 ] T [ C][ U 2 ] [ U 2 ] T [ C][ E] + [ U 3 ] T [ C][ U 2 ] [ U 1 ] T [ C][ U 1 ] [ U 2 ] T [ C][ E] \u2212[ U 3 ] T [ C][ U 2 ] [ U 2 ] T [ C][ U 1 ] [ U 1 ] T [ C][ E] = 0 (32) or a slack membrane ( Wang X.F. et al., 2014 ). Eqs. (31) and ( 32 ) re nonlinear equations on the wrinkling angle and they are sually solved by an iteration algorithm such as Newton method Jarasjarungkiat et al. 2008 and 2009 ). .6. Judgment of the wrinkling state of a membrane A membrane has three states in general, i.e. taut, wrinkling and lack (see Fig. 5 ), and it would be in any one of these states when ubjected to a kind of load or deformation. A membrane behaves ifferently in a different state. When a membrane is taut, it is sotropic and elastic if plastic deformation does not exist, and the lastic constitutive matrix [ C ] is used in the numerical analysis. hen a membrane is wrinkled, its stiffness in the direction perendicular to the wrinkles vanishes (flexural stiffness is assumed o be zero) and the modified constitutive matrix, i.e. Eq. (23) is dopted" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000888_j.promfg.2019.06.223-Figure22-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000888_j.promfg.2019.06.223-Figure22-1.png", "caption": "Fig. 22. An AM system with two gantries for contour crafting of a cottage.", "texts": [], "surrounding_texts": [ "Contour crafting [31] is a large-scale AM process with multi-gantry configuration to construct a building layer by layer. Fig. 20 shows the construction of a single-material cottage by two gantry-style actuators, modelled with the position-based constraint. The rooftop has been set transparent for visualization of internal structures. Fig. 23(a) presents a cottage layer sliced with an X-Y plane. The cross section is a single CF because the walls are inter-connected, but it can be divided into edge segments at the wall intersections to speed up fabrication [31], as in Fig. 23(b). The safety envelopes are based on the sizes of the end-effectors of the gantries, as in Fig. 23(c). The total build time is about 39.64 hours with the concurrent toolpaths, and 59.61 hours with sequential toolpaths." ] }, { "image_filename": "designv11_5_0000200_s00202-019-00882-x-Figure15-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000200_s00202-019-00882-x-Figure15-1.png", "caption": "Fig. 15 Eddy current volumetric loss distribution, nin = 2200\u00a0rpm/nout = 400\u00a0rpm", "texts": [], "surrounding_texts": [ "Distributions of the eddy current losses are presented in Figs.\u00a015, 16, 17, 18 and 19 for the NdFeB magnets [27, 28]. These losses are limited to electrically conductive parts only. Results correspond very well with these for electrical conductivities of permanent magnets, shown in Table\u00a02, and total losses and efficiency in Table\u00a03. Eddy currents are estimated at five rotational speeds, for inner rotor 150, 550, 1100, 1650 and 2200\u00a0rpm and for outer rotor rotational speeds 27, 100, 200, 300 and 400\u00a0rpm, respectively. Output rotor torque reduction due to eddy current demagnetization effect for different rotational speeds is shown in Fig.\u00a020. Eddy currents, neglecting the temperature changes of resistivity, are linearly related to rotational speed [29, 30]. Eddy current losses rise with the square of rotational speed. Also, they are influenced by the huge number of magnetic pole pairs. At rotational speeds greater than 500\u00a0rpm for the magnetic gear construction under consideration, the heat effect of eddy current losses could not be neglected. At speed 10,000\u00a0rpm, only the eddy current losses are overcoming 4% of the transmitted power. Eddy current volumetric loss distributions show that most of the losses are located in the low-speed rotor magnets. They are induced by high-speed rotor magnet movement and are amplified by modulating segments at high harmonics of the magnetic flux. Frequency separation of eddy current losses is important for loss analysis. 1 3 8 Magnetic field harmonic distortion of\u00a0CMG In many existing researches, eddy current losses are often ignored in steady state because of direct analogy with single-rotor electrical machines with permanent magnets. In MGs, these losses still appear because of relative movement of two rotors and modulating segments. Harmonic flux distortion increases eddy current effects. The main rotational frequencies of the magnetic field of the magnetic gear construction for the outer and inner rotors are 55\u00a0Hz and 10\u00a0Hz. The radial components of the flux density in the air gap between the inner rotor and the steel segments of the magnetic gear are shown in Fig.\u00a021. The radial components of the flux density in the air gap between the steel segments and the outer rotor of the magnetic gear are shown in Fig.\u00a022. The fast Fourier transform (FFT) analysis of the radial components of the flux density in the air gap between the inner rotor and the steel segments is shown in Fig.\u00a023. The FFT analysis of the radial components of the flux density in the air gap between the inner rotor and the steel segments is shown in Fig.\u00a024. 1 3 FFT analysis shows significant harmonic distortion. All frequency harmonics different from main 55/10\u00a0Hz are creating torques with no proper frequency of rotation, limiting the output CMG torque. According to these estimations, more than 20% of the total magnetic flux is engaged in these undesired frequencies [31]. This distortion effect does not influence the power efficiency, but it limits the CMG output torque and does not use the available flux from the permanent magnets. The harmonic distortion is not related to direct power losses because we are not using electrical supply for these fluxes, but it limits the output torque value of the CMG design. Some words on hysteresis loss are necessary, considering frequency-dependent hysteresis models, such as Steinmetz equation variations; they are suitable for loss superposition over magnetic flux amplitudes and frequency harmonics, like shown in Figs.\u00a023 and 24. However, a closer look reveals that hysteresis loss is not substantial for MG operation, as a power loss, covered by electrical excitation as it is in rotational machines. In MG, magnetic hysteresis loop causes time-dependent flux non-linearity, decreasing this way slightly the dynamic magnetic torque interaction between rotors. The summarized results for torque reduction and losses according to rotational velocity are presented in Table\u00a04. Losses are estimated for 150\u00a0rpm, 2200\u00a0rpm and 10,000\u00a0rpm of high-speed rotor. Results are showing significant dynamic torque reduction in high rotational speeds and rise of eddy current losses. According to estimated losses, efficiency mapping at MG overload, at torques above 320\u00a0Nm, is shown in Fig.\u00a022. At low-speed overload, efficiency is influenced by rotor slipping, while in high speeds it is influenced by eddy currents (Fig.\u00a025)." ] }, { "image_filename": "designv11_5_0002245_1350650120969003-Figure11-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002245_1350650120969003-Figure11-1.png", "caption": "Figure 11. Rotor tracks of novel GFJB at different speeds. (a) 30,000 rpm; (b) 50,000 rpm.", "texts": [ " The test results quoted here is for the NASA test,27 and the bearing parameters are shown in Table 2. The rounding radius is set as Rr\u00bc 0.3mm according to the actual test on the arch shape. The calculated results exhibit a good correlation with the literature data, and the predicted minimum film thickness considering rounding fillet shows better agreement with the test result, especially when the static load is larger than 60N. This indicates that the correctness of theoretical model can be verified. Figure 11(a) and (b) show the rotor track of air compressor at the running speeds of 30,000 rpm and 50,000 rpm, respectively. The vertical and horizontal displacements of the rotor increase at a higher speed, but the trajectory still converges. This proves that the foil bearing has good supporting performance. By comparing the rotor trajectories supported by the novel and traditional GFJBs, it is found that the insertion of shim foil can make the trajectory slender, and the rotor vibration in both directions is also lower" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002341_iros45743.2020.9341246-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002341_iros45743.2020.9341246-Figure2-1.png", "caption": "Fig. 2. Pictorial description of the multi-mode TO for the task of halting a moving object.", "texts": [], "surrounding_texts": [ "I. INTRODUCTION\nSafe and robust robot manipulation under switching dynamics still poses many challenges. Typically, manipulation tasks require making and breaking contact with objects. This results in challenges in motion planning and control due to, among other factors, (i) the hybrid nature of the problem [1] and (ii) the uncertainties that arises due to contact dynamics [2].\nStrikingly humans are not only competent in object manipulation, but prefer to make contact with non-zero velocities to achieve the task faster and smoother [3]. The two key enablers to realise this are the ability to skillfully switch between free-motion and contact [4], and the capacity to shift between a variety of control mechanisms depending on the stage of the motion and their associated uncertainties [5].\nRecent hybrid Trajectory Optimization (TO) methods in robotics [6], [7], [8] have demonstrated efficient methods for multi-contact manipulation planning. Yet, it is not trivial to transfer these behaviours robustly on to the hardware due to the challenge of regulating the transitions between free motion and motion in contact, as well as dealing with imprecise timing of the transition in the reference motions. To address this, a number of hybrid control [9], [10] and compliance control [11], methods have been proposed. However, given the inherent limitations of the hardware [12], the impacts that a stand-alone controller is capable of dealing with are limited.\n\u2217Denotes equal contribution for both authors. E-mail: theodoros.stouraitis@ed.ac.uk, lei.yan@ed.ac.uk 1 Authors are with School of Informatics, University of Edinburgh, Edinburgh, U.K. 2 Authors are with Honda Research Institute Europe (HRI-EU), Germany. 3 Authors are visiting researchers at the Shenzhen Institute for Artificial Intelligence and Robotics for Society (AIRS), The Chinese University of Hong Kong (Shenzhen), China.\nIn this work, we try to address this problem at the level of \u2018impact-aware\u2019 manipulation planning. We ask ourselves, \u201dHow could we plan hybrid motions, such that they are easily executable by out-of-the-box controllers?\u201d, which can be re-framed as a problem of planning such that consistent contact can be maintained during and after impact\u2014even for tasks with contacts at speed, i.e. moving objects. As a typical example scenario, consider an agent that attempts to stop an object in motion, as shown in Figs. 1 and 2. In such a case, the agent needs to address the following challenges:\n\u2022 Plan discontinuous motions through contact events and\nphysical impacts, which may result in state triggered velocity jumps described by jump maps [13], i.e., jointly plan continuous motions (flows) and contacts (jumps) to perform a task. \u2022 Track discontinuous reference motions, where the actual\ntime of the jumps (impact) may not coincide with the jump time (impact time) of the reference motion.\nA number of motion planning methods have investigated impact related problems. In [14], [15], non-zero velocities at contact were avoided to exclude impact events. In [16], a QP controller ensured that unexpected impacts will not violate joint limits. They all resulted in either conservative motions or required a priori knowledge of the exact time of impact.\nCatching motions were demonstrated based on learned dynamical systems [17] and with TO methods [18]. However, the mass of the intercepted objects was negligible and contact was realized by caging the object. We consider objects with large size and mass, and contacts that can break anytime.\nIn this paper, we address the two challenges mentioned above with a coherent contact-invariant TO method that plans \u2018impact-aware\u2019 hybrid motions, while the control input results from a hybrid controller capable of absorbing impacts.\n978-1-7281-6212-6/20/$31.00 \u00a92020 IEEE 9425\n20 20\nIE EE\n/R SJ\nIn te\nrn at\nio na\nl C on\nfe re\nnc e\non In\nte lli\nge nt\nR ob\not s a\nnd S\nys te\nm s (\nIR O\nS) |\n97 8-\n1- 72\n81 -6\n21 2-\n6/ 20\n/$ 31\n.0 0\n\u00a9 20\n20 IE\nEE |\nD O\nI: 10\n.1 10\n9/ IR\nO S4\n57 43\n.2 02\nAuthorized licensed use limited to: San Francisco State Univ. Downloaded on June 18,2021 at 16:40:34 UTC from IEEE Xplore. Restrictions apply.", "The hybrid controller is based on compliance control that allows to mitigate the peak error due to the mismatch in time between reference and actual impact. Our TO method results in hybrid motions that are inline with the hybrid controller capabilities, while the controller\u2019s parameters (e.g. stiffness) are simultaneously optimized, as in [9].\nThe core insight is based on the duality between the impact model used and the capabilities of compliance controllers available in the latest collaborative robots. By modulating the robot\u2019s end-effector compliance, we can emulate a number of different types of collisions ranging from elastic to in-elastic, and deduce the optimal force transmission model given the system\u2019s limitations, e.g. workspace limits.\nThe contributions of our work are:\n\u2022 A parametric programming technique to encode both\nhybrid dynamics and hybrid control in a single multimode trajectory optimization formulation. \u2022 A generic impact model formulation based on a second-\norder critically damped system to generate smooth contact forces and simultaneously optimize the stiffness. \u2022 A multi-mode trajectory optimization framework that\ncan deal with both multi-contact motion planning and contact force generation for impact-aware manipulation.\nThis paper is organized as follows: Section II provides the background on systems with hybrid dynamics and hybrid control and reviews related work on hybrid motion planning. The details of the proposed contact force transmission model and TO method are given in Section IV. Section V presents the evaluation of the method and the experimental results. Finally, Section VI discusses current limitations and future directions.\nAs described in Section I, motion planning concepts for manipulation are often based on trajectories that guide an object to its desired state. In this work, we consider a class of systems where the trajectories include discontinuous transitions between different contact states. Similar to [6], [19], we describe systems with hybrid dynamics as\nx\u0307(t) = fk (x(t),u(t),v(t)) , if (x(t),u(t)) \u2208 Dk, (1)\nwhere x(t) \u2208 R n is the state of the system, u(t) \u2208 R m is the control actions of the plant, v(t) \u2208 R \u03bd is the control input applied on the environment, n,m, \u03bd \u2208 R define the dimensions of each quantity and k \u2208 {0, 1} indexes to the different sets Dk. Each Dk \u2282 R n\u00d7m defines the domain (relative to x(t) and u(t)) of a contact state, i.e. free-motion or in-contact. Note that (1) defines both the plant\u2019s and environment\u2019s dynamics.\nFor dynamic robot manipulation with contact changes, a number of TO methods [7], [20], [21] have been proposed. The underlying formulations have been borrowed from the locomotion domain [22], [23] and can be separated into two classes. The contact-implicit [22] and the multi-phase [23] or multi-mode [6] approaches. The former requires special\nattention to the relaxation of the problem to avoid spurious local minima [24]. The latter enables us to obtain a smooth NLP [24] given a mode sequence, which can be obtained from an outer-level process [6]. In [6], each mode is associated with the \u201dcontact activity\u201d (physical interaction between objects), is specified via path constraints and furthermore, the modes used can be contact, kinematic and stable. Here, we adopt the latter paradigm as the sequence of modes is fixed and it admits a general notion of modes [6] not limited to on / off contact.\nTO methods for contact planning transcribe the state and the control input of the system, which entails that forces are optimization variables too. Although [25] indicated that simply planning contact forces based only on the contact state does not suffice towards making stable contact and pointed out the importance of an accurate force model during contact transitions, most previous works neglected this aspect of the problem. Typically, a number of assumptions were made to transfer the motion plans to the robots. In [21], purely inelastic collision was assumed to impose no-rebound condition, while in [20], a variable smooth contact model was used that allows virtual forces from distance. Thus, a natural design question arises regarding the choice of the contact force transmission model. Such a model can be used to constrain the control inputs and could also be conditioned on the current mode of the system.\nHybrid control is useful when dealing with systems that require multiple separate controllers, e.g. Ott et al. [26] proposed a hybrid force controller for switching between impedance and admittance controllers. In this work, different control modes are used for the deformation and the restitution phases, shown in Fig. 3. Similar to [8], [27], we describe systems with hybrid control as\nv(t) = hl (u(t)) , if (u(t)) \u2208 Tl, (2)\nwhere l \u2208 Z indexes to a selected set Tl \u2282 R m. Each Tl corresponds to a controller type, e.g. impedance, admittance, direct force-control. (2) specifies the transformation from the plant\u2019s control actions to the environment\u2019s control inputs.\n9426\nAuthorized licensed use limited to: San Francisco State Univ. Downloaded on June 18,2021 at 16:40:34 UTC from IEEE Xplore. Restrictions apply.", "One can observe from (1) and (2), that the investigated system has a variety of different contact states and controllers that can alter the system\u2019s behaviour along the time axis. We refer to a single combination of a contact state and a controller as a mode of the system. The proposed notion for contact-control modes is similar to the notion of physical interaction modes introduced in [6] (see Section II-A). Here, we only consider a limited number of contact states as physical interaction modes, but we extend the notion of mode by considering a variety of different controllers.\nThe sequential arrangement of these modes zj = {(kj , lj)} defines the outline of the trajectory, while for each different sequence of contact-control modes z : {z0, z1, ...zJ} there is a different optimal solution of state \u2217x(t) and control \u2217u(t) trajectories. J \u2208 Z + describes the total number of modes of the trajectory.\nGiven a mode sequence, the multi-mode trajectories described by (1) and (2) can be explicitly expressed as a function of the initial state and the plant\u2019s action sequence. Inspired by [6], [8], [19], we think of impact-aware manipulation planning as a special form of Parametric Programming (PP) [28], where the sequence of modes z is encoded in the problem as\nmin x(t),u(t),v(t)\nc (x(t),u(t),v(t), z) (3a)\ns.t. x\u0307(t) = f (x(t),u(t),v(t), z) , (3b)\nv\u0307(t) = h\u0303 (u(t), z) , (3c) g(x(t),v(t),u(t), z) \u2264 0. (3d)\n(3a) - (3d) are piecewise functions from which the appropriate piece (interval) can be selected based on z. (3a) defines the objective function, and g(\u00b7) in (3d) represents both the equality and the inequality constraints of the system. It is worth pointing out that Optimal Control (OC) problems with hybrid dynamics are usually written as in (3), excluding (3c), while OC problems with hybrid control are usually written as in (3), excluding (3b). The formulation above defines an OC problem where both dynamics and control are hybrid. Further, we enforce (2) as a dynamical system through (3c). The details on this decision are given in the next section.\nNext, we consider one instantiation of such a problem\u2014 characterised as halting a moving object. For this task, the robot has to be initially soft to absorb the impact and then, stiff to accurately manipulate the object.\nIV. IMPACT-AWARE MOTION PLANNING\nA. Impact model\nIn scenarios where two objects collide (contact transition) with non-zero relative velocity, an impulsive force results. The velocity discontinuity between pre-impact and postimpact is described with the following relationship\nM ( v+ \u2212 v\u2212 ) = \u039b\u03b4t, (4)\nwhere M is the mass of the system, v\u2212 and v+ are the preimpact and post-impact relative velocities, respectively. \u039b is the impact force and \u03b4t \u2243 0 is the impact\u2019s duration.\nFor a moving object that experiences an impact, the dissipated energy\u2014due to the impulsive force\u2014during the collision is\nE\u039b = 1\n2 Mv\u2212\n2 \u2212 1\n2 Mv+ 2 . (5)\nIn this paper, we adopt the mass-spring-damper system to model real-world collisions [29]. The equation of motion for a mass-spring-damper system shown in Fig. 3 is written as\nMx\u0308+Bx\u0307+Kx = \u2212Mg, (6)\nwhere K,B,M are the stiffness, damping and mass respectively; g is gravity and x is the state of the system. The energy dissipation of such a system is caused by the damper and can be calculated as follows\nEp =\n\u222b \u03b4t\n0\nBx\u0307dx =\n\u222b \u03b4t\n0\nBx\u03072dt. (7)\nThus, by equating (5) with (7), we can model the energy loss during impact with a spring-damper system [29], where the dissipated energy during deformation and restitution stages (see Fig. 3) is related to both stiffness and damping.\nIn addition, based on [29], [30], the characteristics of the physical system such as duration of impact and restitution coefficient1 can be related to the mass, damping and stiffness parameters of the mechanical system. We utilize this observation to accurately emulate the physical interaction through the impedance controller of the manipulator. As shown in Fig. 3, the negative contact is defined as the deformation stage during which the contact force for making a stable contact is generated. The positive contact is defined as the restitution stage which generates the contact force for the manipulation tasks, such as pushing an object away. We encode these two stages of the contact as\nl =\n{\n\u22121, v\u2212 \u2192 0 1, 0 \u2192 v+.\n(8)\nIn terms of impact-aware manipulation, the stiffness should be minimized during negative contact while it should be maximized during positive contact, to achieve accurate manipulation. These contact stages are encoded in (3) in the form of controllers according to (2). In this way, the controller parameters (stiffness) are optimized to conform with the different stages of the contact.\n1The restitution coefficient value \u01ebr = 1 represents a perfectly elastic collision, 0 \u2264 \u01ebr \u2264 1 represents a real-world inelastic collision. \u01ebr = 0 represents a perfectly inelastic collision.\n9427\nAuthorized licensed use limited to: San Francisco State Univ. Downloaded on June 18,2021 at 16:40:34 UTC from IEEE Xplore. Restrictions apply." ] }, { "image_filename": "designv11_5_0000200_s00202-019-00882-x-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000200_s00202-019-00882-x-Figure6-1.png", "caption": "Fig. 6 Kinematic scheme of the magnetic gear", "texts": [ "8 d1 Inner diameter of the inner rotor mm 54 d2 Outer diameter of the inner rotor mm 94 hPM Height of the permanent magnets mounted on the inner rotor mm 6 Hss Height of the steel segments mm 4 HPM Height of the permanent magnets mounted on the outer rotor mm 6 D1 Inner diameter of the outer rotor mm 130 D2 Outer diameter of the outer rotor mm 140 lstack Stack length of the CMG mm 300 \u03b41 Air gap between the inner rotor and the steel segments mm 1 \u03b42 Air gap between the steel segments and the outer rotor mm 1 Material of the permanent magnets NdFeB 35 Material of the steel segments AISI 1008 where |B| is the module of the flux density vector and \u03c6 represents the angle of rotation of the cylindrical coordinate system. Kinematic scheme of the CMG model is shown in Fig.\u00a06. Dynamic mode for both rotors is described by Eqs.\u00a0(8) and (9). where I1 and I2 are the outer and inner rotors\u2019 moments of inertia; \u03c9In and \u03c9Out are the input and output rotational speeds; kf is the rotor friction coefficient; TM1 and TM2 are the magnetic torques coupling between rotors; and TIn and TOut are the incoming and outgoing torques transmitted through the magnetic gear. Inner rotor mass moment of inertia I2 about axis of rotation is 0.027\u00a0kgm2, and outer rotor moment of inertia I1 is 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001397_1350650120942329-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001397_1350650120942329-Figure5-1.png", "caption": "Figure 5. Illustration of jet impingement depth on pinion (t1\u00bc t).", "texts": [ " The inv a sph a, b\u00f0 \u00deg is equal to the declination angle a sph at the intersection point between the outside circle of the gear or pinion and the involute inv sph , b\u00f0 \u00dep\u00bc 1 sin bp arccos cos p cos bp arccos tan bp tan p inv sph , b\u00f0 \u00deg\u00bc 1 sin bg arccos cos g cos bg arccos tan bg tan g inv a sph a, b\u00f0 \u00deg\u00bc 1 sin bg arccos cos ag cos bg arccos tan bg tan ag 8>>>< >>>: \u00f013\u00de where b \u00bc arcsin sin cos \u00f0 \u00de, \u00bc tan 1 tan n=\u00f0 cos \u00de. , n is the transverse pressure angle and normal pressure angle at pitch radius, respectively. 2. At the moment (t1\u00bc t): the position parameters of the spiral bevel gears and the jet streamline are as illustrated in Figure 5. As the flowing time of the jet steam is equal to the rotation time of the pinion, which is rotating from the angle p1 at the initial time t0 to the angle p2 at time t1, the impingement depth dp can be calculated as rI2i \u00bc rIa dp cos 2 ap 2 \u00bc L2 p \u00fe r2 \u00f014\u00de or dp \u00bc rIa L2 p \u00fe r2 1=2 =cos2 ap \u00f015\u00de where L denotes the impingement distance. And rIa \u00bc 1 2 Np mI n cos I \u00fe hIap and hIap \u00bc h a \u00fe x1 mI n: Since the jet time is equal to the rotation time, then p2 p1 !p \u00bc h Vj \u00f016\u00de where p2 \u00bc tan 1 Lp r \u00fe inv p2 sph p2, b p inv p2 sph p2, b p \u00bc 1 sin bp arccos cos p2p cos bp arccos tan bp tan p2p p2p \u00bc p \u00fe tan 1 hIap dp cos ap RI e 8>>>>>>>>>>>>>< >>>>>>>>>>>>>: \u00f017\u00de where Vj is the oil jet velocity; inv p2 sph p2, b p rep- resents the declination angle p2 sph at the impingement point in the pinion involute" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001495_j.chroma.2020.461540-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001495_j.chroma.2020.461540-Figure1-1.png", "caption": "Fig. 1. (A) OP-SMD (Optan 235 nm LED by Crystal IS) (B) assembled detector. 1: OP-SMD on a starboard; 2: 3D-printed lens adapter; 3: full-ball lens; 4: half-ball lens; 5: adapter for OP-BL LED; 6: Agilent Z-cell; 7: adapter for SiC-PD; 8: SiC-PD; 9: Al housing; 10: cooling fan).", "texts": [ " A test standard mixture of 13 compounds ranging from 50 o 200 mg/L (ppm), including uracil, theophylline, 2-acetylfuran, cetanilide, m-cresol, acetophenone, propiophenone, benzofuran, utyrophenone, valerophenone, hexanophenone, heptanophenone, nd octanophenone, was prepared for chromatographic separaions. .2. Instrumentation An OP-SMD (OPTAN 3535 series) LED with a maximum emision wavelength of 235 nm was provided by Crystal IS (Green Isand, NY, USA). 24 The OP-SMD LED (3.5 \u00d7 3.5 mm) was mounted n a 20 \u00d7 20 mm star board ( Fig. 1 A). The key components f the UV-LED flow-through detector consisted of the OP-SMD ED, a TOCON_ABC2 silicon carbide (SiC) UV broadband photodide (PD) with an integrated transimpedance amplifier (Sglux Solel Technologies GmbH, Berlin, Germany) and an Agilent CE highensitivity cell (Part No.: G160 0-60 027) (Agilent Technologies, Palo lto, CA, USA). A previously described Al-based detector housing, originally deigned for OP-BL and other optical components [26] , was used o house the new OP-SMD, and weighed 70 g in total (shown chematically as Fig. 1 B, and within photograph as Fig. S2). A Sirocco YX2500 fan (5V DC) was attached onto the Al-housed etector for active heat dissipation. For the detector sensitivity study, an AL-10 0 0 syringe pump 12V DC) equipped with a Trajan SGE 1-mL Luer-lock glass syinge (Trajan Scientific and Medical, Ringwood, Australia) was used. Labsmith AV303 6-port injector valve (Labsmith, Livermore, CA, SA) fitted with a 10 cm \u00d7 50 \u03bcm I.D. PEEKsil capillary (360 \u03bcm .D.) loop, provided an injection volume of 296 nL and was used o inject samples onto the capillary liquid chromatography (capLC) olumn for analysis" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002566_j.ymssp.2021.107816-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002566_j.ymssp.2021.107816-Figure1-1.png", "caption": "Fig. 1. Performance analysis of a regular PMDC machine based on working torque-velocity characteristic chart, (a) Four operating quarters of a PMDC described using the torque-velocity chart, (b) rendering suitable quarters of current-position chart for motoring (position tracking) or reverse barking (torque tracking) control when rotational velocity is positive and (c) rendering suitable quarters of current-position chart for braking (torque tracking) or reverse motoring (position tracking) control when rotational velocity is negative.", "texts": [ " In other words, if eh is determined, ei will be a function of the minimum required current for tracking the desired position trajectory accompanied by a given acceptable tolerance. From the fundamentals of power system theory, the performance of a PMDC motor can be categorized using the currentvelocity plane that consists of four quarters and, based on the signs of the current and velocity, it is divided to motoring (i > 0;x > 0), braking (i > 0;x < 0), reverse motoring (i < 0;x < 0) and reverse braking (i < 0;x > 0), (Fig. 1-a). The operation of the PMDC motor in any of the aforementioned quarters is contingent upon external loading on the shaft and also the inertial momentum of the rotation. When the PMDC is observed to be in a condition that the position tracking error is sufficiently small relative to the torque tracking error, it can be concluded that no intense external loading exists which subsequently implies that braking is not required and position control can be pursued. Consequently, the PMDC\u2019s control scheme will be set to position control to competently continue its motoring or reverse-motoring performance when strong braking is not required, (quarters 1 and 2 of Fig. 1-b and quarters 3 and 4 of Fig. 1-c). On the other hand, if the position tracking error suddenly becomes larger than the torque tracking error it can be stated that the PMDC machine shaft is prone to severe external loading. This severe loading significantly affects the output shaft and indicates that braking is urgently required. Correspondingly, the active control strategy must be switched to track the desired torque, (quarters 3 and 4 of Fig. 1-b and quarters 1 and 2 of Fig. 1-c). Furthermore, if the exerted external disturbance pushes the PMDC into the braking operation, the torque tracking algorithm is activated to facilitate and quicken the required braking which, desirably, results in a more precise tracking of the desired position. It is noteworthy to mention that the function of PBTC is similar to the cascade structure but contrary to the cascade control, the PBTC does not impose two additional equations to the control system and consequently, it is rational to expect more accurate and agile performance results from PBTC" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001481_tmech.2020.3022830-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001481_tmech.2020.3022830-Figure1-1.png", "caption": "Fig. 1. SHER and its five degrees of freedom. The first three axis indicate the translation motion and the last two axis represent the rotational motions around the associated axis. The handle frame is shown in the close-up view of the robot end-effector.", "texts": [ " The benefits of robotic intervention include but are not limited to enhancing tool tip position accuracy, suppressing involuntary hand tremor, and provision of various sensing capabilities during the surgery. Robots developed for the purpose of microsurgery can be classified into two main groups of collaborative and tele-operated systems. In the collaborative category, surgeon and robot synergistically share control of the surgical tool. The Steady-Hand Eye Robot (SHER), developed at the Johns Hopkins University, is an example of a collaborative robot [6], [7] (Fig. 1). Other examples of collaborative robots are referenced in [8], [9]. Furthermore, literature documents other studies (e.g., [10]\u2013 [16]) representing recently developed tele-manipulated robots for eye surgery. The most clinically advanced use of robots in eye surgery to date is attributed to the first in-human robotassisted eye surgeries by Edwards et al. [17] and Gijbels et al. [18]. A review of the literature indicates that the focus of the above-mentioned robots have been mostly on hardware development (mainly for providing a tremor-free manipulation) while less attention has been paid to sensing and control, which are required for a reliable and safe robotic assist", " One drawback of this approach, therefore, is the insertion depth estimation instability when the sclera force is small and noisy. Additionally, when the tool shaft is not in contact with the sclera, no insertion depth measurement will be available. These are the present limitations associated with this method of measuring the insertion depth. Of note, using (1) and (2), the presented dual force sensing tool in this study is able to measure sclera force components and insertion depth with an accuracy less than 1 mN and 0.5 mm, respectively [23]. The SHER, as depicted in Fig. 1, is a 5 degrees of freedom (DoF) collaborative robot designed to provide steadier manipulation with less tremor in eye surgical interventions. Additionally, using a fast release mechanism for the robot end-effector, various surgical tools can be integrated with the SHER [7]. Surgeons can grasp the tool handle, which is also held by the robot, to move it to a desired location and orientation in a collaborative way with the robotic assistance. A schematic plot for the tool integrated to the robot endeffector is represented in Fig. 2. As it can be seen, when a tool is inserted into the eye through the sclera, the tool shaft may deflect due to the relative motion between the eyeball and the robot, resulting in sclera forces Fsx and Fsy . Both of these sclera force components are perpendicular to the original (unbent) tool shaft direction as shown by the dashed lines in Fig. 2. As shown in Fig 1, three coordinate frames are incorporated in the robot modelling. The robot base frame is a fixed frame placed next to the robot base while the handle frame and the tip frame are rigidly attached to the robot end-effector. The tip frame is always located at point o (Fig. 2) with z direction along the original (unbent) tool shaft direction. Of note, the tip frame will be used throughout the rest of the paper, particularly in deriving the equations for the KF. The robot control works based on a proportional admittance control which is looped around a low-level embedded joint velocity controller", " Restrictions apply. 1083-4435 (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. elements of Vd correspond to the linear velocity of the origin of the handle frame and the angular velocity of the handle frame, respectively. This desired velocity is generated based on Fh \u2208 R6 representing the forces and torques applied by the surgeon to the tool handle (Fig. 1), such that the surgeon will have an intuitive sense of movement when manipulating the robot. It is noted that the handle frame and the tip frame have a fixed offset along the z direction of the tip frame and they are always aligned. V handle d = KFh \u03b8\u0307d = J\u2020handle(\u03b8)V base d (3) where the matrix K is a diagonal 6\u00d7 6 matrix including the admittance control gains. The variable \u03b8 \u2208 R5 represents the joint angles. The desired joint angles velocity \u03b8\u0307 is then produced using the pseudo inverse of the handle frame Jacobian, Jhandle \u2208 R6\u00d75, which gives the least squares solution for \u03b8\u0307d" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002870_s00500-021-06159-5-Figure10-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002870_s00500-021-06159-5-Figure10-1.png", "caption": "Fig. 10 Various rubber bushing designs (Hei\u00dfing and Ersoy 2015)", "texts": [ " They might also feature an internal crush tube that protects the bushing from being crushed by the fixings that hold onto a threaded spigot. (Bushing (Isolator)\u2014Wikipedia, n.d.). Standard rubber bushings are also used to mount chassis and suspension components such as the control arms, dampers and subframe. Hydraulic bushings and sliding bushings can also be used for these functions. The latter feature a sliding bearing between the rubber and the mounting surface to allow free rotation and rubber damping (Hei\u00dfing and Ersoy 2015). Some of the different types of rubber bushings are shown in Fig. 10. A wide range of different solutions is available depending on the specific application\u2019s requirements. Design A represents the most straightforward and most economical solution. This type of bushing is designed to be pressed into a cylindrical bore in a suspension link or control arm, where it is held in place by radial prestress and integrated axial retaining rings. Design B allows the component\u2019s radial and axial spring rates to be more accurately tuned to match each other. By adjusting the diameter of the outer tube and the height-to-width ratio of the elastomeric inner bushing, the rubber can be preloaded to generate a spring rate ratio cradial/caxial of 1/4 to 1/10" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000427_0954410016643978-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000427_0954410016643978-Figure3-1.png", "caption": "Figure 3. Thin-layer element method and improved thin-layer element method.", "texts": [ " Using the thin-layer element method, the joints are modeled by a layer of solid elements with isoparametric material properties, which may change the joint stiffness. In order to overcome this drawback, the improved thin-layer element method is presented and can consider the interface contact stress distribution of the bolted joints adequately. Based on this, the partitioned thin-layer elements are solid elements with the bolted joints characteristics, and the thin-layer elements in different areas have different material properties, as shown in Figure 3, wherein the yellow area represents the contact interface, and the circle areas represent the different bolts. In Figure 4, the arbitrary node displacement in 8-node hexahedron element is denoted by x \u00bc X8 i\u00bc1 Ni xi y \u00bc X8 i\u00bc1 Ni yi z \u00bc X8 i\u00bc1 Ni zi \u00f01\u00de at RMIT University Library on May 19, 2016pig.sagepub.comDownloaded from where Ni is the shape function of the hexahedron element and xi, yi, and zi are the coordinate values of the ith node. According to the basic equations of the theory of elastic mechanics, the relationship between the element strain and the nodal displacement is ef g \u00bc B\u00bd aef gT\u00bc B1 B8 a1 a8 T \u00f02\u00de where B\u00bd is the geometric element matrix and aef g is the node coordinate value vector in the element", " Yi \u00bc Zi 0 \u00f015\u00de So the axial stiffness of the bolted joints structure is simplified as kN \u00bc FN= l \u00bc Xk m\u00bc1 I V0 \u00f0A1xAjxc11\u00dedV0 X1 l \u00fe \u00fe Xk m\u00bc1 I V0 \u00f0AixAjxc11\u00dedV0 Xi l 8>>>< >>>: 9>>>= >>>; 1 \u00fe Xk m\u00bc1 I V0 \u00f0A\u00f0i\u00fe1\u00dexAjxc11\u00dedV0 X\u00f0i\u00fe1\u00de l \u00fe \u00fe Xk m\u00bc1 I V0 \u00f0AhxAjxc11\u00dedV0 Xh l 8>>>< >>>: 9>>>= >>>; 2 \u00fe Xk m\u00bc1 I V0 \u00f0A\u00f0h\u00fe1\u00dexAjxc11\u00dedV0 Xh\u00fe1 l \u00fe \u00fe Xk m\u00bc1 I V0 \u00f0AnxAjxc11\u00dedV0 Xn l 8>>>< >>>: 9>>>= >>>; 3 \u00f016\u00de In equation (16), the first part represents the thin-layer elements of different bolts (circle area shown in Figure 3); the second part denotes the thin-layer elements of the contact interface (yellow area shown in Figure 3); the third part is the nonthin-layer elements of the remaining structure. The axial stiffness of the bolted joints structure is associated with the axial elastic modulus, the element shape functions and the node displacement. For a certain structure, the section size and the FE element type are determined. Therefore, the axial stiffness kN is determined by the axial elastic modulus of the thinlayer elements in different areas (circle area and yellow area). Bending stiffness. Based on the axial stiffness of the bolted joints structure, the equivalent bending stiffness is obtained when the bolted joints structure suffers from the bending moment" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000890_ilt-04-2019-0125-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000890_ilt-04-2019-0125-Figure1-1.png", "caption": "Figure 1 Schematic sectional view of elastic ring squeeze film damper", "texts": [ " In this paper, a numerical model is proposed to investigate the dynamic characteristics of the ERSFD by the fluid\u2013structure interaction (FSI) approach. The contact behavior between the journal and the elastic ring is investigated based on the assumptions of Greenwood andWilliamson (Popov, 2010). Block iterative method is used to couple the governing equations of the films and the elastic ring. The force coefficients of the ERSFD are identified and comparedwith that of the SFD. The schematic sectional view of the ERSFD (Wang et al., 2017) is shown in Figure 1. Compared with the conventional SFD, the ERSFD is identified by an additional metal ring, with bosses on both sides, located in the lubrication gap between the journal and the housing. The bosses on the ring divide the original lubrication gap into several small chambers. These film chambers between the journal and the elastic ring are called inner film; otherwise, they are called the outer film. According to the geometrical features of the ERSFD, its theoretical model can be simplified as one elastic plate interacting with two films" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002356_978-3-030-50460-1-Figure27.5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002356_978-3-030-50460-1-Figure27.5-1.png", "caption": "Fig. 27.5 (a) Nominal circular arch. (b) A generic sample of the random polycentric arch.", "texts": [ "2) r\u0303i = E[r\u0303i] + \u03c7rp\u0303ri = r(1 + \u03c7p\u0303ri) (27.3) \u03b1\u0303i = E[\u03b1\u0303i] + \u03b5\u03b1/nsp\u0303\u03b1i = \u03b1/nsp\u0303\u03b1i = \u03b1/ns(1 + \u03b5p\u0303\u03b1i ) (27.4) 27 Effect of Shape Uncertainties on the Collapse Condition of the Masonry Arch 459 in which ns indicates the number of arc sectors outlining the polycentric arch. The introduction of the parameter \u03b2\u0303i entails that the centre C1 falls on a position not belonging to the line joining the C0 and the first point of the polycentric arch. On the other hand, the thickness (t) and the span (l) are defined with their nominal values (Figure 27.5). The deterministic value of the span represents a geometrical constraint necessary to make sets of arch samples comparable according to the ratio l/t. Moreover, this condition finds confirmation in the assumption that, usually, the span between the two piers is a predefined property, independently from the uncertainties related to the construction process. Considering the established geometrical rules, the probabilistic model generates the curved line passing through the centres of the joints. Since the thickness is a constant feature, its offsets outline the extrados and the intrados of the final arch profile" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001656_tim.2020.3046051-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001656_tim.2020.3046051-Figure4-1.png", "caption": "Fig. 4. Rolling bearings with different faults.", "texts": [ " 3, it consists of a singlephase asynchronous motor as the power output, a heavy load of 5 kg, and the test bearing on the right-hand side of the rig. In this test, two speed conditions of 300 and 600 rpm are used. The signals collected by the current sensor are stored using the CoCo80 (Crystal Instruments) dynamic signal analyzer. The sampling time is 180 s, and the sampling frequency is 51.2 kHz. In this study, seven ER16K bearings with different fault types and fault degrees are applied, as shown in Fig. 4. They are normal, outer fault (OF) with 2, 4, and 6 mm in length and Authorized licensed use limited to: University of Exeter. Downloaded on May 30,2021 at 01:21:19 UTC from IEEE Xplore. Restrictions apply. inner fault (IF) with 2, 4, and 6 mm in length, respectively. For the collected current signal, we first perform the frequency reduction operation and then use the time shift window to divide it. The window length is fixed at 5120, and the moving step is 256. The total number of samples is 4760, and each category is 680" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002002_012076-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002002_012076-Figure1-1.png", "caption": "Figure 1. A Formula 1 style wheel", "texts": [ " Series: Materials Science and Engineering 872 (2020) 012076 IOP Publishing doi:10.1088/1757-899X/872/1/012076 overall performance, polyethylene plastic has been used. Plastic has very good damping properties and is a low-density material. The use of plastic contributes to the overall strength and stiffness of the tire [5] A circular looking block made up of a durable and hard material that has a bored hole through which an axle is passed and is rotated about the same center when torque is applied to it is called a Wheel. A typical tire model shown in figure 1. Wheels are the main invention which assisted mankind to transport heavy loads from place to place. Earlier logs were used. With recent inventions, tires of rubber have been designed and made lightweight by the use of Pneumatic tires. But the use of Pneumatic tires comes with its shortcomings like three Bursts, deflation, etc. Additionally plastic has also become a threat to modern-day society. To tackle both the problems we have chosen to replace the traditional Pneumatic tires with plastic mainly polyethylene" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000437_s12206-015-1235-2-FigureA.1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000437_s12206-015-1235-2-FigureA.1-1.png", "caption": "Fig. A.1. The diagram of contact cones.", "texts": [ " Venner, Multigrid solution of the ehl line and point contact problems, Ph.D. Thesis, University of Twente, Netherlands (1991). [29] J. I. Pedrero, M. Pleguezuelos and M. Mu\u00f1oz, Critical stress and load conditions for pitting calculations of involute spur and helical gear teeth, Mechanism and Machine Theory, 46 (4) (2011) 425-437. Appendix As shown in Fig. 1, the meshing of helical gears can be transformed approximately as two cones with opposite directions. Each cone is studied separately as shown in Fig. A.1. The radius of the end side of the two cones read 1 1 1 1 1 2 2 1 2 2 tan tan . d b d b r N K r r N K r q q \u00ec = =\u00ef \u00ed = =\u00ef\u00ee (A.1) where \u03b81, \u03b82 are defined in Fig. 1. When the end side point K1 locates at the pitch point P, we have \u03b81 = \u03b82 = \u03b1n, where \u03b1n is the pressure angle; rb1, rb2 are the radius of base circle. The rotational radius of the contact point K for the two cones read 1 1 2 2 sin sin . d b d b r r y r r y b b = -\u00ec \u00ed = +\u00ee (A.2) where y is the coordinate value of point K in the y direction. The relationship of the geometry parameters are shown in Fig. A.2. Then the velocity of point K for cone 1 can be calculated as 1 1 .Kv r w= \u00d7 (A.3) where vk is normal to the plane YOZ, which also represents the line-of-action plane 1 1 2 2N N N N\u00a2 \u00a2 . The velocity of the cone 1 reads ( )1 1 1sin .K d bu v r y b w= = - (A.4) Similarly, as shown in Fig. A.1(b), the velocity of the cone 2 reads ( )2 2 2sin .d bu r y b w= + (A.5) So we get the velocities of the gears ( ) ( ) 1 1 1 2 2 2 sin sin . d b d b u r y u r y b w b w \u00ec = -\u00ef \u00ed = +\u00ef\u00ee (A.6) Mingyong Liu, born in 1985, obtained Ph.D. degree in the Chongqing University, China in 2013. His research fields include gear dynamic and gear tribology. He had published 5 international journal papers and 2 conference papers in international gear conferences." ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001126_j.msea.2020.138918-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001126_j.msea.2020.138918-Figure2-1.png", "caption": "Fig. 2. (a) Geometry of the specimen indicating the symmetry along all three axes, and (b) schematic of the meshed model along with the symmetric boundary conditions.", "texts": [ "1-HEA was obtained from the near-notch region of the notchtensile specimen by focused ion beam (FIB) milling process using FEI Nova 200 Nanolab Dual Beam FIB/FESEM. TEM characterization was carried out using a FEI Tecnai G2 F20 S-Twin 200 keV field emission TEM. Finite element modeling of notch-tensile behavior was performed in a commercially available software ABAQUS to understand the stress distribution around the notch for all the four alloys. The modeling was performed with a 1/8 symmetric geometric model of the specimen by virtue of its symmetry along x, y and z axis as shown in Fig. 2 (a). A schematic of the meshed model is shown in Fig. 2(b). 3D C3D8R brick elements with reduced integration and hourglass control were used for the meshing. The symmetric boundary conditions in x, y and z directions were applied on faces shown in Fig. 2(b). A displacement boundary condition was applied on the face marked as A in Fig. 2(b). An extremely refined mesh, as shown in the inset of Fig. 2(b), was generated around the notch in order to avoid the mesh distortion due to substantial plastic deformation which would develop around the notch. All the alloys were S. Sinha et al. Materials Science & Engineering A 774 (2020) 138918 assumed elastic-plastic with isotropic hardening and the post yield stress-plastic strain response of the materials obtained from the respective quasi-static tensile test was input in the model. Fig. 1 (b) shows the tensile load-displacement curves of the notched and un-notched specimens of Al0" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002583_0957456521999836-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002583_0957456521999836-Figure3-1.png", "caption": "Figure 3. Schematic diagram of machine fault simulator setup.", "texts": [ " The rotor is supported on two rolling element bearings. The rotor system considered in the present study is supported on rolling element bearings, which generally exhibit the same stiffness in the vertical and horizontal planes. The shaft is made of steel and its length between the bearings is 0.43 m, whereas the mass of the centrally mounted disc is 0.68 kg. The mass density of the steel shaft and Young\u2019s modulus have been considered as 7800 kg/m3 and 210 GPa, respectively. The schematic diagram of the MFS is shown in Figure 3. This setup allows the study of unbalance of the rotating shaft having a circular disc at its middle. The unbalance in the rotor is artificially produced by inserting the screw (or eccentric mass) in one of the holes of the disc at a particular radial distance and angle. The rotor disc has 24 holes/slots where the eccentric mass can be inserted to produce the unbalance. Figures 4 and 5 show the rotor in healthy and unbalanced condition, respectively. The present study mainly includes experimental work for unbalance identification. However, it is important to know the critical speed of the rotor system, before operating it at different speeds. The schematic of the rotor system has been given in Figure 3. Critical speed for the rotor system has been calculated analytically. For this purpose, maximum deflection of the rotor shaft has been calculated, which is then used to calculate the natural frequency of the rotor shaft in its fundamental mode. The rotor shaft is considered to be simply supported. The critical speed of the rotor shaft corresponding to its first bending mode is found out to be 5766 r/min. Therefore, it is decided to operate the rotor system at subcritical frequencies to avoid any unwanted conditions" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000093_j.engfailanal.2019.104180-Figure12-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000093_j.engfailanal.2019.104180-Figure12-1.png", "caption": "Fig. 12. A view of the loading bridge\u2019s connection to the excavator at the moment of accident.", "texts": [], "surrounding_texts": [ "Numerical analysis of the event has been conducted with the use of the finite element method. A series of numerical models of the loading bridge and its assembly with the excavator have been prepared. They have been built based on the technical documentation of the machine and measurements conducted on the real object. Elements of the caterpillar undercarriage have been modeled by applying adequate boundary conditions. The discrete model has been composed of both one- and two-dimensional finite elements. The areas located far from the damage have been modeled with the use of beam elements and the damaged regions with shell elements. The model of the loading bridge has been positioned in the same way relative to the excavator as at the moment of failure. It has also included the location of various subassemblies fixed to the load-carrying structure, such as operator\u2019s cabin, transformer stations, drive elements, mechanical elements of a conveyor located on the bridge, and moveable booms of the loading bridge. The elements of the loading bridge that have not been included in the geometrical model have been modeled with the use of mass elements. Thanks to that, the mass of the object and the location of its center of gravity conforming with the machine\u2019s documentation have been obtained. Figs. 11 and 12 present the discrete model with the location of mass elements. The presented discrete models have been used to perform a numerical analysis of the crawler-mounted loading bridge and the part of the excavator\u2019s undercarriage. Calculations have been conducted for two cases in which different loads were acting upon the loadcarrying structure of the crawler-mounted loading bridge. The machine\u2019s load was reflected by static loads (elements that are fixed to the structure of the excavator), the load exerted by mining product located on belt conveyors, and the load exerted as the result of encrustation. The crawler-mounted loading bridge was also under the force caused by the tilt of the machine. The last load was caused by displacement. Table 1 presents in which cases the particular load was taken into account. Calculations have been conducted for two cases: loss of stability and normal operation. All loads for the normal operation case were defined according to the standard [6]. Table 2 presents the results of analyses represented as the maximum values of stress. Figs. 13 and 14 presents regions for which their yield strength has been exceeded." ] }, { "image_filename": "designv11_5_0000685_1077546318818694-Figure11-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000685_1077546318818694-Figure11-1.png", "caption": "Figure 11. Drive train of a wind turbine.", "texts": [ " In Figure 9, the first EM with maximum margin factor corresponds with the one in Figure 8(j). The Fourier spectrum of this EM is shown in Figure 10(b), which occupies a narrow frequency band from 2,800 Hz to 3,300Hz. Then the EM with maximum margin factor is processed by Hilbert transform, and its envelope spectrum is shown in Figure 10(c). The FCF 90Hz and its harmonics are evident, verifying the effectiveness of PEWT combining with margin factor in fault characteristic extraction of the simulated bearing. The drive train of a wind turbine is shown in Figure 11. The tested wind turbine is a double fed induction generator with its rated power 1.5MW. The faulty bearing is at the drive end of the generator. To monitor the health state of the critical subassemblies, seven piezoelectric acceleration sensors are respectively installed on the exterior of the main bearing, gearbox and generator. We analyze the signal of sensor 6 from the bearing at the drive end of the generator. The sampling frequency is 25,600 Hz and the FCF of this bearing is listed in Table 2" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000113_j.finel.2019.103319-Figure12-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000113_j.finel.2019.103319-Figure12-1.png", "caption": "Fig. 12. Pinched cylinder test (R\u00bc 300.0; L\u00bc 300.0; thickness\u00bc 3.0; E\u00bc 30.0E6; \u03bd\u00bc 0.3; F\u00bc 0.25).", "texts": [ " The displacement along y axis of point \u201cP\" (Fig. 10) was collected. As shown in Fig. 11, QUAD elements perform much better than the TRIA meshes. Moreover, when using QUAD elements, the membrane sed displacements of point \u201cP\u201d. locking did not occur because the simulation converged to the normalised reference solution just after the first refinement of the mesh (\u201cQUAD 8 8\u201d in Fig. 11). The pinched cylinder was one of the most severe tests to check the ability of elements to model both inextensional bending and complex membrane states (Fig. 12). It was consisted of a cylinder with rigid end diaphragms (u\u00bc v\u00bc \u03b3\u00bc 0) and loaded with two opposite forces, causing a bending-membrane coupling. Bending is just located in the areas close to the loading point, while membrane strains occurs far from the edges. Because of the symmetry of the problem, just one eighth of the cylinder was modelled using symmetry boundary constraints (same number of elements was assigned on each side). Displacements were calculated at the point of application of the load in the direction of the force" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000286_b978-0-08-102884-1.00001-7-Figure1.3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000286_b978-0-08-102884-1.00001-7-Figure1.3-1.png", "caption": "FIG. 1.3 The face-milling process, where the workpiece is being reduced in height by d, f is the feed speed of the cutter relative to the workpiece, W is the depth of the cut, and N the rotational speed of the milling cutter.", "texts": [ " While the forces will be considerably smaller than those experienced by the spindle, they still must be quantified during any design process. The locations of the forces at the point of cutting also shown in Fig. 1.2; their magnitudes are, in practice, a function of the approach and cutting angles of the tool. Their determination of these magnitudes is outside the scope of this book, but it can be found in texts relating to machining processes or manufacturers\u2019 data sheets. In a face-milling operation, the workpiece is moved relative to the cutting tool, as shown in Fig. 1.3. The power required by the cutter, for a cut of depth, W, can be estimated using; Power \u00bc d f W Rp (1.3) where Rp is the quantity of material removed in m3min 1 kW 1and is a function of the cutter speed and tooling provided by the tool manufacturer. The determination of the cutting forces is outside the scope of this book, because the resolution of the forces along the primary axes is a function of the cutting angle and of the path of the cutter relative to the material being milled, and hence reference should be made to the data provided by the tool manufacturer" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000327_j.mechatronics.2015.06.014-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000327_j.mechatronics.2015.06.014-Figure4-1.png", "caption": "Fig. 4. Force analysis of constraint limb.", "texts": [ " Fdx \u00bc nT i f i \u00fe L x L mLig L x L mLi aLi \u00f011\u00de The axial deformation at point Bi can be obtained by the integration of Eq. (11) along link rod, which yields di \u00bc Z L 0 Fdx A E dx \u00bc nT i A E f i L\u00fe L 2 mLig L 2 mLi aLi \u00f012\u00de where ni is the unit direction vector of rod AiBi. Rearranging Eq. (12) yields di \u00bc f i \u00fe f id0 =ki \u00f013\u00de where f i \u00bc nT i f i \u00bc nT i S \u00fe i F is0 nT i S \u00fe i G T JAi si is the magnitude of force f i along axis, and f id0 \u00bc nT i 1 2 mLig 1 2 mLiaLi 1 12 L mLixLi xLi ni\u00f0 \u00de\u00de; ki \u00bc A E L represents the axial tension/compression stiffness. As shown in Fig. 4, the constraint limb bears gravity mzu g; mzd g and inertia force and moment mzu azu; mzd azd applied on the upper and down rods. Also it bears force f z and moment f r applied on the down rod from the moving platform. For the UPU constraint limb, the force f z produces the limb\u2019s movement while the moment f r , whose orientation is in accordance with direction of upper and down rods, produces the limb\u2019s deformation. Using rigid body dynamics we can acquire the dynamic model of the constraint limb" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001581_s1003-6326(20)65413-9-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001581_s1003-6326(20)65413-9-Figure2-1.png", "caption": "Figure 2 Combustor and burner arrangements.", "texts": [ " The thermal structure of these catalytic flames developed over Pt and Pd catalytic disc burners was examined by measuring the mean temperature distribution in the radial direction at different axial locations along the flames. Also, the axis\u2013symmetric distributions of CO and CO2 along the flames were monitored under the same conditions to clarify the catalytic combustion process performance over the two current catalytic discs. The experimental setup (Fig. 1) was comprised of a vertical cylindrical combustion chamber filled with an arrangement supplying fuel and air. The combustion chamber (Fig. 2) is 150 mm in diameter, 5 mm thick and 1.0 m long. The combustor was equipped with a thermal resistant glass window. The fuel jet was discharged vertically through a nozzle of 2.5 mm diameter connected at the centre of the fuel supply line in the axial direction at the base of the combustor. Commercial LPG fuel having an average composition of: 23% propane, 76% butane, and 1% pentane was used in all experiments. Two catalytic disc burners of Pt and Pd over c-Al2O3 having a diameter of 40 mm, 4 mm thick and perforated with 25 holes had been separately used as a catalytic flame burner" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002426_j.engfailanal.2020.105211-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002426_j.engfailanal.2020.105211-Figure1-1.png", "caption": "Fig. 1. Wheel-hub unit construction. 1. Knuckle 2. Bearing 1# 3. Hub 4. Bearing 2# 5. Preload nut.", "texts": [ " Under three typical driving conditions, N-R iterative method and commercial software RomaxDesigner simulation results are compared to verify the effectiveness of N-R iterative method in calculating bearing contact load. In Section 4, the joint distribution counting is carried out based on random road loads, and then the equivalent damage and life of bearings are calculated based on linear damage theory and L-P theory. Section 5 has carried on the experiment verification. Finally, some conclusions are given in the Section 6. L.-H. Zhao et al. Engineering Failure Analysis 122 (2021) 105211 Fig. 1 shows the hub unit structure diagram of a certain type of commercial vehicle driven wheel. The hub unit uses two SKF singlerow tapered roller bearings (TRB) of different models to pair. The preloading force is applied to bearing 1 (Type 30,313 TRB) through the outer preloading nut, and the preloading force is transferred to bearing 2 (Type 30,311 TRB) through the hub to achieve preloading. The geometric parameters of the bearings are shown in Table 1. In this paper, a six-dimensional force transducer was installed on the test vehicle, the vehicle speed in the test range is within the range of 10\u201350 km/h, and the maximum speed is 72 km/h" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000976_s11044-019-09704-1-Figure40-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000976_s11044-019-09704-1-Figure40-1.png", "caption": "Fig. 40 Load distribution of thin-walled four-point contact ball bearing under radial and axial load", "texts": [], "surrounding_texts": [ "The dynamic performance of thin-walled four-point contact ball bearing is affected by the static and dynamic loads of the wrist, hand, and workpiece or tools of the robots, especially the impact force brought by inertial force and moment under high velocity or acceleration. Considering multi-clearances and multibody contacts, the dynamic contact and impact of ball-to-ring raceway surface contacts, ball-to-cage pocket surface contacts, and cage-toouter ring guidance surface contacts of the bearing are investigated by using the multibody contact dynamic analysis models in this work. The proposed models of the bearing are solved by HHT algorithm with ADAMS software package. The results of dynamic contact forces, impact forces, displacements, and velocities are discussed under the conditions of different complex loads. Due to the effect of multi-clearances and multibody contacts, the impact forces of ball-to-cage small pocket surface contacts are larger in the non-load zone under the conditions of pure radial load or combined rotating radial load, axial load, and positive moment. The dynamic characteristics will be underestimated when ignoring the effect of multibody contact dynamics of the bearing using the empirical design method. The rules of static load distribution and the simulation model of the bearing are verified by the theoretical value. Considering pure radial load, there\u2019re always four-point contacts of ball-to-ring raceway surface contacts in the load zone. The dynamic contact forces are similar to the rule of the static load distribution. The impact forces of ball-to-cage small pockets surface contacts are large, the angular velocities of ball rotation and cage are vary- ing greatly in the non-load zone. The dynamic characteristics of the bearing are stable under the condition of combined radial and axial loads. There\u2019re four-point contacts in the load zone and two-point contacts in the non-load zone. The dynamic contact force of the main contact pair is greatly different from the secondary contact pair. The maximal contact force and the angle of load distribution of the main contact pair are both larger than those of the secondary contact pair. The dynamic contact force of the main contact pair of ball-to-ring raceway surface contacts has one peak and one valley, and is nonzero in the non-load zone. The dynamic contact force of the secondary contact pair is similar to the rules of static load distribution and is zero in the non-load zone. The angular velocities of the ball rotation and cage are stable and periodically varying as a result of the small impact forces of ball-to-cage small pockets surface contacts. The moments are important to the rules of dynamic contact forces of ball-to-ring raceway surface contacts and angular velocities of the ball rotation. The dynamic contact forces of the main contact pair in the load zone are increasing as positive moments increase. The dynamic contact forces of the secondary contact pair in the load zone are increasing as the negative moments increase. The effects are an increase in the radial load and reduction in the preload. The impact forces of ball-to-cage small pockets surface contacts in the non-load zone are large as a result of the effect of rotating radial load. The motion trajectories of outer ring center are as from circular whirling motion. Under the conditions of the proposed loads, the motion trajectories of cage center are always similar to a circular motion trajectory. As a result of the slight impact forces of ball-to-cage big pockets surface contacts, the balls in the cage big pockets are always purely rolling in the load zone and slightly varying in the non-load zone. The motion stability is high and dynamics characteristics are complex for the bearing. In the proposed work, a new approach is presented and the calculated results are illustrated. The dynamics performance and motion accuracy of thin-walled four-point contact ball bearing are complicated. They\u2019re influenced by the complex load conditions, multi-clearances and multibody contacts of ball-to-ring raceway surface contacts and ball-to-cage pockets surface contacts. The influences of the geometrical parameter, multi-clearance and flexibility on multibody contact dynamic analysis, and motion accuracy of thin-walled four-point contact ball bearing will be investigated under varying complex working conditions in the future. Acknowledgements The authors would like to express sincere thanks to the referees for their valuable suggestions. This project is supported by National Natural Science Foundation of China (grant nos. 11462008 and 11002062) and Natural Science Foundation of Yunnan Province of China (grant no. KKSA201101018). This support is gracefully acknowledged. Publisher\u2019s Note Springer Nature remains neutral with regard to jurisdictional claims in published maps and institutional affiliations." ] }, { "image_filename": "designv11_5_0000753_00207179.2019.1590737-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000753_00207179.2019.1590737-Figure1-1.png", "caption": "Figure 1. Underwater towed system.", "texts": [ " The indirect variational formulation is given in the second part and the scheme for its implicit solution in the third section. Finally, the MPC method is illustrated in the fourth section followed by numerical results. A classical towed mechanical system used in ocean exploration and military applications, is composed by a towed vehicle V and a towed sonar line array (TLA) (Abreu, Morishita, Pascoal, Ribeiro, & Silva, 2016; Arrichiello, Antonelli, & Kelholt, 2016). The two subsystems are towed by the towing cable attached to the moving boat, B (Figure 1). The goal of the vehicle V (Figure 2) is to guarantee a TLA\u2019s constant depth of navigation and the reduction of the entire system oscillations (Falls, Mouring, Sridharan, & Loomis, 2015; Paifelman, 2017; Pan & Xin, 2012). V is also called \u2018depressor\u2019 (Vinod, Francis, Prabhasuthan, &Nandagopan, 2016) and itsmotion generatesmemory effects which influence the proper operation of the TLA (Feng & Allen, 2004). As already explained in Section 1, if the vehicle moves close to the free surface, memory effects called addedmass and added damping respectively, are borne" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002614_j.compositesa.2021.106449-Figure9-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002614_j.compositesa.2021.106449-Figure9-1.png", "caption": "Fig. 9. The procedures of updating tow points to incorporate the curvature of mandrel surface.", "texts": [ " In the current work presented in Section 4, seven points were used across the width of each tow. All triangular mesh processing, for the tows and mandrel, was completed in python library trimesh [25], including STL file import of the mandrel CAD surfaces. As stated previously, the tow points not on the tow centreline are extrapolated transversely using the local tow orientation and normal vectors. Hence, each tow is approximated as locally flat and not conforming to the curvature of the mandrel surface (Fig. 9). This effect is negligible for narrow tows and mandrels with limited curvature. In some cases, though, to ensure geometric compatibility and avoid excessive tow distortions, the tow points need to be corrected using the mandrel surface curvature. In the TWM algorithm, the projections of the tow points on the mandrel surface are calculated. The projections are used to remap the tow points and the detailed procedures are described in Fig. 9. Seven interpolation points are used in the transverse direction here to define the curvature of the mandrel surface. The projections can be efficiently determined through a ray-tracing algorithm embedded as a submodule X. Li et al. Composites Part A 147 (2021) 106449 of trimesh. If the ray origin and direction are specified, the intersected mesh and the specific intersection position can be calculated. The intermediate step of offsetting the tow points slightly upwards (in Fig. 9) is to ensure a robust ray-triangle intersection. Finally, the tow points will be updated by shifting the intersections along the normal direction of the mandrel surface by half tow thickness to account for the thickness effect. The updated tow points are used to represent the mid surface of the tow geometry. The resultant tow width/ length will differ slightly from the nominal dimension due to the curved surface projection, but this is a second-order error and is negligible for the majority of real problems" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002828_j.jii.2021.100265-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002828_j.jii.2021.100265-Figure1-1.png", "caption": "Fig. 1. The simulation wind of the AOLP engine.", "texts": [ " Firstly, for each layer in the deposition sequence, a valid robot trajectory must be planned so that the robot is able to deposit the layer without exceeding its kinematic limits or colliding with the surrounding environment (including the partially manufactured component). Additionally, a valid build sequence for all layers needs to be identified. These two planning processes are interwound with each other and present a challenging optimisation problem. This planning process is supported by an automated robot offline programming (AOLP) engine [34], which provides a simulation model for the robotic WAAM system, as shown in Fig. 1. With this offline programming engine, robot motions can be efficiently planned and simulated to test robot reachability and collision. While the most straightforward method is to test all possible combinations of build sequences and robot motion options in a brute-force manner, the exceedingly large number of possible permutations, particularly for more complex parts, renders it prohibitive to do so. To find an acceptable solution in a short time-frame, an efficient search algorithm that breaks the complex search problem into multiple steps, each with much lower DoFs, was developed in this study" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000639_cdc.2016.7798933-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000639_cdc.2016.7798933-Figure4-1.png", "caption": "Fig. 4. Circulant graph used in the simulation in Example 2.", "texts": [ " Indeed, for any choice of k > 0, the synchronization region is SP(k) = C>0 as illustrated in Fig. 1. The synchronization of a network of harmonic oscillators is shown in Fig. 2. Example 2: (Output-feedback-passive) Consider system P = (A,B,C) A = [ 0 1 \u22121 1 ] , B = [ 0 1 ] , C = [ 0 1 ] . P is not passive but can be made passive with the feedback gain h = 1. Lemma 2 implies that SP(1) \u2283 C>b for some b > 0. The synchronization region is SP(1) = C>1 and is illustrated in Fig. 3. We simulate a network of three systems connected by the circulant graph G depicted in Fig. 4. Its spectrum is \u03c1 (G) = {3/2 \u00b1 i \u221a 3/2}. Therefore, G \u2208 G3 >\u03b5, \u03b5 < 3/2. By setting k > 2/3 we guarantee synchronization of (P3, k,G). The dynamics of the network for values of k below, equal, and above this threshold are simulated. The synchronization error is plotted in Fig. 5. Synchronization only occurs for k > 2/3. Example 3: (Neither passive nor output-feedbackpassive) Consider P = (A,B,C) A = [ 0 1 \u22121 0 ] , B = [ 0 1 ] , C = [ \u22121 1 ] . P is not output-feedback-passive, hence, by Lemma 2, the synchronization region SP(1) does not contain the set C>\u03b5, for any \u03b5 > 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000859_1077546319856147-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000859_1077546319856147-Figure3-1.png", "caption": "Figure 3. (a) bearing showing load line and clearance; and (b) deformation in balls and races.", "texts": [ " The angular velocity of the cage is equal to the orbital velocity of the center of the ball which comes out to be Wcage \u00bc rotor ri \u00f0r0 \u00fe ri\u00de \u00f09\u00de The configuration of balls achieves same angular positions after every certain time-interval so the frequency of variation of structural stiffness of bearing is given by vc \u00bcWcage Nb \u00f010\u00de where, Nb is the number of balls in the cage. Combining equation (10) and equation (11) vc \u00bc rotor fNB; where fNB \u00bc ri \u00f0r0 \u00fe ri\u00de Nb \u00f011\u00de The fNB number in equation (12) is the ratio of frequency of self-excitation and rotational speed. In the present study, both the bearings are considered to be synchronized regarding orientation of balls. The angular position of ith ball with respect to reference x-axis is given by i \u00bc 2 Nb i 1\u00f0 \u00de \u00fe \u00f0Wcage t\u00de \u00f012\u00de The center of the inner race is subjected to displacement during motion. Figure 3 (b) shows displaced position O0 of the center of the inner race. The elastic deformation of ball and races along the radial-direction i comes out to be i t\u00f0 \u00de\u00f0 \u00de \u00bc x cos i t\u00f0 \u00de \u00fe y sin i t\u00f0 \u00de \u00f013\u00de Figure 3 (a) shows the geometrical meaning of internal radial clearance ( ) of ball bearing. In case of \u00f0 i\u00f0t\u00de\u00de 0 for any ball, there is deformation of the corresponding ball and races. Reaction force from bearing to rotor is generated only from those ball\u2013race contacts where balls are in compression, which is given by the Hertzian nonlinear loaddeflection relation F i \u00bc Cb\u00f0 \u00f0 i\u00de\u00de n; n \u00bc 3=2 \u00f014\u00de The values of Cb (N/m3/2) and n are determined from elastic analysis (Harris, 1984). If \u00f0 i\u00f0t\u00de\u00de5 0 for any ball then this ball is not under compression and therefore no restoring force is generated from such balls" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002426_j.engfailanal.2020.105211-Figure8-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002426_j.engfailanal.2020.105211-Figure8-1.png", "caption": "Fig. 8. The displacement caused by an angular misalignment.", "texts": [ " The concave curvature of the outer raceway prevents the roller from being skewed, and this resistance and subsequent deformation will change the load distribution of roller - raceway and roller - inner ring guide edge. Therefore, when establishing the bearing deformation coordination equation, the angular misalignment between the inner ring and the outer ring should also be considered. When calculating the bearing load distribution, the angular misalignment of the inner ring roller at 0\u25e6 should be taken as the reference value. The angular misalignment of the roller at other angular positions is as follows \u03b3i = \u03b3cos\u03d5i (15) As shown in Fig. 8, Point O is the center of two TRB. Mo is the center of the roller. R1 is the distance between the roller center and the bearing rotation center. R2 is the distance between point H and point O. \u03b2 is the angle between OM and ON. \u03b4\u03b3i is the total displacement caused by angular displacement. \u03b4r\u03b3i and \u03b4a\u03b3i are the radial and axial displacement components of \u03b4\u03b3i respectively. \u03b4r\u03b3i = R\u03b3isin\u03b2 = R2\u03b3i (16) \u03b4a\u03b3i = R\u03b3isin\u03b2 = R1\u03b3i (17) Besides, in the actual use of hub bearings, in order to improve the bearing performance, to apply preload on the bearing, the axial pre-deformation caused by the bearing preload Fo[22] is as follows \u0394 = ( Fo ZKn(sin\u03b1v) 2" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001290_s11071-020-05510-z-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001290_s11071-020-05510-z-Figure1-1.png", "caption": "Fig. 1 The dynamic model of a spur gear system", "texts": [ " 4, nonlinear dynamics of multi-stable behaviors in a two-parameter plane are studied systematically on the base of the designed method. The distribution and bifurcation characteristics of multi-stable behaviors in the parametric plane are studied in detail in Sect. 4.1. Double-bifurcation points (intersection of different bifurcation curves) are investigated in Sect. 4.2. Probabilities of occurrence of various hidden bifurcation curves are calculated and analyzed in Sect. 4.3. Conclusions follow in Sect. 5. The dynamic model of a spur gear system is illustrated in Fig. 1. It is reasonable to assume that the spur gear pair is supported by rigid mounts, and only the rotation of gears is considered [28,33,38]. In Fig. 1, Ti , \u03b8i , Ii and Rbi are the torque, dynamic angular displacement, mass moments of inertia and base circle radius of the gears, respectively, where i = p and g represent, respectively, pinion and gear. Dm is half of backlash. e(\u03c4 ) is the comprehensive transmission error along the line of action, where \u03c4 indicates time. cm and km(\u03c4 ) represent the meshing damping and time-vary meshing stiffness of the spur gear system, respectively. The studied gears are all standard involute spur gears, and some parameters of the gear system under consideration are shown in Table 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000200_s00202-019-00882-x-Figure19-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000200_s00202-019-00882-x-Figure19-1.png", "caption": "Fig. 19 Eddy current volumetric loss distribution, nin = 150\u00a0rpm/nout = 27.3\u00a0rpm", "texts": [], "surrounding_texts": [ "Distributions of the eddy current losses are presented in Figs.\u00a015, 16, 17, 18 and 19 for the NdFeB magnets [27, 28]. These losses are limited to electrically conductive parts only. Results correspond very well with these for electrical conductivities of permanent magnets, shown in Table\u00a02, and total losses and efficiency in Table\u00a03. Eddy currents are estimated at five rotational speeds, for inner rotor 150, 550, 1100, 1650 and 2200\u00a0rpm and for outer rotor rotational speeds 27, 100, 200, 300 and 400\u00a0rpm, respectively. Output rotor torque reduction due to eddy current demagnetization effect for different rotational speeds is shown in Fig.\u00a020. Eddy currents, neglecting the temperature changes of resistivity, are linearly related to rotational speed [29, 30]. Eddy current losses rise with the square of rotational speed. Also, they are influenced by the huge number of magnetic pole pairs. At rotational speeds greater than 500\u00a0rpm for the magnetic gear construction under consideration, the heat effect of eddy current losses could not be neglected. At speed 10,000\u00a0rpm, only the eddy current losses are overcoming 4% of the transmitted power. Eddy current volumetric loss distributions show that most of the losses are located in the low-speed rotor magnets. They are induced by high-speed rotor magnet movement and are amplified by modulating segments at high harmonics of the magnetic flux. Frequency separation of eddy current losses is important for loss analysis. 1 3 8 Magnetic field harmonic distortion of\u00a0CMG In many existing researches, eddy current losses are often ignored in steady state because of direct analogy with single-rotor electrical machines with permanent magnets. In MGs, these losses still appear because of relative movement of two rotors and modulating segments. Harmonic flux distortion increases eddy current effects. The main rotational frequencies of the magnetic field of the magnetic gear construction for the outer and inner rotors are 55\u00a0Hz and 10\u00a0Hz. The radial components of the flux density in the air gap between the inner rotor and the steel segments of the magnetic gear are shown in Fig.\u00a021. The radial components of the flux density in the air gap between the steel segments and the outer rotor of the magnetic gear are shown in Fig.\u00a022. The fast Fourier transform (FFT) analysis of the radial components of the flux density in the air gap between the inner rotor and the steel segments is shown in Fig.\u00a023. The FFT analysis of the radial components of the flux density in the air gap between the inner rotor and the steel segments is shown in Fig.\u00a024. 1 3 FFT analysis shows significant harmonic distortion. All frequency harmonics different from main 55/10\u00a0Hz are creating torques with no proper frequency of rotation, limiting the output CMG torque. According to these estimations, more than 20% of the total magnetic flux is engaged in these undesired frequencies [31]. This distortion effect does not influence the power efficiency, but it limits the CMG output torque and does not use the available flux from the permanent magnets. The harmonic distortion is not related to direct power losses because we are not using electrical supply for these fluxes, but it limits the output torque value of the CMG design. Some words on hysteresis loss are necessary, considering frequency-dependent hysteresis models, such as Steinmetz equation variations; they are suitable for loss superposition over magnetic flux amplitudes and frequency harmonics, like shown in Figs.\u00a023 and 24. However, a closer look reveals that hysteresis loss is not substantial for MG operation, as a power loss, covered by electrical excitation as it is in rotational machines. In MG, magnetic hysteresis loop causes time-dependent flux non-linearity, decreasing this way slightly the dynamic magnetic torque interaction between rotors. The summarized results for torque reduction and losses according to rotational velocity are presented in Table\u00a04. Losses are estimated for 150\u00a0rpm, 2200\u00a0rpm and 10,000\u00a0rpm of high-speed rotor. Results are showing significant dynamic torque reduction in high rotational speeds and rise of eddy current losses. According to estimated losses, efficiency mapping at MG overload, at torques above 320\u00a0Nm, is shown in Fig.\u00a022. At low-speed overload, efficiency is influenced by rotor slipping, while in high speeds it is influenced by eddy currents (Fig.\u00a025)." ] }, { "image_filename": "designv11_5_0002687_j.mtcomm.2021.102517-Figure11-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002687_j.mtcomm.2021.102517-Figure11-1.png", "caption": "Fig. 11. Schematic diagram of columnar dendrite growth in the molten pool.", "texts": [ " The cellular dendrites in the molten pool would grow almost perpendicular to the boundary against the direction of maximum heat flow, thereby making the shape of the molten pool affect the growth direction of the cellular dendrites and ultimately affect the overall texture of the sample. According to Fig. 3, the molten pools were all in keyhole mode, the angle between the direction of cellular dendrites growth in the molten pool and the horizontal line \u03b1 was significantly small, so the direction of cellular dendrites growth in the molten pool would display a large angle with the BD direction, thereby making the angle between the c-axis of magnesium and the building direction (BD) smaller in the final sample (Fig. 11). Likewise, S.D. Jadhav [42] in the SLM of pure copper and M. Higashi [43] in the SLM of molybdenum alloy reported the effect of the geometry of the molten pool on the texture. Fig. 12 illustrates TEM images of grain boundary. When the solidification occurs, the decrease in the temperature led to the decrease in the solid solubility of the magnesium matrix to the rare earth elements. Thus, the rare earth elements were discharged from the magnesium matrix, and the eutectic phase was finally formed at the grain boundary" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001800_978-3-030-40410-9-Figure8.7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001800_978-3-030-40410-9-Figure8.7-1.png", "caption": "Fig. 8.7 Structure of two beams connected by spherical joint", "texts": [ "56) After all matrices are represented to the base frame, MSA can be applied to find the stiffness of the structure of themanipulator. Themethodology described forMSA assumes that the whole structure is free. In other words, the global stiffnessmatrix for the system will be singular and cannot be inverted. To solve this problem, boundary conditions must be applied to this matrix to convert it to an invertible version to find the displacements: x = K\u22121F (8.57) Assumewe have to find the stiffness matrix of the structure shown in Fig. 8.7a, which consists of two beams connected together utilizing spherical joints. Each physical connection and point of force effect should be represented as a node. Figure 8.7b shows the possible node on the structure, where node 2 is the point of the effect of force Fy, which splits the horizontal beam into two segments. Now, the model is composed of segment 1\u20132, segment 2\u20133, joint 3\u20134, and segment 4\u20135. Note that the boundary conditions for this system are that node 1 and node 5 are fixed with zero translational and rotational displacements. It is worth mentioning that element 3\u20134, which represents the spherical joint, is only a representative scheme where nodes 2 and 3 are considered to form the joint", "9 Final overall stiffness matrix obtained by MSA Node 1 and node 5 are fixed and do not have any translational and rotational displacements. Therefore, all rows and columns related to nodes 1 and 5 should be deleted. The final overall stiffness matrix is shown in Fig. 8.9. For the purpose of applying Eq. (8.2) to find the rotational and translational compliances, we have to find the wrench vector by structural analysis of the system. Example 8.2 For the same manipulator, we may use the MSA method to calculate the global stiffness of the robot using the same procedure described for the structure in Fig. 8.7. First of all, we have to identify each element in this mechanical system, as shown in Fig. 8.10. For this example, the robot consists of 5 segments and 6 nodes starting from the ground where node 1 is attached. Node 1 is fixed to the ground; they have zero translational and rotational displacements, and also the inertial frame is attached to this point. Segment 1\u20132 is considered rigid, and it is assumed that it has a rigidity that is ten times greater than other segments. We have to identify the numbering DOFs of each node, as described in Table 8" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001365_j.cirpj.2020.05.014-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001365_j.cirpj.2020.05.014-Figure1-1.png", "caption": "Fig. 1. Powder-control diagram (not to scale).", "texts": [ " Schematic of powder mass flow rate test (not to scale). Note: As drawn, schematic shows measurement of mass flow rate for mechanicallygenerated powder (Hopper 2); mass flow rate was also measured for gas-atomized powder (Hopper 1) using this technique. the statistical significance of the difference in the mean flow rates [22]. A study of the flowability of the powders in the LENSTM process was also conducted. The flow rate of the powder in the Optomec MR7 is controlled by adjusting the RPM of the powder auger shown in Fig. 1; increasing the RPM increases the mass flow rate of the powder. The RPM of the powder auger on the DED system used in this study is controlled with an open-loop control system. This leads to inherent differences between the RPM input by the user and the actual RPM of the auger in each hopper. Since the two powders for this test were to be loaded into different hoppers, it was necessary to correlate the programmed input to a specific auger to the actual output RPM for that auger. With no powder in the hoppers, RPM settings were varied and two measurements of the actual time it took for the auger to make a revolution were taken at each setting, in each hopper; this data is provided in supplemental material" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000390_j.protcy.2016.03.068-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000390_j.protcy.2016.03.068-Figure4-1.png", "caption": "Fig. 4. Photographs and schematic diagram of strain gauges indicating their locations.", "texts": [ "1 \u00b1 1% and Gauge length \u2013 5 mm) are placed on the master leaf along longitudinal direction and a portable 5 channel strain indicator (Make \u2013 Syscon Instruments, Bangalore) is used. Locations of the SGs ( ix , iy ), i = 1, 2, \u2026, Nsg (number of SGs) with respect to the centre of the drill hole are given in table 1, both in Cartesian and polar coordinates. Individual SG readings are obtained by using half bridges for each strain gauge, formed by mounting a dummy gauge on a similar piece of material as that of the master leaf spring. As it is shown in Fig. 4, SGs 1 and 2 are located on the top surface of the master LS in left side, whereas SGs 3 and 4 are placed on the bottom surface in the same side. Similarly SGs 5 and 6 are placed on the top surface and SGs 7 and 8 are placed on the bottom surface in the right side of the drill hole. During experimentation, the applied load on the master leaf is increased until the induced bending stress equals to 75% yield stress value of the spring material. The limit load is calculated by using bending stress equation of Winkler-Bach curved beam theory", " At location x ( Lx0 ) in plane load is given by xxxxLtx VPP sincos , where x is slope of the beam profile at that location. This in plane load produces axial stress which is given by bhPtxtx . (4) Presence of circular hole in master LS gives rise to stress concentration. Overall stress distributions around the circular hole in an infinite plate subjected to nominal stress , as given by [2], are )]2cos()431()1[(2/ 242 r , )]2cos()31()1[(2/ 42 and 2/r 431[( )]2sin()2 2 , where a is the radius of hole (refer Fig. 4), ra / and txbx . Due to the state of stress ( rr ,, ) at point ( ,r ), the induced stress along x axis is back calculated using the analytical form of Mohr\u2019s circle, )180(2sin)180(2cos]2/)[(]2/)[( rrrx (5) Profiles of the master LS under each of the load steps are exported from AutoCAD\u00ae figures and stored in MATLAB\u00ae computational platform. Now locations of SGs are taken on the LS profile and at each of these sgN points shear force and bending moment are calculated from equations 1 and 2 respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002402_0016-0032(55)91068-1-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002402_0016-0032(55)91068-1-Figure1-1.png", "caption": "FIG. 1 .", "texts": [ "OTES FROM THE FRANKLIN INSTITUTE LABORATORIES FOR RESEARCH AND DEVELOPMENT NICOL H. SMITH, DIRECTOR TURBULENCE IN A TILTING-PAD THRUST BEARING BY STANLEY ABRAMOVITZ During experiments with a tilting-pad thrust bearing (see Fig. 1) using water as a lubricant, we found that friction torque increased abnormally when the rubbing speed exceeded what appeared to be a critical U RUNNER /////////////////i///////// value . To examine this in further detail, tests were conducted over a range of water temperatures (i5 to 200\u00b0 F .) and bearing velocities (25 to 100 fps .) . With the torque values obtained, a plot of coefficient of friction versus \u00b5 U/PB was made ; its deviation from the classic, straight-line, log-plotted relationship was clear " ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000619_asjc.1433-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000619_asjc.1433-Figure5-1.png", "caption": "Fig. 5. BAB System.", "texts": [ " According to Lyapunov equation Ai BiLsj TeP \u00fe eP Ai BiLsj \u00bc Qi; where Qi \u2265 0 and eP \u2265 0, and using the Schur complement, equation (38) can be written as: _V x\u00f0 \u00de \u00bc \u2211 r i\u00bc1 hi z\u00f0 \u00de\u2211 r j\u00bc1 hj z\u00f0 \u00de\u2211 qi s\u00bc1 ms j J t\u00f0 \u00de\u00f0 \u00de xT eSTh i Qi ePBiNs jePBiNs j T 0 24 35 xeS \u2264 0 (39) This result implies that the system trajectory could reach the sliding surface in finite time and maintain on the sliding surface [12]. IV. SIMULATION RESULTS In this section, we examine the performance and effectiveness of the proposed controller developed above. 4.1 Example 1 We consider the ball and beam (BAB) system, shown in Fig. 5 [31], which is used as standard test for many control techniques [32]. The aim of this work was to stop the movement of the ball at any distinct location. In BAB system, the ball falls quickly and this makes the control very difficult. The dynamics of the BAB system are given as: \u20acr t\u00f0 \u00de \u00fe g sin\u03b8 t\u00f0 \u00de \u00fe \u03b2 _r t\u00f0 \u00de r t\u00f0 \u00de _\u03b82 t\u00f0 \u00de \u00bc 0 r2 t\u00f0 \u00de \u00fe 1\u00f0 \u00de\u20ac\u03b8 t\u00f0 \u00de \u00fe 2r t\u00f0 \u00de_r t\u00f0 \u00de _\u03b8 t\u00f0 \u00de \u00fe gr t\u00f0 \u00de cos\u03b8 t\u00f0 \u00de \u00bc \u03c4 t\u00f0 \u00de ( where r(t) is the position of the ball; \u03b8(t) is the angle of the beam; \u03c4(t) is the torque that moved the beam; g is the gravity constant and \u03b2> 0 is the friction coefficient" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0003110_520237-Figure11-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0003110_520237-Figure11-1.png", "caption": "Fig. 11 - Actual shape of tooth contdct as determined by etching tests", "texts": [ " It is shown (Appendix D) that: mo 3 TOO3 + 2 V(m0 2 - 4)3 where C, = Ratio of load carried on line of contact VW to total load mo = Mismatch contact ratio = V mF* + m,2 mp = Face contact ratio m, = Profile contact ratio 316 SAE Quarterly Transactions Note that equation (1) is to be applied when there are always at least two teeth in contact. (This is the case when m0 is greater than 2. When m0 is less than 2, all the load will be carried on one tooth and C, will have a value of 1.) Position of Point of Load Application The next step is to consider at what position on the tooth a concentrated load is to be assumed. The true shape of the area of contact on a spiralbevel-gear tooth, as determined by etching tests made at General Motors Research Laboratories,\" will be approximately as shown in Fig. 11. This differs from the previous assumption that the area was a symmetrical ellipse. Note that the center of pressure is toward the heel of the tooth. Actually, there is a distributed load, but for simplicity a point K is chosen to represent a concentrated load replacing the distributed load. Point K has been selected at a distance along the line of contact two-fifths of the way toward the heel of the tooth from the center of the line of contact O' to V. This represents an arbitrary decision based upon what seems to be reasonable thinking", " when mo < 2 (2) when MO > 21 cos 4 (cos2 $ + tan2 $I) p, = Mean normal circular pitch = Normal pressure angle ii = Mean spiral angle Layout in Normal Section The next step is to make a layout of the tooth in the mean normal section to determine the tooth 3 \"Spiral-Bevel-Gear Etching Tests,\" report from General Motors Research Laboratories December 1932. (Not published.)4 See proceeding; of Engineers' Club of Philadelphia, Vol. 10, January. 1893, pp. 16-24: \"Investigations of Strength of Gear Teeth,\" W. Lewis.5 \"Calculated Bending Stress in Spur Gears\" by A. H. Candee. Presented at Twenty-Fourth Semiannual Meeting of AGMA, October, 1941. _ strength factor. The procedure followed is basically similar to that used by Wilfred Lewis4 However, instead of applying the load at the tip of the tooth, the load is to be applied at point K' (Fig. 11). The tooth strength factor XLN takes into account only the tangential component of the load. With a high pressure angle the radial component of the load, which was neglected by Lewis, may amount to as much as 25% of the total load. It shoukId', therefore, not be neglected, especially when making comparisons between two tooth designs. It has been shown 4, 5 that the tooth strength factor Xh for the tension fillet, including the combined tan- April 1952 Vol. 6 No. 2 317 It is generally recognized today that the actual stresses in most machine parts cannot be calculated exactly", " Tangential Load Since the load has been considered concentrated at point K, which usually will not lie in the mean section, it is desirable to consider the load tangent to the circle passing through this point. This may be expressed as follows: (6) where: T = Total torque transmitted by gear C, = Ratio of load carried on nlost heavily loaded tooth to total load. The product TC, is the portion of the torque transmitted by the one tooth RG = Radius in plane of rotation to point of lond application A + .To cos y R = Mean transverse back cone distance A = Mean cone distance row = Distance from mean section to point K measured in the lengthwise direction on the tooth (see Fig. 11) Fin, when mo > 2 Fmr I\" 2m ?_ mo~ L amr when mo < 2 niF Concave ]Convex Afii.y = Distance from pitch circle to point of load application L. This value is obt,ained from the layout y = Pitch angle 318 SAE Quarterly Transactions WI = Impact load to allow for the effect of inaccuracies in the gears. For static loading WI will be zero. This item will be discussed more fully under dynamic loads Effective Face Width The stress is distributed along the length of the tooth in a complicated manner. Referring to Fig. 11, the load is distributed along the line of contact VW. An attempt to analyze the stress over the tooth length based on the load distribution shown in this diagram would involve complications not in keeping with the general attempt at simplicity.\" Certainly, if the tooth is very long and the line of contact is relatively short, it is not reasonable to consider the stress at the fillet as uniform along the entire length of the tooth. If the line VW is projected to the root of the tooth, the projected distance will be FK", " yo1 = \u2014 sin $6 = o 2~ma 5MO2 and from equation (23) in Appendix C: Yo = z 2 f COS x Z mp (2 \u2014 F2 sin2 $b + Z2 cos2 where: f = VN - /o = F Z + Z2 Substituting these values in equation (34) : 2m , iW,F 1 + -Z77 - \u2014 5m:\u21220 ' \" 0 OirtQ mo -1 - 1 J Concave (344 Conves SAE Quarterly Transactions When the contact ratio is sufficiently large such that m0 will be greater than 2, the line of contact VW will pass through the center 0. In this case f = 0 and y0 = 0. and: Z r 2 [ 211 1~ Concave 5m0 J Convex (34b) The corresponding distance in the normal section (see Fig. 11) is p3. Then: _ V\\ cos \\pi, When mo is less than 2: p \\ I .. \" ' 0 \" ' 0 ZirlF \u00bb'o / L 2 2?/i,, 5mp When m0 is greater than 2: / m , \\ ! r ?\u00ab0 2 m.F \u00bbic P3 = ?2 I I \u00b1 \\ \u00bb'o / L 2??!p 5mp (35) ]Conc?ve (35 1Convex a Concave Convex (35b) where: cos 4 (cos2$ + tan2q5) Equations (35a) and (35b) are the same as those given in equation (2). APPENDIX F Layout Procedure The following data are necessary for the layout (Fig. 13) : tp, tG = Mean normal circular thickness of pinion and gear (tp + tG = pn) ap aG = Mean addendum of pinion and gear bp, bG = Mean dedendum of pinion and gear TTP, TTG = Tool edge radius of pinion and gear cutters RNP, RNG = Equivalent pitch radius in the normal plane 2 cos 7 cos21/- /" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002035_j.measurement.2020.108224-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002035_j.measurement.2020.108224-Figure7-1.png", "caption": "Fig. 7. Centrifugal fan and naming of different part of blade root.", "texts": [ " The centrifugal fan selected is mainly used in the ventilation and cooling system of high-speed trains and nameplate parameters are shown in Table 5. The centrifugal fan consists of a volute and an impeller as shown in Figs. 6 and 7. The impeller consists of a front disc, 60 blades, a hub and a rear disc. The hub and the rear disc are connected by bolts. The front disc is pressformed with blade by a bending roller of the clasps, the same as the rear disc. The definition of buckle teeth on blade is shown in Fig. 7. This connection mode is simple and has good stiffness. Then the finite element model of centrifugal fan is established. The elements type is SOLID186 and elements size is 0.4\u20135 mm. A total of 1,102,276 hexahedral grids are connected with 1,536,162 nodes. There are 16,674 grid elements for a single blade and 60 blades with a total of 1,000,440 grids. At the same time, mesh refinement was carried out at the blade root to obtain higher precision stress. The finite element model and material parameters, as in Table 4, are imported into ANSYS Workbench software to calculate the strain of the fan at rated speed" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002380_s0368393100069789-Figure14-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002380_s0368393100069789-Figure14-1.png", "caption": "FIGURE 14. General arrangement of projected aircraft. For main dimensions\u2014see Table III.", "texts": [ "UC6 atB Vile puvweas '/3?! PLY e o v e a t M Q WEIGHT.Ib COM PLC T 6 . 80 variation in chord and span on wing weight is shown in Fig. 11. No allowance has been made for the provision of control surfaces; indeed their form and size has not yet been established to enable this. Likewise, no careful consideration has been made of the other items of structure weight but tentative estimates of these are given in Table II, referring to the projected layout envisaged in making the performance calculations later (see Fig. 14). It has been considered that all the surfaces should be rigid to preserve a low-drag form; owing to its relatively small size, the tail-plane would probably have to be shaped from solid birch; the propeller blades are also supposed solid. Details of the other items will emerge from the following paragraphs. 5. General Layout The fuselage layout presents a considerable prob lem; we shall suppose that it is necessary to accom modate two men in tandem in the cycling position, two men being the least required\u2014one to ensure power continuity while the other is temporarily abstracted by controlling the aircraft\u2014and three or more men seeming somehow to invite derision\u2014although it may well be true that an \"eight\" is the optimum in flying as in rowing", "-B1 BFHCIENCX ADES j I shown by Fig. 12. We shall see later that the high propeller position is troublesome to the longitudinal stability and control of the aircraft, but it appears to be an essential feature, and indeed appeared on all the man-powered aircraft designs of the pre-war era. The length of the fin extension above the fuselage should be kept as short as possible, the critical interest lying in providing just an adequate tail arm rather than a large one (see Sections 9 and 10). The general arrange ment shown in Fig. 14 has been assumed for initial assessment and its stability characteristics are discussed in later Sections, where certain important changes are suggested. A summary of the leading geometrical features of this layout appears in Table III. It involves an optimised wing span and chord based on an examination of the effect of these variables on the minimum power needed to maintain level flight at an altitude of 35 ft. The 85 8Ca B7 86 EFFICIEMCV \u00b0/O 8 3 SO FIGURE 13. Variation of propeller efficiency with blade diameter", " To determine the heights and durations possible, it has been assumed that the air crew are of average size and develop 90 per cent of the National record level (0-49 h.p. plus 13,000 ft. lb.), that 4 per cent of this is lost in the transmission to the propeller, and that the efficiency of the latter is 87 per cent. In the ground run the thrust supplied by the wheels is reduced by a fraction 0006 of the ground reaction, to account for rolling and mechanical friction(3). The assumed profile drag breakdown of the aircraft shown in Fig. 14 is given in Table IV and speed and height variations are shown in Fig. 16. It will be appreciated that items \" interference \" and \" leaks and gaps\" can only be guessed at this stage. Control surfaces are likely to provide the bulk of the drag under the latter heading. With the foregoing assumptions it can be deduced that in the flight path for maximum gain of height, some 15 seconds of maximum power are available after take off during which time the aircraft reaches a height of 22 ft. and the speed drops from 55 to 50 ft", " The stick-free stability has not yet been studied but the spiral damping could easily be improved in this condition (if considered necessary) by allowing the rudder to trail in sideslip113'\u2014so reducing n, still further. It should be noted that the foregoing estimates (as well as those of the next paragraph) have included only a token allowance for propeller-fin effect based on a guessed value of yv=i\u20140-02. Its effect on nv (and the spiral motion) is relatively important, but a more accurate assessment could only be made when its design is fixed. 10. Longitudinal Stability The estimated stick-fixed logitudinal stability deriva tives of the layout given in Fig. 14 are quoted in Table VI relative to flight at 6 ft. height and CL = 0-94 (corresponding to a speed of 50 ft./sec). The aircraft has been designed to provide a static margin of 005c without propeller-fin effect (and 0 08c with). The value of mu is zero for gliding flight, and on the debatable assumption that constant power is maintained in any disturbed motion, equals 0 009 in the cruise condition, and 0022 at full power (1.25 h.p.). If merely r.p.m. is kept constant, the value of mu would be even higher" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001048_0142331219883056-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001048_0142331219883056-Figure1-1.png", "caption": "Figure 1. The schematic diagram of the considered double-rod hydraulic system.", "texts": [ " Section 3 gives the time-varying BLF and dual ESO-based backstepping controller design procedure as well as corresponding theoretical results. Contrastive simulation results can be obtained in Section 4. The conclusions as well as future prospects can finally be found in Section 5. Nonlinear model of the considered hydraulic system The double-rod hydraulic servo system under consideration is the same as that in Yang et al. (2016). The schematic of the double-rod cylinder hydraulic control system is shown in Figure 1. The goal is to have the inertia load to track any smooth motion trajectory as closely as possible and constrain the output at the same time. The dynamics of the inertia load can be described by m\u20acy=PLA B _y+ d(t, y, _y) \u00f01\u00de where d(t, y, _y) represents lumped uncertain nonlinearities due to unmodeled friction forces, external disturbances, and other hard-to-model terms. The cylinder dynamics can be modeled as (Yang et al., 2019) V1 be _P1 = A _y CtPL +Q1 Qe(t) V2 be _P2 =A _y+CtPL Q2 +Qe(t) \u00f02\u00de where Qe(t) represents the time-varying modelling error due to unmodeled pressure dynamics" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001398_s12540-020-00779-6-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001398_s12540-020-00779-6-Figure1-1.png", "caption": "Fig. 1 a Schematic of the steps of the pre-cold welding procedure, b pre-cold chamber of sample", "texts": [ " Heat treatment was performed to increase the weldability of the specimens and to remove the liquation-sensitive phases during welding according to the GTD-111 solidus temperature (1295 \u00b0C) in accordance with Table\u00a02. For this purpose, an electric tube furnace with argon gas shielding (with a purity of 99.999%) was used. By microstructural investigations of heat treated specimens, the optimum specimen was selected to investigate the influence of precold and pre-heat treatment on it. To pre-cool the samples, a sealed chamber with an approximate temperature of \u2212 25 \u00b0C was used in accordance with Fig.\u00a01a. The frame of this chamber was made of plastic material (due to no heat loss) and the surface was made of copper (to increase thermal conductivity). Dry ice was in direct contact with the underside of the work-piece during welding and the heat transfer was continuous (Fig.\u00a01b). In order to investigate the effect of pre-heating on the welding behavior of the alloy, an electric heater was used below the sample. Preheating temperature of the specimens during welding was measured by a K-type thermocouple with a tolerance of \u00b110 \u00b0C. Table\u00a03 shows the pre-heat and pre-cold conditions of the samples. The welding was performed by a Nd:YAG pulse laser machine, model IQL-10 with a maximum power of 400 W. Selected parameters were selected based on previous studies by Taheri et" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000539_s00170-016-9337-2-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000539_s00170-016-9337-2-Figure1-1.png", "caption": "Fig. 1 Tool path and cut field", "texts": [ " Tool-axis A :R\u2192R2 is a two-dimensional unit sphere; it denotes the orientations of the cutter at the corresponding CC points [28]. When parameterized by t ( t can be any parameter, and except for in the CNC controllers, t is usually not time parameter), the tool path can be represented as F t\u00f0 \u00de \u00bc P t\u00f0 \u00de;A t\u00f0 \u00de\u00f0 \u00deT \u00f01\u00de The cut vector Q(t) =dP/dt is the tangent vector of the 3D surface S(u, v) with parameters (u, v). The union of cut vectors forms the cut vector field Q, which is the tangent vector space of S(u, v), shown in Fig. 1. In order to improve the machining quality, the tool path should not cross with each other or be self-intersected. In order tominimize the machining time, the tool path length should be minimized and the tool path should be smooth. So it is ideal that the tool paths are evenly distributed on the machining area, and the values of step-overs tend to be uniform [11]. Figure 2 shows examples of poor and ideal tool path shapes. We assume that the trajectories of the tool path never cross each other (an optimized tool path should be like that) and each trajectory is either a closed loop or intersects the boundary of the machining region even times" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000471_j.ijsolstr.2016.07.004-Figure19-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000471_j.ijsolstr.2016.07.004-Figure19-1.png", "caption": "Fig. 19. An annular membrane under torsion ( Akita et al., 2007 ).", "texts": [ " 18 (a) shows that when the thickness increases, the neutral xis goes downward with the maximum tensile stress diminishing. ig. 18 (b) illustrates that the wrinkling strain decreases when the hickness of the membrane is augmented. It is possible to infer hat the longitudinal normal stress will be linearly distributed f the thickness increases to the extent that the wrinkling strain isappears. .2. An annular membrane under torsion An annular membrane under torsion studied by Akita et al. 2007) (see Fig. 19 ) is reproduced here to verify the accuracy of he new wrinkling model and to further study the influences of he prestress, Poisson\u2019s ratio, Young\u2019s modulus and thickness on he wrinkling behaviors of a membrane. The annular membrane shown in Fig. 19 is attached to a rigid ub along the inner edge and to a fixed guard ring along the uter edge. Its thickness, inner and outer radii are t, a and b , espectively. The membrane is pretensioned by a uniform radial tress \u03c3 0 and then subjected to a twisting moment M at the rigid ub. When the hub is rotated by an angle \u03d5, wrinkles will appear round the hub with a radius R . According to the analytical solution presented by Mikulas 1964) to this problem, the relation between the radius of wrinles R and the twisting moment M can be determined by the C C \u03d5 w \u03d5 6 t t t r a m e D r following equations: 1 A + 1 B \u2212 ln ( B A ) \u2212 2 3 = 0 (63) 2 1 \u2212 [ 1 + 2 C 1 ( a b )2 ]2 R\u0304 4 + M\u0304 2 = 0 (64) 2 = { R\u0304 + [ 1 R\u0304 + 2 \u0304R ( a b )2 ] C 1 }2 + ( M\u0304 R\u0304 )2 (65) where M\u0304 = M 2 \u03c0a 2 \u03c30 t , R\u0304 = R a , A = C 2 M\u0304 2 \u2212 1 , B = R\u0304 2 C 2 M\u0304 2 \u2212 1 (66) It is clear that Eqs", " (67) \u2013( 69 ) as \u03c31 \u03c30 = \u23a7 \u23a8 \u23a9 C 2 / r \u221a C 2 \u2212M 2 / r 2 r < R 2 C 1 a 2 / b 2 + 1 + \u221a C 2 1 + M 2 2 r > = R (67) r \u03c32 \u03c30 = { 0 r < R 2 C 1 a 2 / b 2 + 1 \u2212 \u221a C 2 1 + M 2 r 2 r > = R (68) \u00af = 3 M\u0304 8 ( 1 \u2212 a 2 b 2 )[ 1 R\u0304 2 \u2212 1 B + ln ( B A ) + 1 R\u0304 2 \u2212 8 3 ( a b )2 + 5 3 ] (69) here \u00af = 3 E\u03d5 4 \u03c30 ( 1 \u2212 a 2 b 2 ) (70) .2.1. Verification of accuracy In this sub-section, the annular membrane is analyzed with he new wrinkling model and the results are compared with he theoretical solutions obtained from Eqs. (63) to ( 65 ), ( 67 ) o (69). The finite element model is shown in Fig. 20 . Its outer adius b and inner radius a (shown in Fig. 19 ) are 1250 mm nd 500 mm, respectively. The thickness t is 1 mm. The Young\u2019s odulus E = 10 0 0 MPa and the Poisson\u2019s ratio \u03c5 =1/3. The inner dge is restrained in the radial direction and the circumferential OFs of nodes on the edge are coupled to have the same torsional otation. The outer edge is firstly pretensioned to introduce a u c t r r a m p M i a s c b p r o a p a w fl v t 6 m ( 0 8 As for the numerical examples in Section 6.2 , their geometrical and physical parameters are the same as given in Section 6", " The horizontal axis in the figure represents the ratio f the radial coordinate r to the inner radius a and the vertical xis denotes the major or minor principal stresse ( \u03c3 1 or \u03c3 2 ) to the restress \u03c3 0 . Figs. 21 and 25 indicate that the wrinkling model is ccurate and the results agree well with the theoretical solutions. Fig. 21 also manifests that the major principal stress in the inkled region is much larger and its distribution curve becomes at in the taut area. Conversely, the minor principal stress almost anishes in the wrinkled region while it increases gradually with he radial coordinate in the taut zone. .2.2. Influence of the prestress The annular membrane shown in Fig. 19 subjected to a twisting oment M = 10,0 0 0 N \u2022mm is analyzed 8 with different prestresses \u03c3 = 0 . 1 \u03c4, 0 . 2 \u03c4, 0 . 3 \u03c4, 0 . 4 \u03c4 and 0.5 \u03c4 , where \u03c4 is the uniformly o u m \u223c s i r distributed shear stress along the circumferential direction at r = a and obtained from the boundary condition with \u03c4 = M 2 \u03c0a 2 t ). The influences of the prestress on the major and minor principal stresses and the wrinkling strain are illustrated in Fig. 26 (a) and (b), respectively. The vertical axis in Fig. 26 (a) represents the ratio of the principal stress \u03c3 1 or \u03c3 2 to the shear stress \u03c4 and the vertical axis in Fig", " 26 (a) represent the major principal stresses nder different prestresses while the discrete symbols denote the inor stresses. Fig. 26 (a) shows that when the prestress increases from 0.1 \u03c4 0.5 \u03c4 , the major and minor principal stresses are augmented lightly in the taut area. Fig. 26 (b) indicates that the prestress s able to decrease the wrinkling strain notably in the wrinkled egion. 6 r i m s p i 6 t t s Y I E m f d d a m o a i 7 p k d m t m a e a r b A f a g P R A A A A B D D D E E F H H H .2.3. Influence of the poisson\u2019s ratio The annular membrane (shown in Fig. 19 ) with the Poisson\u2019s atio varying from 0.1 to 0.5 is studied and the results are plotted n Fig. 27 . The solid or dashed lines in Fig. 27 (a) represent the ajor principal stress while the discrete symbols denote the minor tresses. The Poisson\u2019s ratio is noted to have little effect on the rincipal stresse distributions along the radial direction. But it nfluences the wrinkling strain as revealed in Fig. 27 (b). .2.4. Influence of the bending stiffness The annular membrane with different Young\u2019s modulus or hickness are analyzed in this Subsection to study the effect of he bending stiffness on the stress distribution and wrinkling train" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000294_ijvp.2019.097098-Figure9-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000294_ijvp.2019.097098-Figure9-1.png", "caption": "Figure 9 Automatic mesh refinement in footprint region (see online version for colours)", "texts": [ " Thus, the LEPP algorithms proceeds by exploring the neighbouring triangle(s) and checking if the longestedge rule can be applied. In Figure 8(c) one can see that this propagates the edge splitting to outer triangles, but the final result is well balanced and the LOD gradually increases from the outer areas to the initial t0 triangle. The LEPP algorithm can run in real time without much overhead, as it increases the LOD only for the triangles that are marked as \u2018touching\u2019 and have \u03c3 > 0, as shown in Figure 9. The only requirement from the user side is a LOD threshold value, i.e., the maximum allowed length of edges in the contact area. Although not currently implemented, a similar strategy could be used to coarsen the mesh back to the original LOD as the vehicle moves away from a previously refined area. We implement an heuristic algorithm for reproducing the effect of soil build-up at the sides of the rolling tyre, inspired to what proposed in Krenn et al. (2008b). This is especially useful in the case of soft sandy soils, where it creates more realistic ruts" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000028_s40192-019-00145-4-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000028_s40192-019-00145-4-Figure3-1.png", "caption": "Fig. 3 Schematic of the substrate with all ten tracks [21] with the dimensions in millimeters", "texts": [ " (4) This process of mapping the melt pool at each time step and the growth and solidification of CA cells is repeated for the entire length of each track. The AM-Bench experimental work was performed by NIST. The experiment was designed to isolate the effect of individual tracks printed on a bare plate of alloy IN625. The plate dimension is 24.82 mm \u00d7 24.08 mm \u00d7 3.18 mm [21]. The tracks were printed at the center of the plate and were separated by a distance of 0.5 mm. The plate geometry and the track positions are shown in Fig. 3. To avoid heat buildup, each track was printed at 5-min intervals. The nominal temperature of the base plate was 25 \u00b0C at the start of experiment. Three different process parameter sets were used to produce the tracks as shown in Table 2. The laser was a continuous-wave ytterbium fiber (Yb-fiber) laser with a wavelength of 1070 nm. The experimental work was performed using twomachines, namely Additive Manufacturing Metrology Testbed (AMMT) and Commercial Built Machine (CBM) for which the laser spot sizes were 100 \u03bcm and 59 \u03bcm, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001738_j.mechmachtheory.2019.103629-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001738_j.mechmachtheory.2019.103629-Figure7-1.png", "caption": "Fig. 7. Trajectory for validation of the inverse dynamic models.", "texts": [ " (44) into the latter equation, yields F a = \u2212 \u02dc F c \u2212 J \u2212T ( F E + 4 \u2211 i =1 J T i, j F i, j + 4 \u2211 i =1 J T si, j F si, j ) (50) where F a = [ f A 1 , 1 f A 1 , 2 f A 2 , 1 f A 2 , 2 ]T is the vector of actuation forces and \u02dc F c is: \u02dc F c = \u23a1 \u23a2 \u23a2 \u23a3 m c1 , 1 (g \u00b7 \u02c6 i \u2212 p\u0308 1 , 1 ) m c1 , 2 (g \u00b7 \u02c6 i \u2212 p\u0308 1 , 2 ) m c2 , 1 (g \u00b7 \u02c6 i \u2212 p\u0308 2 , 1 ) m c2 , 2 (g \u00b7 \u02c6 i \u2212 p\u0308 2 , 2 ) \u23a4 \u23a5 \u23a5 \u23a6 (51) Eq. (50) represent the dynamic model of the mechanism of interest. In this section, a simulation study is performed to verify the correctness of the presented model. To do so, for a certain trajectory, the computed actuator forces using the proposed analytical model, are compared with those obtained by using MATLAB Simscape Multibody. For the simulation, the motion below which is shown in Fig. 7 is given to the reference point of the end-effector: \u23a7 \u23aa \u23a8 \u23aa \u23a9 x = 80 cos ( \u03c0 80 t) y = 1 . 4 t sin ( \u03c0 3 t) z = \u22125 + 0 . 5 t cos ( \u03c0 3 t) \u03c6 = \u03c0 4 sin ( \u03c0 20 t) (52) The required joint space trajectory was calculated using the inverse kinematic model, then the obtained results were used to calculate the required forces through the inverse dynamics. The computed forces using the Virtual Work model and Simscape Multibody are shown in Fig. 8 . The main blocks of the Simscape Multibody model are also shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001872_icmre49073.2020.9065189-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001872_icmre49073.2020.9065189-Figure2-1.png", "caption": "Fig. 2: Shock absorber installed on IRB360 Delta robot, top view, bottom view, suspension.", "texts": [ " To overcome these issues, we employ a custom made shock absorber mechanism attached at the end of the robotic arm, in order to gradually press the suction cup on the object and facilitate sealing. The shock absorber has a triangular shape with dimensions of 70mm for each side and 30mm of height, weighting 700gr. It consists of two parallel plates of plywood with thickens of 4mm that are linked together by three aluminum rods, of 110mm length, able to move on linear bearings with pretension spring, one on each corner, giving it a 50mm travel capability (Fig.2). The system is able to move freely on the z axis, keeping the two plates always parallel to each other. The top plate has been made compatible with the end effector of ABB IRB360 DELTA robot, making additionally sure that there is enough space for the axial rods of the linear bearing to not collide with the robot. The bottom plate is matching in dimensions to the top plate, being also partially adapted to house the suction cups. The two plates have a spacing of 100mm, thus providing enough room for the installation of the suction tubing" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000976_s11044-019-09704-1-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000976_s11044-019-09704-1-Figure4-1.png", "caption": "Fig. 4 Accurate ring raceways of thin-walled four-point contact ball bearing", "texts": [ " (2), the subscript \u2217 denotes the inner ring \u2018i\u2019 or outer ring \u2018o\u2019, R\u2032 i = Di/2 + ri, R\u2032 o = Do/2 \u2212 ro, R\u2032 i and R\u2032 o denote the center of the circle of ring raceway section, known as the \u201cmajor radius\u201d of ring raceway torus, ri and ro denote the radius of curvature of ring raceway, known as the \u201cminor radius\u201d of ring raceway torus; 0 \u2264 \u03d5\u2217 \u2264 2\u03c0 , where \u03d5\u2217 is the revolution angle of the circle around the axial direction of the bearing; \u03b31 \u2264 \u03b3\u2217 \u2264 \u03b32, where \u03b3\u2217 is the angle of partial ring raceway torus surface, starting at \u03b31 and ending at \u03b32; \u03bei = \u2212(ri \u2212d/2) sin\u03b1i and \u03beo = (ro \u2212d/2) sin\u03b1o denote the offset distances of ring raceway torus along the axial direction of the bearing. According to Eq. (2), the accurate ring raceway of the bearing in the ring body reference is established by using user\u2019s subroutines in ADAMS software package, shown in Fig. 4. Taking INA CSXC045 as an example, the multibody contact dynamics of the bearing is investigated in the proposed work. The parameterized three-dimensional modeling with the crown-type cage is performed using user\u2019s subroutine in ADAMS. From Eq. (1), the accurate ring raceway is implemented as a partial geometric surface of the torus using ADAMS\u2019s dummy part. The characteristics and modeling of the bearing are presented in Table 1 and Figs. 5\u20136. There\u2019re 29 pockets on the crown-type cage: 14 small pockets and 15 big pockets, respectively", " The developed approach is free of these disadvantages and is summarized as follows: (see Flow chart of Fig. 14) (1) Based on ADAMS, we apply users\u2019 subroutines of a microprogram to establish the parametric design of inner ring, outer ring, ball, and cage of the bearing. (2) The parametrized geometry models are achieved using the input parameters of the bearing shown in Figs. 5 and 6. Using equations of the surfaces of the torus, we may represent analytically the volume of the designed ring raceways shown in Fig. 4. (3) The definition of contacting surfaces for the contact algorithm of ADAMS computer programs is automatic using the parameterized model as well, and requires definition of the action and target surfaces. We are using the equations of contact stiffness to calculate the load\u2013deflection factor of contact points. Also we apply the sphere-to-partial ring torus surface contacts, sphere-to-plane contacts, sphere-to-cylindrical contacts, and cylindrical-to-cylindrical contacts to define for sphere-to-race contacts (point contact), sphere-to-pocket contacts (point contact), and cage-to-outer guidance contacts (line contact) in the multibody contact models" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000351_s00034-015-0229-8-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000351_s00034-015-0229-8-Figure1-1.png", "caption": "Fig. 1 Quadrotor motion obtained by varying the speed of the four rotors", "texts": [ " Roll motion is produced when the speed of the left rotor is different from that of the right rotor. Pitch motion is produced when the speed of the front rotor is different from that of the rear rotor. Yaw motion is produced when the speed of the front and rear rotors is different from that of the left and right rotors. The front and rear rotors rotate counterclockwise, whereas the left and right rotors rotate clockwise. Thus, anti-torque can be offset because of the symmetry of the rotors, as shown in Fig. 1. Modeling a quadrotor is not an easy work because of the shaft coupling and serious nonlinearity. In this paper, in order to simplify the modeling without loss of generality, some assumptions are given as follows: Assumption 1 (1) The structure is rigid and strictly symmetrical. (2) The body fixed frame origin coincides with the center of mass. (3) DC voltage and torque have a proportional relationship. (4) Gyroscopic effect can be ignored when flying at low speed. Define \u03b7 = ( \u03c6 \u03b8 \u03c8 ) \u2208 R3 as roll, pitch and yaw angles respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002317_mmsp48831.2020.9287100-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002317_mmsp48831.2020.9287100-Figure4-1.png", "caption": "Fig. 4. (a) Niryo One robot arm in Gazebo simulator (b) Niryo One robot arm in RViz with transform frames", "texts": [ " Ji = (1 \u2212 \u03bbi)Ji+1 + \u03bbiJi () Ji+1 = (1 \u2212 \u03bbi)Ji + \u03bbiJi+1 () According to [32], the algorithm converges in fewer iterations, lower computational cost and generates more realistic and smoother poses. Constraints can also be easily integrated within FABRIK. B. System Architecture: Server-Side ROS is a versatile, multilingual middleware. It provides a collection of tools and libraries for robotic developers to simplify their effort of designing complicated and robust robot actions [33]. ROS is open-source and is language agnostic. Gazebo, as shown in Fig. 4(a), serves as a robotic simulator while ROS is the interface to the robot. The key benefit of using Gazebo is its ready integration with ROS. Similar to ROS, Gazebo is also open-source. It is one of the most popular robotic simulators [22]. RViz, as shown in Fig. 4(b), is not a simulator but a 3d visualization tool for ROS applications. It offers realtime sensor information from robot sensors and a view of the Authorized licensed use limited to: University of Prince Edward Island. Downloaded on May 27,2021 at 13:32:07 UTC from IEEE Xplore. Restrictions apply. robot model. It can be used to display 2D and 3D point clouds as well as camera or lasers data [34]. A node (as per Fig. 5) in ROS represents a separate program running a specific task. It can be written in any programming language, although it is primarily developed in C++ and Python [35]. Each standalone task can be divided into different nodes which communicate with each other through unique named buses. These so-called unique named buses are described as topics in [35]. A publisher can broadcast messages into a ROS topic, whilst a subscriber listens and retrieves the messages from a topic. In order to successfully manipulate the robot arm in Gazebo as shown in Fig. 4(a), the first terminal will launch 6 nodes, namely: the file server; Gazebo simulator server; Gazebo simulator GUI client; ROS API libraries; ROSBridge WebSocket; and finally, the Gazebo URDF model node. The URDF model node will spawn the Niryo One robot arm model in the Gazebo simulator. The controller manager node and the robot state publisher node are launched. The ROS control framework is defined inside a YAML [36] configuration file. It is used to implement and manage the simulated gazebo robot controllers" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001538_jestpe.2020.3029802-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001538_jestpe.2020.3029802-Figure1-1.png", "caption": "Fig. 1. Structure diagram of EDS maglev. (a) Overall structure diagram of EDS maglev. (b) Local structure diagram of suspension and guidance system.", "texts": [ " Based on the space harmonic method, combined with the equivalent circuit of suspension and guidance loop, this paper analyzes the characteristics of the suspension and guidance system of EDS maglev by using a new M.M.F distribution model. Finally, the analytical results are compared with the three-dimensional finite element method (3-D FEM) simulation results and the experimental data of the Yamanashi test line to verify the effectiveness and accuracy of the method. II. OPERATION PRINCIPLE OF SUSPENSION AND GUIDANCE SYSTEM OF EDS MAGLEV EDS maglev consists of suspension, guidance and driving system. As shown in Fig.1(a), the primary winding combined with the vehicle coils forms the linear synchronous motor to drive the maglev. And the vehicle coil interacts with the figure-eight-shaped coil to form the suspension and guidance system. The magnetic field level is reduced to 0.5mT by adopting various magnetic shielding methods in maglev [22]. As shown in Fig.1(b), the suspension system consists of ground figure-eight-shaped coils and vehicle coils, and the guidance system consists of ground figure-eight-shaped coils, vehicle coils and zero-flux cables connecting two groups of ground figure-eight-shaped coils on the opposite side [15]. The figure-eight-shaped coil is composed of upper and lower unit coils connected in reverse. This special structure leads to the phenomenon that when there is current flowing in the figure-eight-shaped coil, the direction of the magnetic flux of upper and lower unit coils is opposite" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000976_s11044-019-09704-1-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000976_s11044-019-09704-1-Figure5-1.png", "caption": "Fig. 5 Three-dimensional model of crown type cage", "texts": [], "surrounding_texts": [ "Figure 1 shows a thin-walled four-point contact ball bearing with crown-type cage applied to harmonic gear drive in industrial robots. A thin-walled four-point contact ball bearing has four contact points of ball-to-ring raceway surface contact as theoretically shown in Fig. 2. There\u2019re four ring raceways for four-contact points of ball-to-ring raceway surface contact: two outer-ring raceways and two inner-ring raceways, respectively. The section consisting of four ring raceways is similar to an ellipse; d , Di and Do are the diameters of balls and inner and outer raceway, respectively; ri and ro are radii of curvature of the inner and outer ring raceway, respectively; \u03b1i and \u03b1o are nominal contact angles of the inner and outer ring raceway, respectively; gi and go are the distances to the center of curvature of the inner and outer ring raceway, respectively; hi and ho are the distances of ball to vertex of the inner and outer ring raceway, respectively. Actually, a thin-walled four-point contact ball bearing is a typical multibody system consisting of balls, inner ring, outer ring, and crown-type cage. The radial, axial, or combined radial and axial loadings, and especially overturning moment loadings, are considered in most bearing applications. The internal contact interactions of ball-to-ring raceway, ball-to-cage pocket and cage-to-ring guidance are complicated and time-varying. The load distribution of all components, dynamic performance, contact stresses, and fatigue endurance of the bearing are changing significantly. The assumptions of three-dimensional multibody contact dynamics simulation analysis of thin-walled four-point contact ball bearing are presented as follows: (1) The inner ring is fixed rigidly to the ground and the outer ring is rotated around the axial direction Z axis. The six degrees of freedom of balls, cage and outer ring are considered. (2) Without structural elastic deformation, the rings, balls and cage are ideal rigid bodies, only the nonlinear Hertz contact deformations are considered for ball-to-ring raceway surface contacts, ball-to-cage pocket surface contacts, and cage-to-ring guidance surface contacts. (3) To simplify the complexity, the normal contact force and the Coulomb friction force of ball-to-ring raceway, ball-to-cage pocket, and cage-to-ring guidance are considered. The influences of elasto-hydrodynamic lubrication (EHL) and temperature on the bearing are neglected. The accuracy of ring raceway of thin-walled four-point contact ball bearing is very important to calculate the contact forces of ball-to-ring raceway surface contacts correctly. Without consideration of structural elastic deformation of ring raceway, the rigid surface of ring raceway is part of the torus surface. It\u2019s a surface of revolution generated by revolving a circle in three-dimensional space about coplanar axis as shown in Fig. 3. By ooxoyozo we mark the coordinates at the center of the outer ring, ootcxotcyotczotc mark the coordinates at the center of curvature of the outer ring raceway, sor is the location of the curvature center of the outer ring raceway going from the center of the outer ring, sot is the location of point k going from the center marker of the curvature of the outer ring raceway, while sok is the location of point k going from the center marker of the outer ring. The position vector of point k on the surface of partial torus of the outer ring raceway can be expressed as sok = sor + sot (1) where sor and sot are constant for a given ball bearing, sor = Do/2 \u2212 ro, sot = ro. The equation of the ring raceway in the ring\u2019s body reference frame is obtained as S \u239b \u239c\u239d x \u2032\u2217 y \u2032\u2217 z\u2032\u2217 \u239e \u239f\u23a0 = \u239b \u239c\u239d (R\u2032\u2217 + R\u2217 cos\u03b3\u2217) cos\u03d5\u2217 (R\u2032\u2217 + R\u2217 cos\u03b3\u2217) sin\u03d5\u2217 R\u2217 sin\u03b3\u2217 + \u03be\u2217 \u239e \u239f\u23a0 . (2) In Eq. (2), the subscript \u2217 denotes the inner ring \u2018i\u2019 or outer ring \u2018o\u2019, R\u2032 i = Di/2 + ri, R\u2032 o = Do/2 \u2212 ro, R\u2032 i and R\u2032 o denote the center of the circle of ring raceway section, known as the \u201cmajor radius\u201d of ring raceway torus, ri and ro denote the radius of curvature of ring raceway, known as the \u201cminor radius\u201d of ring raceway torus; 0 \u2264 \u03d5\u2217 \u2264 2\u03c0 , where \u03d5\u2217 is the revolution angle of the circle around the axial direction of the bearing; \u03b31 \u2264 \u03b3\u2217 \u2264 \u03b32, where \u03b3\u2217 is the angle of partial ring raceway torus surface, starting at \u03b31 and ending at \u03b32; \u03bei = \u2212(ri \u2212d/2) sin\u03b1i and \u03beo = (ro \u2212d/2) sin\u03b1o denote the offset distances of ring raceway torus along the axial direction of the bearing. According to Eq. (2), the accurate ring raceway of the bearing in the ring body reference is established by using user\u2019s subroutines in ADAMS software package, shown in Fig. 4. Taking INA CSXC045 as an example, the multibody contact dynamics of the bearing is investigated in the proposed work. The parameterized three-dimensional modeling with the crown-type cage is performed using user\u2019s subroutine in ADAMS. From Eq. (1), the accurate ring raceway is implemented as a partial geometric surface of the torus using ADAMS\u2019s dummy part. The characteristics and modeling of the bearing are presented in Table 1 and Figs. 5\u20136. There\u2019re 29 pockets on the crown-type cage: 14 small pockets and 15 big pockets, respectively. The materials of ring and cage are GCr15 and brass, the densities are 7800 and 8500 kg/m3, respectively." ] }, { "image_filename": "designv11_5_0000347_2816820-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000347_2816820-Figure6-1.png", "caption": "Fig. 6. Example scenario.", "texts": [ " Let us consider \u03ba number of sensor nodes deployed after assigning each node a unique ID such that id = {1 . . . \u03ba}. Due to the occurrence of adverse environmental conditions, if a sensor node gets isolated from the network, it starts to behave as the BEGIN node, and initiates connectivity reestablishment. A BEGIN node reestablishes connectivity with a STOP node, which is the nearest activated neighbor node of the BEGIN node, if it was connected before with the BEGIN node but has ceased to be connected. Figure 6 shows how the BEGIN node, U , reestablishes connectivity with the STOP node, V . Initially, the BEGIN node functions as the FORWARD node, which activates its neighbor node and coalesces them to participate in the connectivity reestablishment process. Gradually, the activated neighbor nodes take turns acting as the FORWARD node and progress with the connectivity reestablishment process until the STOP node becomes a FORWARD node. The FORWARD node broadcasts the node ACTIVATION message within its single-hop neighbor to activate the neighbor nodes in the sleep state", " Nodes receiving a REQUEST packet remain active for the next trep time and calculate their own benefit and cumulative benefit values. The parameter trep is the estimated time for receiving a REPLY packet from the STOP node. Thereafter, each of the nodes receiving a REQUEST packet functions as the FORWARD node and continues the same process until the STOP node is reached. Based on the highest cumulative benefit value, the nodes receiving REQUEST packets from multiple FORWARD nodes select one of the FORWARD nodes as the downstream node to the BEGIN node. Figure 6 shows that Node 8 sends a REQUEST packet to the upstream Node 9 after increasing its communication range (indicated as a dotted line), because there is no node to reach the STOP node within its reduced communication range. Similarly, Nodes 16 and 18 also increase their communication range to get a neighbor node. Upon receiving REQUEST messages from multiple FORWARD nodes, the STOP node selects one as the downstream node to the BEGIN node and sends a REPLY packet to it. The REPLY packet contains the STOP node ID, STP_ID, STOP node address, STP_ADDR, BEGIN node ID, BGN_ID, and BEGIN node address, BGN_ADDR, as shown in Figure 5(b)" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000437_s12206-015-1235-2-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000437_s12206-015-1235-2-Figure1-1.png", "caption": "Fig. 1. A pair of meshing helical gears [24].", "texts": [ " The main purpose of this work was to investigate the effects of work conditions and design parameters on lubrication performances for helical gear pair. It is known generally that the helical gear lubrication is a transient process. However, the difference between a transient solution and a quasi-steady solution is insignificant for the meshing process of a helical gear pair [9]. So, the transient squeezing effect is not considered in this work. Besides, the tooth modification is not taken into account. The meshing process of a helical gear pair are shown in Fig. 1. According to the principle of gear meshing, the meshing of helical gears can be transformed approximately as two cones with opposite directions in each meshing position. As shown in Fig. 1, the half cone angle is the helix angel bb and the helical gear parameters of the selected by Ref. [24] are applied here which is listed in Table 1. The detailed information about the geometrical and kinematic analysis could be found in Appendix A. For helical gear lubrication, the non-Newtonian fluid effect cannot be neglected. Especially, the effect of rheological properties of fluid on friction coefficient is remarkable. In last decades, the researchers proposed several constitutive equations, among which the Eyring model is commonly applied in EHL studies [25]", "Wen, A generalized reynolds equation for non- newtonian thermal elastohydrodynamic lubrication, ASME Journal of Tribology, 112 (4) (1990) 631-636. [28] C. H. Venner, Multigrid solution of the ehl line and point contact problems, Ph.D. Thesis, University of Twente, Netherlands (1991). [29] J. I. Pedrero, M. Pleguezuelos and M. Mu\u00f1oz, Critical stress and load conditions for pitting calculations of involute spur and helical gear teeth, Mechanism and Machine Theory, 46 (4) (2011) 425-437. Appendix As shown in Fig. 1, the meshing of helical gears can be transformed approximately as two cones with opposite directions. Each cone is studied separately as shown in Fig. A.1. The radius of the end side of the two cones read 1 1 1 1 1 2 2 1 2 2 tan tan . d b d b r N K r r N K r q q \u00ec = =\u00ef \u00ed = =\u00ef\u00ee (A.1) where \u03b81, \u03b82 are defined in Fig. 1. When the end side point K1 locates at the pitch point P, we have \u03b81 = \u03b82 = \u03b1n, where \u03b1n is the pressure angle; rb1, rb2 are the radius of base circle. The rotational radius of the contact point K for the two cones read 1 1 2 2 sin sin . d b d b r r y r r y b b = -\u00ec \u00ed = +\u00ee (A.2) where y is the coordinate value of point K in the y direction. The relationship of the geometry parameters are shown in Fig. A.2. Then the velocity of point K for cone 1 can be calculated as 1 1 .Kv r w= \u00d7 (A.3) where vk is normal to the plane YOZ, which also represents the line-of-action plane 1 1 2 2N N N N\u00a2 \u00a2 " ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002663_j.matpr.2021.05.271-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002663_j.matpr.2021.05.271-Figure2-1.png", "caption": "Fig. 2. S500 carbon fiber X-shaped frame.", "texts": [ "SFiN) (ii) SFij is termed as subject factors to explain the important factors of autonomous UAVs such as speed, wind direction, behavior and position, air pressure, path planning and human input command are part of these subject factors. The mechanical part plays a key role in the development of autonomous quadcopter. This design includes a frame with flexibility in nature, vibration-free and cross-shaped structure. 3D printing of such frames was not so successful because of their different weight ratios to maintain the balance of the quadcopter. Depending on this an S500 carbon fiber X-shaped frame with an integrated power distribution board (PDB) and the landing gear was selected as shown in Fig. 2. The dimensions of the frame are 11.42*7.09*2.36 in. The specification of motors for the quadcopter is Emax MT3506 650Kv as shown in Fig. 3 (a). These brushless motors are equipped with a small circuit for the supply of energy to the motors. These motors are lightweight in design with a longer duration of the flight. 1355 carbon fiber propellers as shown in Fig. 3 (b) are paired with these motors and observed that each motor is generating 9,620 revolutions per minute (RPM) and produces a maximum thrust of 1100 g" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002002_012076-Figure8-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002002_012076-Figure8-1.png", "caption": "Figure 8. Stress on tire with air package", "texts": [ " The calculated forces like bump force were applied by the road to the tire of the wheel with a force of 2500N and the analysis is carried out in the no-slip condition as shown in figure 7. According to the simulation study mentioned, the analysis was carried out in ANSYS workbench and the following results were obtained. The individual analysis was carried out for the existing model and the proposed model with element size was 5mm to get the most efficient results in the stipulated amount of time. The Von-Mises stress, strain and deformation plots are as shown in figure 8, 9, 10 for air package as material for existing tire. Figure 11, 12 and 13 explain the von mises stress, strain and deformation plots for polyethylene as a material. ICMSMT 2020 IOP Conf. Series: Materials Science and Engineering 872 (2020) 012076 IOP Publishing doi:10.1088/1757-899X/872/1/012076 Case 1 \u2013 a) Stress plot b) Strain plot c) Total deformation plot ICMSMT 2020 IOP Conf. Series: Materials Science and Engineering 872 (2020) 012076 IOP Publishing doi:10.1088/1757-899X/872/1/012076 Case 2 \u2013 a) Stress plot b) Strain plot c) Total deformation plot The von-mises stress plot results seem to be accurate based on the convergence theory and compatibility equation which is decided based on the number of iterations carried out for each of the case [10, 11]" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001397_1350650120942329-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001397_1350650120942329-Figure7-1.png", "caption": "Figure 7. Illustration of jet impingement depth on gear (t1\u00bc t).", "texts": [ " A plane OII g perpendicular to the gear axis and passing the impingement point on the gear meshing surface is drawn; then on plane OII g , a line perpendicular to the projection line of the jet stream from the central point OII g is drawn, and the perpendicular length R is R \u00bc Rx cos \u00f025\u00de where: \u00bc 3 2 \u2019 Rx \u00bc x0 y0 sin \u00bc x0 \u00fe y0 cos \u2019 \u00f026\u00de The initial angular positions of the pinion and gear at the initial time t0 are p4 and g2, respectively, thus g2 \u00bc p4=u\u00fe inv sph , b\u00f0 \u00deg\u00fe \u00f027\u00de p4 \u00bc cos 1 XII =rIIa inv a sph a, b\u00f0 \u00dep \u00fe inv sph , b\u00f0 \u00dep \u00f028\u00de where inv a sph a, b\u00f0 \u00dep denotes the declination angle a sph at the intersection point between the outside circle of the pinion and the involute, and inv a sph a, b\u00f0 \u00dep\u00bc 1 sin bp arccos cos ap cos bp arccos tan bp tan ap . 2. At the moment (t1\u00bc t): the position parameters of the spiral bevel gears and the jet streamline are as illustrated in Figure 7. The impingement depth dg at the time t1 can be calculated from RII i 2 \u00bc RII a dg cos 2 ag 2 \u00bc L2 g \u00fe R2 or dg \u00bc RII a L2 g \u00fe R2 1=2 . cos2 ag \u00f029\u00de where RII a \u00bc 1 2 Ng mII n cos II \u00fe hIIag Since the jet time is equal to the rotation time, then g3 g2 !g \u00bc h Vj \u00f030\u00de where g3 \u00bc tan 1 Lg R \u00fe inv g3 sph g3, b g inv g3 sph g3, b g \u00bc 1 sin bg arccos cos g3g cos bg arccos tan bg tan g3g g3g \u00bc g \u00fe tan 1 hIIag dg cos ag RII e 8>>>>>>>>>>>>< >>>>>>>>>>>>: \u00f031\u00de where inv g3 sph g3, b g represents to the declination angle g3 sph at the impingement point in the gear involute" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000187_s10846-019-01113-y-Figure15-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000187_s10846-019-01113-y-Figure15-1.png", "caption": "Fig. 15 The trajectory of base (red line) and end-effector(blue line) a CMM b EMM", "texts": [ " The second simulation is a comprehensive operation involving collision-prediction for all situations, crawling function, cascade connection and configuration-to-Cartesianpose planning. The CMM is supposed to crawl through platforms twice to reach the experiment station. Then EMM should raise its end-effector and get prepared. After that, CMM attaches EMM and carries it to the cargo station where EMM can finish its ORU transport operation. The entire operation is briefly demonstrated in Fig. 14. Video \u2019Comprehensive Operation. avi\u2019 shows the detail. Figure 15 demonstrates the trajectory of base and end-effector for both CMM and EMMwith blue standing for the base and red for the end-effector. In comparison, the trajectory provided by the non-any-angle-path algorithm (basic A*), illustrated in Fig. 16, is neither smooth nor reasonable. Note that the collision between CMM and EMM as well as EMM against the environment should also be considered especially when CMM is approaching and carrying EMM. The collision between ORU and the station should also be avoided when being transported" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001364_7.0000111-Figure29-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001364_7.0000111-Figure29-1.png", "caption": "FIG. 29. Build job for trials on part orientation. FIG. 30. Features investigated by 3D laser scanning.", "texts": [ " Additional material was added to the interface areas for machining tolerance compensation. Figure 28 shows the topology-optimized bracket with thickened interfaces and without holes in the interface. Due to the strong interdependencies between the part orientation and the occurring distortion of manufactured components, three different orientations were investigated with respect to shape deviation. The defined target is to determine the orientation showing the least degree of shape deviation. The Materialise Magics build file including all tested orientations can be seen in Fig. 29. All components of each orientation were successfully manufactured but differed in their individual distortion. The manufactured parts were characterized concerning shape tolerance using the 3D laser scanning system GOM Atos Core. In order to ensure successful integration of the demonstrator, the interfaces need to be postmachined. Therefore, the interfaces were investigated by 3D laser scanning. The investigated features are shown in Fig. 30. All features were analyzed concerning shape deviations in diameter, x-coordinate, y-coordinate, and z-coordinate" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001612_j.oceaneng.2020.108224-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001612_j.oceaneng.2020.108224-Figure1-1.png", "caption": "Fig. 1. The trimaran scale model (a), (b) (different views), and the side view (c).", "texts": [ ", 2019b) (m+ a33)x\u03083 + b33x\u03073 + c33x3 + a35x\u03085 + b35x\u03075 + c35x5 = F3 (1) (I + a55)x\u03085 + b55x\u03075 + c55x5 + a53 x\u03083 + b53x\u03073 + c53x3 = M5 (2) where m is the trimaran mass, I is the longitudinal inertia of the trimaran, x3 represents the trimaran\u2019s heave displacement, x5 represents the pitch angle of the trimaran, aij is the added mass, bij is the damping coefficients, and cij is the restoring coefficient. F3 and M5 are the amplitude of wave force and moment, respectively. To investigate the stabilization method for reducing motion of the trimaran, a model trimaran model was built. Its parameters are listed in Table 1, and Fig. 1 shows the trimaran model itself. Fig. 1(a) and (b) show views of the model from different perspectives, and Fig. 1(c) shows a side view of the trimaran. The T-foil and flap were used as the appendages for the vessel\u2019s vertical stabilization control. A three-view drawing of the appendages is presented in Fig. 2, and the parameters of the appendages are listed in Table 2. Nomenclature aij Added mass bij Damping coefficients cij Restoring coefficients F3 Wave force M5 Wave moment FT Force of T-foil MT Moment of T-foil \u03b8T Attack angle of T-foil lT Distance from the center of the T-foil to the COG of the trimaran FF Force of flap MF Moment of flap \u03b8F Attack angle of the flap lF Distance from the center of the flap to the COG of the trimaran U Speed of the trimaran Hs Significant wave height T1 Wave period in irregular waves SSN Sea state number HSW Heave of the bare trimaran HSY Heave of the trimaran with active appendages \u03b8SW Pitch of the bare trimaran \u03b8SY Pitch of the trimaran with active appendages RHC Reduction rate of the heave motion under active control RPC Reduction rate of the pitch motion under active control S\u03b6(\u03c9) Power spectrum density RFMD-K Resultant force and moment distribution with Kalman filter A State transition matrix P Covariance matrix K Gain of the Kalman filter Q Process noise matrix R Measurement noise matrix H Measurement coefficient matrix Z Measurement vector t Sample time T Wave period in regular waves Because the motion of the trimaran is mainly caused by the wave Z" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001038_0142331219879338-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001038_0142331219879338-Figure5-1.png", "caption": "Figure 5. The horizontal rotational vector-based potential field.", "texts": [ " Define ra as the detection range of obstacle avoidance, and the distance between the UAV to the obstacle as follows ro = ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi x xo\u00f0 \u00de2 + y yo\u00f0 \u00de2 + z zo\u00f0 \u00de2 q \u00f028\u00de The virtual accelaration control input for obstacle avoidance is defined as fnr =(fxnr , fynr , fznr ), and can be developed as follows If ra\\ro fr = fd + fdj jfnr r2 o 1 ro 1 ra else fr = fd \u00f029\u00de where fd = fxl , fyl , fzl for the leader UAV and fd = fxnc + fxn , fync + fyn , fznc + fzn for the follower UAVs. Horizontal obstacle avoidance. Consider the obstacle avoidance situation that the UAV is affected by the horizontal potential field, the virtual control input can be designed as follows: for clockwise direction fxrh = ko v2 v1 y yo\u00f0 \u00de fyrh = ko v1 v2 x xo\u00f0 \u00de fzrh = 0 \u00f030\u00de and for counterclockwise direction fxrh = ko v2 v1 y yo\u00f0 \u00de fyrh = ko v1 v2 x xo\u00f0 \u00de fzrh = 0 \u00f031\u00de where ko is a positive constant. The horizontal vector field is depicted in Figure 5. rn denotes the angle between the line linking the obstacle and the UAV, and the x-axis. fn is the angle between the UAV\u2019s velocity and the x-axis. xn is the angle between the tangent direction of vector field and the x-axis. The expressions of the angles are as follows fn =arctan _y, _x\u00f0 \u00de xn =arctan v2 1xo, v 2 2yo rn =arctan yo y, xo x\u00f0 \u00de \u00f032\u00de Vertical obstacle avoidance. Then consider the obstacle avoidance strategy that the UAV is affected by the vertical potential field. The virtual control input are still designed in two cases: for upward direction fxrv = ko v1v2 v3 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi v1 2 + v2 2 p z zo\u00f0 \u00de cos fn\u00f0 \u00de fyrv = ko v1v2 v3 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi v1 2 + v2 2 p z zo\u00f0 \u00de sin fn\u00f0 \u00de fzrv = ko v3 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi v1 2 + v2 2 p v1v2 x xo\u00f0 \u00de cos fn\u00f0 \u00de v3 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi v1 2 + v2 2 p v1v2 y yo\u00f0 \u00de sin fn\u00f0 \u00de \u00f033\u00de or in downward direction fxrv = ko v1v2 v3 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi v1 2 + v2 2 p z zo\u00f0 \u00de cos fn\u00f0 \u00de fyrv = ko v1v2 v3 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi v1 2 + v2 2 p z zo\u00f0 \u00de sin fn\u00f0 \u00de fzrv = ko v3 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi v1 2 + v2 2 p v1v2 x xo\u00f0 \u00de cos fn\u00f0 \u00de + v3 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi v1 2 + v2 2 p v1v2 y yo\u00f0 \u00de sin fn\u00f0 \u00de \u00f034\u00de The vertical rotational vector field is shown in Figure 6" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001980_1099636220931479-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001980_1099636220931479-Figure6-1.png", "caption": "Figure 6. Typical load-displacement curves of PFRCs and PFRALs.", "texts": [ " Besides, a good metal-composite bonding level could allow more global plastic deformation through the delaying of the localised strain that results in the crack initiation [37]. Owing to the high level of global plastic deformation, PFRALs has undoubtedly shown an outstanding tensile strain. For the PALF fabrics, the fraying effect was observed in the fibres in the loading direction. The 3-point flexural test was performed on those of PFRCs and PFRALs to determine their flexural strength and modulus. Figure 6 shows the load-displacement curves of PFRCs and PFRALs. The flexural properties of PFRCs and PFRALs demonstrated a similar trend to their respective tensile properties. PFRALs exhibited a flexural strength of 181.88MPa which was 309.55% higher than those of PFRCs having a flexural strength of 44.41MPa. This result indicates that the flexural strength was drastically increased with the presence of aluminium as the skin layers of FMLs. The enhancement also could be noticed in the flexural modulus with an increment of 400% when the aluminium layers were bonded to the composites" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0003104_0095-8522(51)90037-2-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0003104_0095-8522(51)90037-2-Figure2-1.png", "caption": "FIG. 2. W0rker-consistometer, cross-section. A cylinders, B lock mechanism, C piston, D orifice plate, E electric heaters, F liquid heat exchanger, G piston rods, H ball and socket joints, I cam, J 'over load device, and K stirrup suspension.", "texts": [ " The nature and magnitude of the piston correction required with greases in this type of instrument and the extent of the temperature rise resulting from shear in highly viscous systems also require evaluation before the validity of the concept of greases as Bingham bodies can be tested. EXPERIMENTAL PROCEDURES AND MATERIALS The worker-consistometer described by McKee and White (1) has been modified to obtain a uniform flow rate throughout the power-driven working stroke and to provide close temperature control over a wide working range by means of electric heaters and liquid coolant channels in the cylinder block itself. The machine (Fig. 2) consists of two opposed cylinders (A) 0.5044 in. in diameter connected by a breech lock mechanism (B) ; each cylinder is fitted with a matching piston (C) of not greater than 0.0002 in. clearance. The cylinders are separated by various 0.2500-in.thick orifice plates (D) containing 0.0135-in.-diameter holes. One plate contains 1 hole and another, 10 holes. A third plate 0.125 in. thick~ containing 11 holes of the same diameter as the previous plates, and an auxiliary spacer or clearance plate of equal thickness, with an opening of 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002236_j.isatra.2020.10.053-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002236_j.isatra.2020.10.053-Figure4-1.png", "caption": "Figure 4: Agonist-antagonist mechanism with SMA springs", "texts": [ " Elongation-rotation relationship Notice that the change in the angular positions the joint depends on the elongation of the SMA sprin actuators, this means that the change in the elong tion produced in the spring modifies the angle in t joint for the SLEMR. This relation can be proposed b geometrical relation as follows: q = sin\u22121 l r , (1 where l is the length of the spring elongation chang r is the know length from the center of the link to t placement of the wire and q is the angle in the joint Based on the configuration seen in Figure 4 can defined that when the spring is elongated at half it total capacity the angular position of the joint equivalent to zero grades. Because the mechanism based on two SMA actuators with a uniqueness w to movement each one, when a negative angular mov ment is required the corresponding actuator elong tion is used and the other is not consider. In oth hand when a positive angular movement is requir the elongation of the corresponding spring is used an the process is repeated. 4. Dynamic model The SLEMR behavior required two different appro imations to model the dynamical aspect, the actu tor dynamic which has been treated extensively in [3 and the dynamic for the mechanical structure that treated in this work using the Euler-Lagrange stru ture" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001128_c9sm02087j-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001128_c9sm02087j-Figure1-1.png", "caption": "Fig. 1 (a) Schematic showing side view of molding procedure used for casting two phase periodic structure (TPPS). (b) Schematic showing top view of two phase periodic structure (TPPS) under film layer. (c) Schematics of control structures, top shows stiff control structure, bottom shows compliant control structure. (d) Optical micrograph of TPPS (e) Schematic of lubricated sliding experiment.", "texts": [ " We find that one can significantly enhance friction of such a two phase periodic structured sample compared to unstructured controls composed of only the higher or the lower modulus material suggesting new mechanisms for friction enhancement. Two possible mechanisms for the lubricated sliding friction enhancement were considered, one due to shear deformation of the material in the sliding direction and the second due to energy loss through the fluid as the indenter transitions periodically from contact with one phase to the other. Samples were made from poly(dimethylsiloxane) (PDMS, Dow Sylgard 184, Dow Corning). The molding process for fabrication of TPPS samples is shown schematically in Fig. 1.a(i\u2013iv). A mold was machined out of aluminum with a ridge channel geometry, as is shown in Fig. 1.a.i. Mold channel width, w1, was 1.2 mm, ridge width, w2, was 0.8 mm, and ridge height, d, was 1 mm. The stiff portion of the structure was fabricated by casting a 10 : 1 base to cross linker mixture of PDMS into the mold as is shown in Fig. 1.a.ii. The PDMS was cured on the mold at 80 1C for 2 hours, then brought to room temperature and removed from the mold as shown in Fig. 1.a.iii. Next a 30 : 1 base to cross linker mixture of PDMS was poured into the cured PDMS to form the soft portion of the structure as is shown in Fig. 1.a.iv. The structure was cured again at 801 for 2 hours with a weighted glass slide on top of the 30 : 1 mixture, resulting in only a thin layer of the 30 : 1 PDMS mixture covering the surface of the sample, making the entire surface chemically homogeneous. The final dimensions for the TPPS samples are labeled in Pu bl is he d on 0 6 Ja nu ar y 20 20 . D ow nl oa de d by U ni ve rs ite P ar is D es ca rt es o n 2/ 14 /2 02 0 9: 45 :2 4 PM This journal is\u00a9The Royal Society of Chemistry 2020 Soft Matter, 2020, 16, 1627--1635 | 1629 Fig. 1.a.iv (side view) and Fig. 1b (top view, under thin film layer) and are as follows; structure height, d, of 1 mm, thin film thickness, t, of approximately 70 microns, total sample height, h, of 2 mm, stiff phase width, w1, of 1.2 mm, soft phase width, w2, of 0.8 mm, and structure period, w of 2 mm. As the thin film was obtained using compression during curing, there is a small undulation of the surface in the direction normal to the stripes, with amplitudes less than 100 microns. The Young\u2019s modulus of the stiffer phase (10 : 1 base to cross linker ratio) is denoted by E1 and is approximately 3 MPa,50 while the Young\u2019s modulus of the softer phase (30 : 1 base to cross linker ratio) is denoted by E2 and is approximately 190 kPa.51 Two controls were fabricated to test friction properties of the two phases, and are shown schematically in Fig. 1c. The compliant control sample is shown in the bottom of Fig. 1c and is a 2 mm thick slab of 30 : 1 base to cross linker mixture of PDMS, cured at 80 1C for two hours. The stiff control sample is shown in the top of Fig. 1c, and was fabricated by first making a 2 mm thick slab of 10 : 1 base to cross linker mixture of PDMS, cured at 80 1C for two hours. Then, in a procedure similar to the TPPS fabrication, a layer of 30 : 1 base to cross linker PDMS mixture was poured onto the cured control sample and cured at 80 1C for two hours while under a weighted slide, resulting in a thin layer of the softer material coating the top of the sample. Sliding friction in a direction orthogonal to the stripes was measured under lubricated conditions. A schematic of the sliding experiment is shown in Fig. 1.e. The surface of the sample was coated with a layer (approximately 1 mm thick) of a highly wetting lubricant (PDMS base, viscosity Z = 5.1 Pa s). A spherical glass indenter (R = 0.5, 2 mm, or 3 mm) was brought into contact with the sample surface under a normal load ranging from 18.6 to 238.1 mN and the sample was moved perpendicular to the applied load using a variable speed motor (Newport ESP MFA-CC) with velocities ranging from 0.025 to 1 mm s 1. The indenter was connected to a load cell (Honeywell Precision Miniature Load Cell) measuring the friction force resisting sample motion" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001237_j.ymssp.2020.106783-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001237_j.ymssp.2020.106783-Figure1-1.png", "caption": "Fig. 1. 13 DOFs vehicle dynamic model with semi-active hydraulic damping strut.", "texts": [ " Section 5 provides the analysis results of applying the optimized parameters of the semi-active HDS in the 13 DOFs vehicle dynamic model when engine start and automatic start. Section 6 provides some works on verification experiments, the simulated and experiment results are analyzed to confirm the validity of the model. Finally, Section 7 draws conclusions. The 13 DOFs vehicle dynamic model is established in order to analyze the effect of the semi-active HDS on the vehicle vibration when engine start-stop, which includes the PMS, vehicle body, tires, the sprung mass, the suspension and the semi-active HDS, as shown in Fig. 1. The vehicle uses a front engine and front wheel drive system with a transverse engine layout. Under engine excitation, the displacement of the center of gravity (CoG) for the PMS can be described as qp T = (xp, yp, zp, ap, bp, cp). The displacement of the CoG for the vehicle body can be described as qb T = (zb, ab, bb). The vertical displacement of each unsprung mass can be described as qu T = (qu1, qu2, qu3, qu4). In this paper, the dynamic reaction force from the mounting system to the powertrain is written by the sum of two parts, which is generated by the mount stiffness and mount damping" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000787_rnc.4560-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000787_rnc.4560-Figure1-1.png", "caption": "FIGURE 1 The simple pendulum system", "texts": [ " To avoid the high gains, Sun et al32 employed the novel disturbance observer to estimate external disturbances, which makes that smaller controller gains can be chosen. In this paper, we combine the novel disturbance observer with finite-time control and ensure the more general system (1) to obtain a better property of PFTS. In this section, to demonstrate the effectiveness of the designed scheme, we try to use the presented controller to stabilize the inverted pendulum system and a numerical system. Example 1. Consider the inverted pendulum system shown in Figure 1, where g= 9.8 m/s2 is the acceleration due to gravity, M is the mass of the cart, m is the mass of the pole, \ud835\udf17 is the angle that the pendulum makes with the vertical, l is the half length of the pole, and u is the applied force (control). As shown in the work of Slotine and Li,6 the dynamic equation is given by A particularly interesting task is to design a controller to bring the inverted pendulum from a vertical-down position at the middle of the lateral track to a vertical-up position at the same lateral point, ie, the pendulum angular position moves clockwise or counterclockwise until it reaches the equilibrium point (\ud835\udf17, " ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000976_s11044-019-09704-1-Figure39-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000976_s11044-019-09704-1-Figure39-1.png", "caption": "Fig. 39 Load distribution of thin-walled four-point contact ball bearing under radial load", "texts": [], "surrounding_texts": [ "The dynamic performance of thin-walled four-point contact ball bearing is affected by the static and dynamic loads of the wrist, hand, and workpiece or tools of the robots, especially the impact force brought by inertial force and moment under high velocity or acceleration. Considering multi-clearances and multibody contacts, the dynamic contact and impact of ball-to-ring raceway surface contacts, ball-to-cage pocket surface contacts, and cage-toouter ring guidance surface contacts of the bearing are investigated by using the multibody contact dynamic analysis models in this work. The proposed models of the bearing are solved by HHT algorithm with ADAMS software package. The results of dynamic contact forces, impact forces, displacements, and velocities are discussed under the conditions of different complex loads. Due to the effect of multi-clearances and multibody contacts, the impact forces of ball-to-cage small pocket surface contacts are larger in the non-load zone under the conditions of pure radial load or combined rotating radial load, axial load, and positive moment. The dynamic characteristics will be underestimated when ignoring the effect of multibody contact dynamics of the bearing using the empirical design method. The rules of static load distribution and the simulation model of the bearing are verified by the theoretical value. Considering pure radial load, there\u2019re always four-point contacts of ball-to-ring raceway surface contacts in the load zone. The dynamic contact forces are similar to the rule of the static load distribution. The impact forces of ball-to-cage small pockets surface contacts are large, the angular velocities of ball rotation and cage are vary- ing greatly in the non-load zone. The dynamic characteristics of the bearing are stable under the condition of combined radial and axial loads. There\u2019re four-point contacts in the load zone and two-point contacts in the non-load zone. The dynamic contact force of the main contact pair is greatly different from the secondary contact pair. The maximal contact force and the angle of load distribution of the main contact pair are both larger than those of the secondary contact pair. The dynamic contact force of the main contact pair of ball-to-ring raceway surface contacts has one peak and one valley, and is nonzero in the non-load zone. The dynamic contact force of the secondary contact pair is similar to the rules of static load distribution and is zero in the non-load zone. The angular velocities of the ball rotation and cage are stable and periodically varying as a result of the small impact forces of ball-to-cage small pockets surface contacts. The moments are important to the rules of dynamic contact forces of ball-to-ring raceway surface contacts and angular velocities of the ball rotation. The dynamic contact forces of the main contact pair in the load zone are increasing as positive moments increase. The dynamic contact forces of the secondary contact pair in the load zone are increasing as the negative moments increase. The effects are an increase in the radial load and reduction in the preload. The impact forces of ball-to-cage small pockets surface contacts in the non-load zone are large as a result of the effect of rotating radial load. The motion trajectories of outer ring center are as from circular whirling motion. Under the conditions of the proposed loads, the motion trajectories of cage center are always similar to a circular motion trajectory. As a result of the slight impact forces of ball-to-cage big pockets surface contacts, the balls in the cage big pockets are always purely rolling in the load zone and slightly varying in the non-load zone. The motion stability is high and dynamics characteristics are complex for the bearing. In the proposed work, a new approach is presented and the calculated results are illustrated. The dynamics performance and motion accuracy of thin-walled four-point contact ball bearing are complicated. They\u2019re influenced by the complex load conditions, multi-clearances and multibody contacts of ball-to-ring raceway surface contacts and ball-to-cage pockets surface contacts. The influences of the geometrical parameter, multi-clearance and flexibility on multibody contact dynamic analysis, and motion accuracy of thin-walled four-point contact ball bearing will be investigated under varying complex working conditions in the future. Acknowledgements The authors would like to express sincere thanks to the referees for their valuable suggestions. This project is supported by National Natural Science Foundation of China (grant nos. 11462008 and 11002062) and Natural Science Foundation of Yunnan Province of China (grant no. KKSA201101018). This support is gracefully acknowledged. Publisher\u2019s Note Springer Nature remains neutral with regard to jurisdictional claims in published maps and institutional affiliations." ] }, { "image_filename": "designv11_5_0001824_j.euromechsol.2020.103994-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001824_j.euromechsol.2020.103994-Figure3-1.png", "caption": "Fig. 3. Basic geometry of the nozzle with details omitted [dimensions in mm].", "texts": [ ", 1995; Zu et al., 2010b) FTW \u00bc \u03c31F1 \u00fe \u03c32F2 \u00fe \u03c32 1F4 \u00fe \u03c32 2F5 \u00fe \u03c32 6F6 \u00fe 2\u03c31\u03c32F7 (33) where F1 \u00bc 1 S1t 1 S1c F2 \u00bc 1 S2t 1 S2c F4 \u00bc 1 S1tS1c F5 \u00bc 1 S2tS2c F6 \u00bc 1 S12 F7 \u00bc 1 2 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi S1tS1cS2tS2c p (34) in which SXY corresponds to the strength in X direction (1 in fiber direction and 2 parallel to the fiber direction) regarding Y type of stress (\u2018t\u2019 and \u2018c\u2019 as tensile and compressive stress, respectively). Also, S12 defines the in-plane shear strength. Fig. 3 depicts the schematic design of a rocket motor composite nozzle, which is a prototype for a propulsive system of a re-entry satellite. The internal geometry defines the mandrel dimensions, which was machined in AISI 1045 steel with a deposited chrome layer. The equation of the radius with respect to the axis of revolution (generatrix) is r\u00f0x\u00de \u00bc 17:9\u00fe 0:480329x 0:00226208x2 \u00fe 5:5731:10 6x3 \u00bdmm\ufffd (35) The internal pressure on the nozzle is p \u00bc 4:5 MPa. This pressure is due to the expansion of combustion gases of the propulsive system" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000861_012046-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000861_012046-Figure3-1.png", "caption": "Figure 3. Picture of the fracture surface of the leaf spring of number 8.", "texts": [ " Figure 2 shows a schematic drawing that explains the arrangement of springs from number 1 to number 8. In this case the broken one is spring number 8 as shown in Figure 2. International Seminar on Metallurgy and Materials IOP Conf. Series: Materials Science and Engineering 541 (2019) 012046 IOP Publishing doi:10.1088/1757-899X/541/1/012046 Visual observations on the fracture surface of leaf spring evaluated were around the bolt hole area. From visual observation, it was found that there was an initial crack on the side of the bolt hole as shown in Figure 3. Furthermore, the beach mark can be seen on the fracture surface. This finding indicates that the spring has experienced fatigue fracture [8]. Figure 4 shows that the image taken from initial crack tip area as shown in Figure 3. It can be seen that the crack initiationoccurred here which is characterized by relatively smooth fracture surfaces which indicates that fatigue cracking has occurred. Fracture surfaces of leaf springs (Figure 3 and 4) show that fatigue crack propagation has occurred which starts from the initial crack, then crack propagation occurs in line with increasing dynamic load so that the final fracture is split in two parts [8][9]. International Seminar on Metallurgy and Materials IOP Conf. Series: Materials Science and Engineering 541 (2019) 012046 IOP Publishing doi:10.1088/1757-899X/541/1/012046 In this study also conducted a hardness test using Rockwell Type Zwick / Roell ZHR to measure hardness on the surface and core of leaf spring specimens [10]" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001960_j.mechatronics.2020.102366-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001960_j.mechatronics.2020.102366-Figure1-1.png", "caption": "Fig. 1. Schematic of the prototype two-wheeled self-balancing vehicle.", "texts": [ " Section V is devoted o the comparative simulations and experiments. Finally, conclusion is rawn in Section VI. . Problem statement: two-point BVP .1. Dynamic model According to the outward configuration, the TWMIP self-balancing ehicles can be classified into standing types [ 1 , 3 ] and wheelchair types 2 , 4 ]. Among the underactuated mechanical systems, it belongs to the crobot [20] , where the inverted pendulum is indirectly activated by the eaction to the wheel torque input. As the schematic is shown in Fig. 1 , he dynamic model of a self-balancing vehicle has three movable rigid odies, the left and right wheels and the inverted pendulum body pivted about the wheel axis. The overall chassis of the vehicle is driven by he wheel rotations and it performs forward/back and steering motion hile the pitch motion of the inverted pendulum is caused by the reacion torque. Hence, the three-degrees-of-freedom motion is implemented ust by the two inputs of both wheels. The details on the mathematical odeling and the dynamic characteristics can be consulted in [28] ", " During the initial transient phase when a self-balancing vehicle acelerates for speed up or decelerates for speed down, its postural stability s governed by the nonlinear inertial coupling effect between the pitch nd forward motion. In contrast, the cross coupling to or from the yaw otion is relatively weak. In general, the yaw motion planning does not ontribute to the pitch stability. Accordingly, the central issue in designng a reference input for the control of a self-balancing vehicle must be ow to find a set of pitch and forward motion trajectories satisfying the oupled dynamics under the underactuated input constraint. For the generalized coordinates designated in Fig. 1 and the parameers in Table 1 , the longitudinal dynamic model of TWMIP without yaw otion can be organized as \ud835\udc40 \ud835\udc65 \ud835\udc40 \ud835\udf03 ] \ud835\udc5e + [ \ud835\udc36 \ud835\udc65 0 ] + [ 0 \ud835\udc3a \ud835\udf03 ] = [ \ud835\udf0f\u2215 \ud835\udc5f \u2212 \ud835\udf0f ] w w \ud835\udf0f f 2 p t t( w s h w \ud835\udf03 o l \ud835\udc65 d b t ( v t m u t t i \ud835\udf03 e t o s t d t t d s t v s r i a e p c a v p t fi 3 3 c i t t m t v t o [ p m o h a i c a r f ith \u23a7 \u23aa \u23aa \u23a8 \u23aa \u23aa \u23a9 \ud835\udc40 \ud835\udc65 = [ \ud835\udc5a \ud835\udc35 + 2 \ud835\udc5a \ud835\udc4a + 2 \ud835\udc3d\u2215 \ud835\udc5f 2 \ud835\udc5a \ud835\udc35 \ud835\udc59 cos \ud835\udf03 ] \ud835\udc40 \ud835\udf03 = [ \ud835\udc5a \ud835\udc35 \ud835\udc59 cos \ud835\udf03 \ud835\udc3c 2 + \ud835\udc5a \ud835\udc35 \ud835\udc59 2 ] \ud835\udc36 \ud835\udc65 = \u2212 \ud835\udc5a \ud835\udc35 \ud835\udc59 \u0307\ud835\udf03 2 sin \ud835\udf03, \ud835\udc3a \ud835\udf03 = \u2212 \ud835\udc5a \ud835\udc35 \ud835\udc59\ud835\udc54 sin \ud835\udf03 \ud835\udc5e = [ \ud835\udc65 \ud835\udf03 ]\ud835\udc47 (1) here the first inertia matrix is always invertible and the single input is equivalent to the sum of left and right wheel torques, while the rictional effects from the ground and wheel axis are neglected", " The notion and rocedure of the semi-online trajectory planning proposed in this paper an be extended to other types of underactuated system without much oss of generality, as far as the dynamic model can be transformed to an nput-output normal form like (2). Finally, it can be said that the closedorm solution (10) to (12) corresponds to the first result on the dynamic odel-based online trajectory planning for an underactuated system. Fig. 9. Distribution of the algebraic coefficients ( \ud835\udc61 \ud835\udc53 = \ud835\udc47 ). 5 5 T g p r c r v F e t o a t j m c s v a m B d r w t c . Numerical examples .1. Simulations First, a quick start problem is considered for the prototype in Fig. 1 . he semi-online reference trajectories according to (10) and (12) are enerated with the same boundary conditions given in Fig. 4 and aplied to the velocity and pitch control loop in Fig. 5 . According to the eference inputs, the translational movement and posture variation are ompared in Fig. 11 . When the target velocity is assigned as the step eference input, the pitch motion and the arrival time at the target elocity are greatly influenced by the feedback gains as indicated in ig. 11 (b) and (c)", " In the case of human-riding transporters [1\u20134] , the vehicle motion an be actively stabilized by the rider\u2019s visual feedback because the rider s to generate reference inputs in real time. For example, if the rider moothly inclines the pitch angle of the vehicle forward or backward, he forward velocity can be increased or decreased in the course of sta- ilizing the pitch angle to the zero reference. The resulting trajectory rofiles during the transient phase are expected to be quite close to the olutions in this paper. .2. Experiment The hardware control system of the prototype in Fig. 1 consists of two heel motors, encoders to estimate the forward velocity, attitude sensor rate gyro) to detect the pitch angle, and PC-based controller to imple- ent the feedback control law and the semi-online trajectory generation lgorithm. The reference command and control input are updated every 0 msec synchronized with the attitude sensor output. The comparative experimental results are shown in Figs. 15 to 17 . he vehicle is accelerated from the stationary point to the target velocity nd finally decelerated to zero along the straight path in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002098_j.automatica.2020.109192-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002098_j.automatica.2020.109192-Figure1-1.png", "caption": "Fig. 1. Engagement Geometry.", "texts": [ ", 2020; V\u00e1\u0148a & Faigl, 2015) even n the scenarios where the assumption that the target regions re separated at least by four times the minimum curvature is elaxed. This is demonstrated by the numerical simulations in ection 6. At a preliminary stage, some results of the paper appear in Jha, hen, and Shima (2020). In this work we have expanded on these esults by proposing more geometrical properties of the optimal ath and establishing analytical solutions for all the categories of he optimal path. . Problem formulation Consider a circular region of radius r \u2208 R+ with its center at he origin of an inertial coordinate system OXY , as represented in Fig. 1. For notational simplicity, hereafter we denote this circle by C, i.e., C = {(x, y) \u2208 R2 |x2 + y2 = r2} In the frame OXY , the state x = [x, y, \u03b8] \u22a4, also called configration, represents the position (x, y) \u2208 R2 and the orientation ngle \u03b8 \u2208 S1 of a Dubins vehicle whose minimum turn radius is \u2208 R+. The kinematics of this vehicle is expressed as dx dt = f (x, u), u \u2208 [\u22121, 1] (1) where f (x, u) is given as, f (x, u) = [ cos \u03b8 sin \u03b8 u \u03c1 ]\u22a4 ere, u \u2208 [\u22121, 1] is the control input and t denotes time. The paper aims to find the shortest path for Dubins vehicle from an initial configuration x0 = [x0, y0, \u03b80]\u22a4 at time t0 \u2208 R+ to a given final configuration xf = [ xf , yf , \u03b8f ]\u22a4 via the boundary of the circle C" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002068_b978-0-12-819661-8.00011-1-Figure11.11-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002068_b978-0-12-819661-8.00011-1-Figure11.11-1.png", "caption": "FIG. 11.11 Basic structure of a typical DSSC (Mohiuddin et al., 2017).", "texts": [ "10C and D presents schematic of the resulting OPT array and a photo of the free standing device which was highly flexible with thickness of less than 4 mm. Dye-sensitized solar cell (DSSC) is one type of organic-based PV solar cells. PV cells of the heterojunction structure, on the other hand, are known to provide conversion efficiency higher than that of the other solar cell types. This is mainly because such structure can provide larger area of interface between the electroactive donor and acceptor biopolymers. One example of the basic structure and principle of operation of DSSCs is illustrated in Fig. 11.11 (Mohiuddin et al., 2017). When the photosensitive dye is exposed to sunlight, some of the valence band electrons in the p-type material will have sufficient energy to overcome the material\u2019s energy gap and jump to the conduction band to be free to move. Once an electric load is connected across the cell, an electric current resulting from the movement of these conduction electrons will be established. In spite of the many advantages they enjoy, biopolymer-based PV solar cells still have some weakness points that form real challenges to this industry" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000024_1350650119858243-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000024_1350650119858243-Figure2-1.png", "caption": "Figure 2. The tilt motion of the rotor-bearing system. (a) xyz plane, (b) xy plane, (c) yz plane, and (d) xz plane.", "texts": [ " Since the axial velocity of the journal in the bearing bush is zero, the steady Reynolds equation can be obtained as @ @x ph3 @p @x \u00fe @ @z ph3 @p @z \u00bc 12 @ ph\u00f0 \u00de @t \u00fe 6 u @ ph\u00f0 \u00de @x \u00f02\u00de The following dimensionless parameters are introduced h \u00bc h=c0, \u00bc x=R, z \u00bc z=R, p \u00bc p pa, \u00bc 6 !R2 pac02 , \u00bc !t, \u00bc !s=! Equation (2) can be written as @ @ h 3 @p2 @ \u00fe @ @z h 3 @p2 @z \u00bc 2 @ ph @ \u00fe 4 @ ph @ \u00f03\u00de where pa is the ambient pressure, ! is the rotating speed, c0 is the radial clearance between the journal and bearing, R is the journal radius, !s is the journal perturbation rotating speed, and is the perturbation frequency ratio. The tilt motion of rotor and the misaligned position of bearing can cause the change of film clearance, as shown in Figure 2. The change of clearance can lead to the change of pressure distribution, which has a great influence on the dynamic performance of the rotor-bearing system. Figure 2(a) shows the rotor is in equilibrium and the initial position of the rotor is (x0, y0, \u2019x0, \u2019y0). The eccentricity and attitude angle of the bearing in the middle cross section (xy plane) are e0 and 0, respectively, as shown in Figure 2(b). The rotor tilts around the x axis and y axis, respectively, with respect to the tilt center zm, as shown in Figure 2(c) and (d). Figure 3 shows the change of air film thickness due to a small perturbation motion. Figure 3(a) gives the change of air film thickness due to translating perturbation of x and y, the corresponding eccentricity and attitude angle become em and m. Figure 3(b) gives the change of air film thickness due to rotating perturbation of \u2019x and \u2019y, the corresponding eccentricity and attitude angle become e and . When a small perturbation of x, y, \u2019x, \u2019y is applied to the rotor-bearing system, the gas film thickness of the journal bearing at any position can be determined using the following equations16 h \u00bc h0 \u00fe x\u00fe \u00f0z zm\u00de \u2019y cos \u00fe y \u00f0z zm\u00de \u2019x\u00f0 \u00de sin h0 \u00bc c0 \u00fe e0 cos 0\u00f0 \u00de \u00fe z zm\u00f0 \u00de \u2019y0 cos \u2019x0 sin \u00f04\u00de The following dimensionless parameters are introduced h0 \u00bc h0=c0, \" \u00bc e0=c0, x \u00bc x=c0, y \u00bc y c0, 0 \u00bc arctan 2c0\u00f0 \u00de L, \u2019y \u00bc \u2019y 0, \u2019x \u00bc \u2019x= 0 Equation (4) can be written as h \u00bc h0 \u00fe l1 z zm\u00f0 \u00de \u2019y cos \u2019x sin \u00fe x cos \u00fe y sin h0 \u00bc 1\u00fe \" cos 0\u00f0 \u00de \u00fe l1 0 z zm\u00f0 \u00de \u2019y0 cos \u2019x0 sin ( \u00f05\u00de where l1 \u00bc R 0=c0" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002347_b978-0-12-803581-8.12081-8-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002347_b978-0-12-803581-8.12081-8-Figure3-1.png", "caption": "Fig. 3 Complex structure made from RBSiC by the BJ process with dimensions: 880 750 300 mm, volume 9.000 cm3. Image courtesy of Schunk Ingenieurkeramik GmbH, Germany.", "texts": [ " Infiltration of liquid silicon at 15001C in vacuum finally lead to the RBSiC material, also designated as SiSiC. Moon et al. (2001), instead, printed by BJ a carbon preform which had 48% open porosity, by using a binder based on furfuryl resin. Following infiltration of Si at 14501C in nitrogen led to the desired RBSiC composite. It is interesting to notice that in this latter example part of the carbon was derived from the polymeric binder, and this allowed the generation of functionally graded parts by varying the dosage of carbon-yielding binder and thus controlling the amount of SiC formed. Fig. 3 shows a voluminous SiSiC part, which has been obtained by BJ of SiC powders followed by Si infiltration during sintering. The first two approaches may be however difficult to apply to pure technical ceramics, such as alumina or silicon nitride. A third approach has been developed which consists of cold isostatic pressing (CIP) or warm isostatic pressing (WIP) the green body printed by BJ, before sintering. Yoo et al. (1993) obtained almost dense parts (99.2% relative density) by applying WIP to alumina parts (with MgO dopant) which had a flexural strength of 324 MPa after sintering" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000976_s11044-019-09704-1-Figure10-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000976_s11044-019-09704-1-Figure10-1.png", "caption": "Fig. 10 Cylindrical surface compressive stress distribution of ideal line contact [1]", "texts": [ "15 \u00d7 105 ( \u03b4\u2217 c )\u22123/2 (\u2211 \u03c1 )\u22121/2 c , (13c) where Ki, Ko, and Kc are contact stiffnesses of ball-to-inner ring raceway, ball-to-outer ring raceway, and ball-to-cage pocket, respectively. The calculation procedure for roller bearings is similar to that of ball bearings. The major difference is the load\u2013deformation relation. Jones\u2019 model [2] was improved by Harris [1] and it was called as the Jones\u2013Harris method (JHM). Harris detailed the Jones\u2019 model by applying a variety of loading conditions and bearings. The ideal line contact of cylindricalto-cylindrical surface contacts (cage\u2013outer ring guidance contact) is shown in Fig. 10. For an ideal line contact, the max stress is expressed as \u03c3max = 2F \u03c0Lb , (14) where L is the effective width of cage\u2013outer ring guidance contact. For steel roller bearings, the semiwidth of the contact surface may be approximated by b = [ 4F \u03c0L \u2211 \u03c1 ( 1 \u2212 v2 I EI \u2212 1 \u2212 v2 II EII )]1/2 , (15) where E and v are modulus of elasticity and Poisson\u2019s ratio. The contact deformation for a line contact condition was determined by Lundberg and Sjovall [42] to be \u03b4 = 2F(1 \u2212 v2) \u03c0EL ln [ \u03c0EL2 F(1 \u2212 v2)(1 \u2213 \u03b3 ) ] , (16) where \u03b3 is shear strain" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001738_j.mechmachtheory.2019.103629-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001738_j.mechmachtheory.2019.103629-Figure5-1.png", "caption": "Fig. 5. Direct singular configuration.", "texts": [ " In this type of singularities, by locking the end-effector, the i th actuator can have an infinitesimal motion and an infinitesimal movement of the end-effector along particular direction cannot be accomplished. Direct kinematic singularity arises when the direct Jacobian matrix ( J D ) is rank deficient. In this type of singularities, as opposed to the other one, the end-effector of the manipulator gains one or more uncontrollable degrees of freedom while the actuators are locked. Referring to Eq. (25) , direct kinematic singularities arises when the determinant of J D is zero which happens in one of the following conditions: \u2022 when l 1;1 is parallel to l 1;2 ( Fig. 5 (a)). \u2022 when l 2;1 is parallel to l 2;2 ( Fig. 5 (b)). \u2022 when s 1 is parallel to s 2 . When this happens, all links of the rhombus becomes parallel to each other and the rhombus changes it shape and becomes a line ( Fig. 5 (c) and Fig. 5 (d)). The reachable workspace of a manipulator is the region in the space that every point can be reached by the manipulated platform in at least one orientation [45] . It consists of the points that the inverse kinematic problem has solution. However, a workspace that includes singularities or a workspace position that can only be reached with some orientation is typically not useful for a robot manipulator. Fig. 6 , illustrates the largest singularity-free box-shaped workspace that can be reached with any orientation of tool" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000226_s10704-019-00417-2-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000226_s10704-019-00417-2-Figure2-1.png", "caption": "Fig. 2 Image of the extensometer position in relation to the specimen crack tip", "texts": [ " Crack growth rates before and after overloading were determined by the slope of the line of the two consecutive points. The results were displayed as da/dN versus K curves. The load-displacement data was acquired at specific crack length increments, for both cases (constant amplitude and overloading conditions). The displacement was obtained through two different systems: using a extensometer, model MTS, with a a maximum displacement of \u00b12.5 mm, and an Allied Vision Stingray camera (20 + 75 mm) to take images and posteriorly process them by digital image correlation (DIC) with the GOM correlate software. Figure 2 shows the position of the extensometer in relation to the specimens. This figure also shows the face of the specimen painted with the black and white pattern required for the DIC process. Using the load and displacement data, it is possible to quantify the crack-opening load (Pop) through the correlation coefficient maximization (Allison et al. 1988). It consists of taking the upper 10% of the loaddisplacement data and calculating the least-squared correlation coefficient. The next data pair is then added, the correlation coefficient is again computed, and this procedure is repeated for the whole data set" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000557_aim.2016.7576989-Figure8-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000557_aim.2016.7576989-Figure8-1.png", "caption": "Figure 8. Straight-fiber-type artificial muscle", "texts": [ " Upon application of air pressure, the artificial muscle expands on the outside, and the cylindrical tube expands on the inside. Fig. 7 shows the unit in the normal and pressurized states. Table II shows the specifications of the peristaltic conveyor. We made a peristaltic pump with diameter, maximum length, and minimum length of 30 mm, 235 mm, and 220 mm, respectively. This conveyor comprises 4 units. Moreover, it is possible to increase the number of units because each unit is connected to a different air supply pipe for independent and smooth operation. Fig. 8 shows a straight-fiber-type artificial muscle [14]. The artificial muscle comprises circularly arranged natural rubber latex and a carbon sheet. The carbon sheet is arranged parallel to the axial direction of the conduit and is made from thin carbon fibers. When the artificial muscle is pressurized, it expands only in the radial direction because the carbon fibers do not readily expand axially. Thus, the artificial muscle contracts in the axial direction. The motion of the peristaltic pump is specified by three parameters: wavelength, propagation speed, and wave number" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001300_j.dt.2020.04.010-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001300_j.dt.2020.04.010-Figure1-1.png", "caption": "Fig. 1. Schematic structure of the ammunition manipulator.", "texts": [ " The effectiveness and practicability of the proposed BLF-DSC strategy is verified by simulation and experimental results. We discuss the mathematical model and problem formulation and the design procedure and stability proof of the proposed BLFDSC controller in Section 2 and Section 3 respectively. The simulation results are demonstrated in Section 4 followed by the experimental results in Section 5. Finally, the sixth section draws the conclusions. 2. Mathematical model and problem formulation The position map of ammunition manipulator is depicted in Fig. 1. The trunnion is fixed at point O. Both ends of hydraulic Please cite this article as: Nie S-C et al., Barrier Lyapunov functionsammunition manipulator electro-hydraulic system, Defence Technology, cylinder are hinged spectacularly with top carriage at point M and coordinator body at point N. The mathematical model was established through theoretical formulation and derivation in this section. The load moment balance equation of ammunition manipulator can be presented as J\u20aca \u00bc Fh\u00f0a\u00de b _xph\u00f0a\u00de Ghg\u00f0a\u00de \u00fe Td (1) where J is the equivalent rotational inertia of the load" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001469_tmag.2020.3019821-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001469_tmag.2020.3019821-Figure4-1.png", "caption": "Fig. 4. No-load flux distribution concerning rotor armature winding of 12- stator-slot/14-rotor-slot DF-NSNS (a) Rotor position at 0 degree. (b) Rotor position at 7.5 degrees (1/4 stator pole pitch). (c) Rotor position at 15 degrees (2/4 stator pole pitch). (d) Rotor position at 22.5 degrees (3/4 stator pole pitch).", "texts": [ "4 degrees (1/4 rotor pole pitch). (c) Rotor position at 12.9 degrees (2/4 rotor pole pitch). (d) Rotor position at 19.3 degrees (3/4 rotor pole pitch). Fig. 3. Stator phase flux linkages of DF-NSNS and DF-N/Fe/N/Fe under noload condition. When analyzing the operation principle of rotor armature winding, the DF-NSNS machine can be regarded as an outerrotor surface-mounted PM (SPM) machine. The rotor phase flux linkage is almost zero when the rotor tooth is aligned with the stator tooth or stator slot opening, as shown in Fig. 4(a) and Fig. 4 (c), while it reaches maximum values when the rotor position is shown as Fig. 4 (b) and Fig. 4 (d). For DF-NSNS and DF-N/Fe/N/Fe machines with 12 stator slots and 14 rotor slots, the mechanical angle corresponding to one rotor electrical period is 30 degrees owing to 12 stator pole pairs and the fundamental frequency of rotor flux linkage can be expressed as / 60r sf Z n (2) where Zs is the number of stator teeth. -80 -40 0 40 80 0 5 10 15 20 25 S ta to r fl u x lin k a g e ( m W b ) Rotor position (mech. deg.) DF-NSNS DF-N/Fe/N/Feb d ca -80 -40 0 40 80 0 5 10 15 20 25 30 R o to r fl u x l in k a g e ( m W b ) Rotor position (mech", ") DF-NSNS DF-N/Fe/N/Fe b d a c Authorized licensed use limited to: Carleton University. Downloaded on September 21,2020 at 00:08:11 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. Fig. 5. Rotor phase flux linkages of DF-NSNS and DF-N/Fe/N/Fe under noload condition. As presented in Fig. 5, the labels in rotor phase flux linkage waveform of DF-NSNS correspond to the subgraphs in Fig. 4. It worth mentioning that the rotor flux linkages of DF-NSNS and DF-N/Fe/N/Fe are almost the same, since the effect of consequent-pole structure on SPM machines is very little. The four aforementioned machines have similar winding configurations. All of them share the same stator star of slots as shown in Fig. 6(a) and for DF-FRPM machines, both DFNSNS and DF-N/Fe/N/Fe share the same rotor star of slots as presented in Fig. 6(b). In this section, the electromagnetic performances of the four machines are analyzed and compared by finite element (FE) method in terms of no-load, on-load and over-load capabilities" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002543_j.fusengdes.2021.112309-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002543_j.fusengdes.2021.112309-Figure7-1.png", "caption": "Fig. 7. Semi-detached ribs installed inside of cooling channel.", "texts": [ " The semi-detached ribs are positioned onto riveting pins machined on the bottom side (heat flux facing surface during operation) of the cooling channel and fixed by plastic deformation of the rivet pin. The precise distance in between the wings of the ribs and the channel internal surface (~1/10 mm) is provided by a spacer ring. The ring is machined into the channel surface located below the riveting pin. The semi-detached ribs, the bottom side of the channel with spacer rings and riveting pins as well as the installed semi detached ribs inside of the channel are shown in Fig. 7. The installation of the ribs can be added into the process chain shown in Fig. 1 directly after step 1) and has been already demonstrated to be compatible to the following process steps by production of another test part as shown in Fig. 4 however equipped with semi-detached ribs inside. A huge variety of channel surface conditioning can be offered by installation of semi detached structures which can be also combined with the option of Laser surface structuring already demonstrated in 2019 [2]" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001287_j.ijmecsci.2020.105639-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001287_j.ijmecsci.2020.105639-Figure3-1.png", "caption": "Fig. 3: Considered geometries for the geometrically exact determination of the relative", "texts": [ "25 and 30\u00b0 the influence of intersecting wall segments attains a higher influence, thus falsifying the predictions of the presented analytical formulas. As a consequence, especially for structures with higher wall thicknesses the effect of overlapping segments becomes pronounced and thus affects the results accuracy. Thus, an exact model is presented in the following to address this shortfall. Given that we consider structures that are periodic concerning two directions while being invariable in the third direction, a two-dimensional geometrically exact consideration of representative unit cells is sufficient. Fig. 3 shows the geometries considered in this work, and Tab. 1 includes the exact equations for the determination of the relative density */S. Fig. 2 shows the influence of the overlap in the intersections for higher wall thicknesses and relative densities. Fig. 4 presents a schematic representation of the relative density */S for the structures considered here. It is useful to note at this point that diamond, square, rectangular and triangle structures are all special cases of the general hexagonal shape so that the results for */S can be represented in one plot" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002002_012076-Figure11-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002002_012076-Figure11-1.png", "caption": "Figure 11. Stress on tire with polyethylene package", "texts": [ " The calculated forces like bump force were applied by the road to the tire of the wheel with a force of 2500N and the analysis is carried out in the no-slip condition as shown in figure 7. According to the simulation study mentioned, the analysis was carried out in ANSYS workbench and the following results were obtained. The individual analysis was carried out for the existing model and the proposed model with element size was 5mm to get the most efficient results in the stipulated amount of time. The Von-Mises stress, strain and deformation plots are as shown in figure 8, 9, 10 for air package as material for existing tire. Figure 11, 12 and 13 explain the von mises stress, strain and deformation plots for polyethylene as a material. ICMSMT 2020 IOP Conf. Series: Materials Science and Engineering 872 (2020) 012076 IOP Publishing doi:10.1088/1757-899X/872/1/012076 Case 1 \u2013 a) Stress plot b) Strain plot c) Total deformation plot ICMSMT 2020 IOP Conf. Series: Materials Science and Engineering 872 (2020) 012076 IOP Publishing doi:10.1088/1757-899X/872/1/012076 Case 2 \u2013 a) Stress plot b) Strain plot c) Total deformation plot The von-mises stress plot results seem to be accurate based on the convergence theory and compatibility equation which is decided based on the number of iterations carried out for each of the case [10, 11]" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002298_tec.2020.3045063-Figure10-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002298_tec.2020.3045063-Figure10-1.png", "caption": "Fig. 10. The temperature distribution of the turns of the rotor winding (longitudinal section is shown) along the axial rotor length depending on the profile of the sub-slot ducts with a 320 MW rated power: (a) variable; (b) smoothly tapering towards the rotor center to 17% of the inlet section.", "texts": [ " It must be mentioned that when air velocity in the radial ducts becomes significantly uniform, the non-uniformity of the rotor winding temperature distribution starts to be notably affected by the air heat up in the sub-slot duct (heat up value in the center of the rotor sub-slot duct is 14\u00b0\u0421). Thus, forced non-uniformity of air in the radial ducts can be used to offset the heat up and, as a consequence, further reduce the rotor winding temperature. This underlies the creation of the sub-slot duct smoothly tapering towards the rotor center to 17% (the best value found in this research) of the inlet section, Figure 9. Figure 10 shows the distribution of the rotor winding temperature along the axial rotor length in stepped sub-slot ducts and sub-slot ducts smoothly tapering towards the rotor center to 17% of the inlet section. Authorized licensed use limited to: UNIVERSITY OF CONNECTICUT. Downloaded on May 18,2021 at 02:32:19 UTC from IEEE Xplore. Restrictions apply. 0885-8969 (c) 2020 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001105_s00170-019-04688-w-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001105_s00170-019-04688-w-Figure7-1.png", "caption": "Fig. 7 Standard deformed model from ANSYS CFX", "texts": [ " 6 Matching result analysis and comparison columns in LQ need to be deleted and bQ should be updated with these fixed positions qi 0 i\u2208B\u00f0 \u00de. The starting guess about Q\u2032 is obtained by minimizing \u2016 LQq \u2032 \u2212 LQq\u2016 2 under the constraints of fixed positions qi 0 i\u2208B\u00f0 \u00de. The proposed method was applied to the remanufacturing of a twisted compressor blade with the blade tip damage which accords with most cases in blade repairing [27]. The fluid\u2013structure coupling calculation is used to simulate the deformation of the nominal model by the ANSYS CFX as shown in Fig. 7, since working deformation largely reflects the actual blade deformation under high-temperature and pressure working conditions. Therefore, this deformed model from the computational fluid dynamics (CFD) software tool, ANSYS CFX, is regarded as the deformed blade for comparison. As mentioned before, the damaged blade suffers from deformation and being worn. To make the input data as realistic as possible, we simulated the deformed blade data as the deposited blade by adding the welding-like part on the blade tip as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002175_icra40945.2020.9197401-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002175_icra40945.2020.9197401-Figure4-1.png", "caption": "Figure 4. (a) Components of the manipulator: i) inextensible nylon fiber as a reinforcement element, ii) unreinforced silicone actuator, iii) encapsulation layer, iv) SCS. (b) completed manipulator.", "texts": [ " Thanks to the modular design of the rigid links, the total rotation angle can be increased or decreased by changing the number of rigid links. All the rigid parts of the SCS and molds for silicone casting were designed using SolidWorks (Dassault Syst\u00e8mes, France) and fabricated by ProJet 3600 HD Max (3D Systems, USA) 3D printer using UV curable plastic (VisiJet M3 Crystal, 3D Systems, USA). The fabrication consists of four steps: i) fabrication of the FRAs; ii) preparation of the SCS and its encapsulation layer; iii) integration; iv) capping. After fabrication, the manipulator is properly assembled as shown in Fig. 4. A. FRA fabrication Each FRA is fabricated in 3 steps: i) first silicone layer fabrication; ii) fiber winding; iii) second silicone layer fabrication. The A and B parts of the DragonSkin 10 were mixed equally by weight and degassed to eliminate entrapped air bubbles. The mixture was poured into the first mold (Fig. 5a) for the first layer of the FRA. When the silicone is cured 697 Authorized licensed use limited to: University of New South Wales. Downloaded on September 20,2020 at 15:57:05 UTC from IEEE Xplore" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002276_j.neunet.2020.11.016-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002276_j.neunet.2020.11.016-Figure3-1.png", "caption": "Fig. 3. Schematic diagram of CartPole-v0.", "texts": [ " The dotted lines in Fig. 2 show the maximum of the trajectories in Example 1 will not exceed the boundaries we obtained. In addition, the figures of this example only show that Theorem 2 is better than Theorem 1 in this case. But in some other cases, Theorem 1 may be better than Theorem 2. In the next example, we will show the feasibility of Theorem 3. Example 2. In this example, we consider a classic CartPole-v0 experiment mentioned in some research (Zhang, Hu, Zhuo and Duan, 2018). As shown in Fig. 3, a trolley is equipped with a balance bar with a hinge, and The balance bar can rotate around the axis. Under the action of force F, the car moves to the left or right in a discrete time interval, thereby changing the position and bar of the car angle. The state of this model has 4 variables that need to be set. The position of the car on the track x, angle between the balance bar and the vertical direction \u03b8 , car speed v, balance bar angular velocity w. The agent will obtain the value of the next state through observation when the car moves every time step" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002766_j.procs.2021.06.071-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002766_j.procs.2021.06.071-Figure1-1.png", "caption": "Fig. 1. General view of the spider robot Fig. 2. Spider robot leg", "texts": [ " Currently, a huge variety of industrial robots with open kinematics similar to the \u201cKuka\u201d robots have found application in the industry [10]. The advantage of these robots is their relatively simple kinematic structure and good positioning accuracy. As a driving force, as a rule, electric motors are used, less often rotary pneumatic motors, for example [11]. Similar schemes of open kinematics are used in biorobots that copy the mode of movement of both humans [12] and insects [13, 14], in particular, spiders robotic (Fig. 1.). In these mechanisms, each leg has three degrees of freedom. Such insects are quite popular in the search for alternative movement mechanisms on difficult surfaces. Indeed, leg systems theoretically offer the potential for better rugged terrain than traditional wheeled or tracked designs, due to the low ground contact and lack of slippage. The use of robots with a large number of legs implies a better adaptation to uneven terrain, but at the same time a large number of legs inevitably lead to a more complex control system", " Currently, a huge variety of industrial robots with open kinematics similar to the \u201cKuka\u201d robots have found application in the industry [10]. The advantage of these robots is their relatively simple kinematic structure and good positioning accuracy. As a driving force, as a rule, electric motors are used, less often rotary pneumatic motors, for example [11]. Similar schemes of open kinematics are used in biorobots that copy the mode of movement of both humans [12] and insects [13, 14], in particular, spiders robotic (Fig. 1.). In these mechanisms, each leg has three degrees of freedom. Such insects are quite popular in the search for alternative movement mechanisms on difficult surfaces. Indeed, leg systems theoretically offer the potential for better rugged terrain than traditional wheeled or tracked designs, due to the low ground contact and lack of slippage. The use of robots with a large number of legs implies a better adaptation to uneven terrain, but at the same time a large number of legs inevitably lead to a more complex control system", "\ud835\udefc2 = \ud835\udfcf\ud835\udfcf \ud835\udfd0\ud835\udfd0 (\ud835\udf46\ud835\udf46\ud835\udfcf\ud835\udfcf \ud835\udfd0\ud835\udfd0 \u2219 \ud835\udc8e\ud835\udc8e\ud835\udfcf\ud835\udfcf + \ud835\udc71\ud835\udc71\ud835\udc68\ud835\udc68\ud835\udfcf\ud835\udfcf + \ud835\udc71\ud835\udc71\ud835\udc88\ud835\udc88\ud835\udfcf\ud835\udfcf)?\u0307?\ud835\udf36\ud835\udfd0\ud835\udfd0; (6) \ud835\udc7b\ud835\udc7b(\ud835\udfd0\ud835\udfd0) = \ud835\udc49\ud835\udc49\ud835\udc34\ud835\udc342 2 \u2219\ud835\udc5a\ud835\udc5a2 2 + \ud835\udc3d\ud835\udc3d\ud835\udc34\ud835\udc342 2 (?\u0307?\ud835\udefc2 + ?\u0307?\ud835\udefd2) + \ud835\udc3d\ud835\udc3d\ud835\udc54\ud835\udc542 2 ?\u0307?\ud835\udefd2. The speed of the point \ud835\udc34\ud835\udc342 is a function of the generalized coordinates \u03b1 and \ud835\udefd\ud835\udefd, thus \ud835\udc49\ud835\udc49\ud835\udc34\ud835\udc342 = \ud835\udc49\ud835\udc49\ud835\udc34\ud835\udc342(\ud835\udefc\ud835\udefc, \ud835\udefd\ud835\udefd), therefore: \ud835\udc49\ud835\udc49\ud835\udc34\ud835\udc342 2 = (\ud835\udc49\ud835\udc49\ud835\udc34\ud835\udc342\ud835\udefc\ud835\udefc + \ud835\udc49\ud835\udc49\ud835\udc34\ud835\udc342\ud835\udefd\ud835\udefd)2 = (\ud835\udc3f\ud835\udc3f1?\u0307?\ud835\udefc + \ud835\udf0c\ud835\udf0c2?\u0307?\ud835\udefd)2 = \ud835\udc3f\ud835\udc3f1 2?\u0307?\ud835\udefc2 + \ud835\udf0c\ud835\udf0c2 2?\u0307?\ud835\udefd2 + \ud835\udc3f\ud835\udc3f1?\u0307?\ud835\udefc \u2219 \ud835\udf0c\ud835\udf0c2?\u0307?\ud835\udefd \u2219 \ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc50(90) = \ud835\udc3f\ud835\udc3f1 2?\u0307?\ud835\udefc2 + \ud835\udf0c\ud835\udf0c2 2?\u0307?\ud835\udefd2 S.Yu. Misyurin et al. / Procedia Computer Science 190 (2021) 604\u2013610 607 Author name / Procedia Computer Science 00 (2019) 000\u2013000 3 Fig. 1. General view of the spider robot Fig. 2. Spider robot leg Let's go to the kinematic description of the robot's leg (Fig. 3). \ud835\udc42\ud835\udc421, \ud835\udc42\ud835\udc422, \ud835\udc42\ud835\udc423 are rotary kinematic pairs with angles of rotation. The \ud835\udc42\ud835\udc421\ud835\udc42\ud835\udc422 link rotates in the \ud835\udc42\ud835\udc421\ud835\udc65\ud835\udc65 plane, \ud835\udefc\ud835\udefc is the angle of the link deviation from the \ud835\udc42\ud835\udc421\ud835\udc65\ud835\udc65 axis. The axis of rotation of the kinematic pairs \ud835\udc42\ud835\udc422, \ud835\udc42\ud835\udc423 are parallel to the \ud835\udc42\ud835\udc421\ud835\udc65\ud835\udc65 plane. We introduce the following notation: |\ud835\udc42\ud835\udc421\ud835\udc42\ud835\udc422| = \ud835\udc59\ud835\udc591, |\ud835\udc42\ud835\udc422\ud835\udc42\ud835\udc423| = \ud835\udc59\ud835\udc592, |\ud835\udc42\ud835\udc423\ud835\udc42\ud835\udc424| = \ud835\udc59\ud835\udc593, |\ud835\udc42\ud835\udc421\ud835\udc34\ud835\udc341| = \ud835\udf0c\ud835\udf0c1, |\ud835\udc42\ud835\udc422\ud835\udc34\ud835\udc342| = \ud835\udf0c\ud835\udf0c2, |\ud835\udc42\ud835\udc423\ud835\udc34\ud835\udc343| = \ud835\udf0c\ud835\udf0c3; \ud835\udc3d\ud835\udc3d\ud835\udc34\ud835\udc341, \ud835\udc3d\ud835\udc3d\ud835\udc34\ud835\udc342, \ud835\udc3d\ud835\udc3d\ud835\udc34\ud835\udc343 are the moments of inertia of links 1, 2 and 3 relative to their centers of mass; \ud835\udc5a\ud835\udc5a1,2,3 are masses of these links; \ud835\udf0c\ud835\udf0c1,2 are the coordinates of their centers of mass relative to the kinematic pairs; \ud835\udc3d\ud835\udc3d\ud835\udc51\ud835\udc511, \ud835\udc3d\ud835\udc3d\ud835\udc51\ud835\udc512, \ud835\udc3d\ud835\udc3d\ud835\udc51\ud835\udc513 are the moments of inertia of the engine rotors, installed in the joints of the links (if the movement from the actuator to the link is transmitted through the reducer, then the moment of inertia of the is substituted into the equation given to the output shaft of the gearbox); \ud835\udefc\ud835\udefc, \ud835\udefd\ud835\udefd and \ud835\udefe\ud835\udefe are generalized coordinates \u2013 the angles of rotation of the links, calculated according to the scheme in figure 3" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001934_j.mechmachtheory.2020.103895-Figure10-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001934_j.mechmachtheory.2020.103895-Figure10-1.png", "caption": "Fig. 10. Kinematically equivalent leg of the Ragnar robot.", "texts": [ " The Ansys\u00a9 model A1 behaves differently than expected when a load is applied to the end-effector along the y -axis. This behavior is due to a rigid motion as it is signaled by a warning message. To better understand this problem, a second Ansys\u00a9 model, referred to as A2, has been implemented. The model A2 refers to a kinematically equivalent solution in which M is rigidly connected using bonded contacts to the lower links at the point C and the two lower revolute joints are replaced by two universal joints (U) as shown in Fig. 10 . The solution A2 allows using the MPC formulation, not available in the first solution A1 in which the augmented Lagrangian formulation has been implemented. In the same model A2, the constraint type target normal couples the displacements to the rotations of target and contact bodies, here respectively denoted by the lower links and M . To perform the comparison, the corresponding fully flexible model has been developed using the SHELL 9WDR. Applying the same load-cases, the results reported in Table 7 have been obtained" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001118_tsmc.2019.2960803-Figure8-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001118_tsmc.2019.2960803-Figure8-1.png", "caption": "Fig. 8. Chaotic behavior of fuzzy model-based Lorenz system.", "texts": [ "4, \u03b11 = \u03b12 = 0.01, and \u03b21 = \u03b22 = 0.001 are obtained as follows: K1 = [\u221278.5441 \u2212 10.0438 \u2212 0.0140 ] K2 = [\u221295.7668 \u2212 10.0869 \u2212 0.0330 ] and L\u0302 = 104 \u00d7 [ 1.6773 \u2212 1.0411 0.1738 \u22121.0266 0.6316 \u2212 0.0781 ] . For the simulation purposes, the membership functions are taken as \u22071 i (\u03bei(t)) = (1/2)(1 + [(yi1(t))/25]) and \u22072 i (\u03bei(t)) = 1 \u2212 \u22071 i (\u03bei(t)) for i = 1, 2, 3, 4. Moreover, the external disturbances wi(t) and \u03b7i(t) are assumed the same as in the previous example. With the parameters mentioned above, Fig. 8 gives the phase portrait graph of the considered fuzzy model-based Lorenz system to realize its chaotic behavior. Based on the designed control strategy, the trajectories of the synchronization states are presented in Figs. 9\u201311, wherein a satisfactory synchronization performance is achieved. This event happened only because of the perfect disturbance estimation of unknown disturbance di(t). To illustrate the efficiency of the proposed anti-disturbance control strategy, the evolutions of ||ei1(t)|| and ||ei2(t)|| are in depicted Figs" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001066_1.a34565-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001066_1.a34565-Figure3-1.png", "caption": "Fig. 3 CTM profile.", "texts": [ " If there is no tipmounted instrument, the mechanism can be retained or jettisoned, depending on debris requirements. If the mechanism is jettisoned, an undamped mechanism may be sufficient. The opening shock will be minimal because only a small portion of the boom is traveling at the end of deployment. The collapsible tube mast (CTM) boom was selected for use with this mechanism. CTM booms have excellent torsional stiffness and axial strength [10]. In addition, North Carolina State University (NCSU) has extensive experience manufacturing these booms [11]. The CTM profile is shown in Fig. 3. The booms used in this study were manufactured in the Ballistic Loading and Structural Test (BLAST) Laboratory atNCSUusing 3K spread tow HTS40 carbon fiber from Composite Envisions (3000 fibers/tow). Booms were manufactured in a single cure cycle using Fibre Glast\u2019s 2000-series epoxy with 2060 hardener [11]. The boom geometry and additional layup details are provided in Table 1. The MUSE mechanism uses a rotary damper to limit deployment speed. An ideal damper for this application would maintain deployment speed at exactly the desired rate", " The energy stored by bending a flattened boom is found by integrating the moment through the displaced angle: Vbending Z Mbending d\u03b8 (16) The curvature is inversely proportional to radius: \u03ba 1 r \u03b8 s (17) Equations (16) and (17) can be combined to express bending energy in terms of curvature: Vbending s Z Mbending d\u03ba (18) To measure the strain energy stored by bending, a boom sample with a 15.5 mm gauge length was placed in a balanced bending fixture [13]. The boom has a flattened width of 65 mm, which is too wide for the fixture. To fit in the fixture, the boom sample was cut in half down the linemarked r1 in Fig. 3 to a flattenedwidth of 32.5mm. The half-boom sample was bent until it reached a curvature of 26.2 m\u22121. This corresponds to themaximumcurvature of the deploymentmechanism: the inner wrap that conforms to the 38.1mm radius of the hub. The test fixture is shown in Fig. 10. The load cell\u2019s vertical displacement can be used to find the fixture angle \u03b8 and the moment arm x: 0 20 40 60 80 100 Pinion speed, rad/s 0 5 10 15 20 25 T or qu e, N /m m Centrifugal Viscous Fig. 8 Damper torques. D ow nl oa de d by M A C Q U A R IE U N IV E R SI T Y o n D ec em be r 6, 2 01 9 | h ttp :// ar c" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001406_ur49135.2020.9144853-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001406_ur49135.2020.9144853-Figure2-1.png", "caption": "Fig. 2. Comparison of two grasping methods for paper gripper.", "texts": [ " The rest of this paper is organized as follows: Section II describes the basic principle and structure design of the paper grippers. Experimental validations and discussions based on the experiment result are presented in Section III and Section IV. Conclusions and future work are talked in Section V. Most of the robotic grippers have similar structures of stationary palm cases and movable multilink fingers, and they perform the grasping movement in a bio-inspired way by closing the finger tips\u2019 distance [3]. However, still two different grasping modes can be observed for design of paper-made grippers. Fig. 2(a) presents a force-closure grasping mechanism. The gripper uses a human-like picking movement with finger tips for fixation of the grasping objects, in which the friction forces between the gripper and object help to maintain the grasping result. Although the gripper may have deformable structures for grasping adaptation, the contacting condition for the gripper fingers and the objects will still be small regions on the finger tips, so that the objects posture can be adjusted precisely by controlling the finger tips movements. However, large grasping force will be required for a secured grasping as compromise, which may cause possible damage to the grasping targets as a result. Things can be even worse when the grasping targets have delicate structures on their surface, such as cream donuts and rice cakes. Fig. 2(b) presents a shape-closure grasping mechanism. The gripper is more suitable for the special soft food objects, in which the gripper fingers will be inserted bellow the objects to provide the support, and the objects will be enclosed with a confined gripper-shape for fixation. On the one hand, the objects are carried from the bottom surface, avoiding the delicate top structures; on the other hand, the objects are wrapped around with the multiple fingers, which will be a more secured holding posture during the grasping process" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000718_978-1-4471-7278-9-Figure6.4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000718_978-1-4471-7278-9-Figure6.4-1.png", "caption": "Fig. 6.4 The scaffolding structure for printing a ball and the Mickey Mouse ear-shaped rafts", "texts": [], "surrounding_texts": [ "Scaffolding is a temporary structure used to support a work crew and materials to aid in the construction, maintenance, and repair of buildings, bridges, and other manmade structures. The use of scaffolding can be traced back as far as 17,000 years. Archeologists found sockets in the walls of Paleolithic cave paintings at Lascaux, suggesting that a scaffolding system was used for painting the ceiling. As a tool, scaffolding has unique characteristics. It provides support for accessing a work area without damaging it, and it can be made out of inexpensive materials. After work is complete, the scaffolding is removed, leaving virtually no trace on the finished product. The idea is to provide a temporary support for the tool-making or problemsolving processes and the support can be removed afterwards." ] }, { "image_filename": "designv11_5_0001089_icasert.2019.8934748-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001089_icasert.2019.8934748-Figure4-1.png", "caption": "Fig. 4. Servo Motor", "texts": [ " The body parts made by an aluminum sheet. We utilized the legs for the development of the robot. Later on, we will utilize legs for strolling the robot. 1) Microcontroller: The microcontroller shows in Fig.3. It controls any fringe device. like servo motor, DC motor, Ultrasonic sound Sensors and so on. We control here 18 Servo motors, 2 Ultrasonic sound Sensors, and some LED. 2) Servo motor: Servo motor to a blunder detecting input control which is utilized to remedy the execution of a framework. Fig. 4 shows the picture of the servo motor. Servo or RC Servo Motors are DC machine furnished with a servo system for exact control of the precise position. The RC servo engines, as a rule, have a revolution confine from 90\u00b0 to 180\u00b0. A few servos additionally have turn farthest point of at least 360\u00b0. In any case, servos don\u2019t turn constantly. Their pivot is limited in the middle of the settled points. We are utilizing servo Motor here for control robot hands. 3) Camera: The camera shows in Fig. 5 which is a picture sensor gadget" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002197_tase.2020.3024033-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002197_tase.2020.3024033-Figure6-1.png", "caption": "Fig. 6. IMU-MK sensor design.", "texts": [ " The proposed sensor was designed to be truly affordable, compact, and easy-to-use. The IMU-MK sensor is based on one MPU-6050 (0.9e, Invensense) IMU embedding a 3-D gyroscope and a 3-D accelerometer. A microcontroller (Atmel ATmega328, 2.4e) was associated with each IMU to collected its data and transmit it to a master microcontroller sending all three IMU data to a computer at a maximum rate of 150Hz. A 4 \u00d7 4 cm AR marker was fixed on the top of the IMU enclosure, which was 3-D printed with dimensions of 4 \u00d7 4 \u00d7 2 cm and a weight of 37 g (see Fig. 6). Velcro strap was used to ensure that the sensor remains firmly attached to the body segment. Moreover, a standard 60 Hz high-definition Authorized licensed use limited to: Carleton University. Downloaded on November 04,2020 at 05:19:35 UTC from IEEE Xplore. Restrictions apply. RGB USB camera (30e, ELP, USBFHD01) was set up on a tripod to be able to visualize the whole scene and to detect all three AR markers (see Fig. 5). A custom software was written in C++ to collect and record synchronously all data at a frequency of 54 Hz" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002847_j.mechmachtheory.2021.104521-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002847_j.mechmachtheory.2021.104521-Figure4-1.png", "caption": "Fig. 4. Skeleton representation of the 3D Cartesian flexible hinges: (a) Filleted design (FD); (b) Non-filleted design (ND).", "texts": [ " 3 with respect to the fixed global frame Oxyz by combining the known, closed-form compliances of the straight-axis segments and of the quarter-circle, circular-axis segments. The 7 segments of the filleted hinge design (referred to by the acronym FD) of Fig. 3(a) are connected in series and so are the 5 segments composing the non-filleted hinge design (the acronym ND denotes this configuration) of Fig. 3(b). As a result, the full-hinge compliance matrices of these hinges are determined by adding the individual-segment compliance matrices after appropriate transformations from the segments local frames to the global frame. Fig. 4 identifies the local frames positions for all component segments of the FD and ND hinges. The local-frame compliance matrix[C(i) Oi ]of a generic segment i (i = 1, 2, \u2026, n where n = 7 for FD hinges and n = 5 for ND hinges) is: [ C(i) Oi ] 6\u00d76 = \u23a1 \u23a3 [ C(i) Oi ,ip ] 3\u00d73 [0]3\u00d73 [0]3\u00d73 [ C(i) Oi ,op ] 3\u00d73 \u23a4 \u23a6 (1) The compliance submatrix[C(i) Oi ,ip] is defined based on the in-plane displacement vector[u(i) Oi ,ip] = [ u(i) xi u(i) yi \u03b8(i) zi ]T , while the sub- matrix[C(i) Oi ,op] is formulated with respect to the out-of-plane displacement vector[u(i) Oi ,op] = [ \u03b8(i) xi \u03b8(i) yi u(i) zi ]T where u represents a N", " (5) for a segment of radius ri and with constant cross-section of diameter d are: \u23a7 \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23a8 \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23a9 C(i) ux \u2212 fx = (3\u03c0 \u2212 8)r3 i 4EI + \u03c0ri 4EA + \u03c0\u03b1sri 4GA ; C(i) ux \u2212 fy = \u2212 r3 i 2EI + ri 2EA ; C(i) ux \u2212 mz = (\u03c0/2 \u2212 1)r2 i EI ; C(i) uy \u2212 fy = \u03c0r3 i 4EI + \u03c0ri 4EA + \u03c0\u03b1sri 4GA ; C(i) uy \u2212 mz = \u2212 r2 i EI ; C(i) \u03b8z \u2212 mz = \u03c0ri 2EI ; C(i) \u03b8x \u2212 mx = C(i) \u03b8y \u2212 my = \u03c0ri 4EI + \u03c0ri 4GIp ; C(i) \u03b8x \u2212 my = \u2212 ri 2EI + ri 2GIp ; C(i) \u03b8x \u2212 fz = \u2212 \u03c0r2 i 4EI + (2 \u2212 \u03c0/2)r2 i 2GIp ; C(i) \u03b8y \u2212 fz = r2 i 2EI + r2 i 2GIp ; C(i) uz \u2212 fz = \u03c0r3 i 4EI + (3\u03c0 \u2212 8)r3 i 4GIp + \u03c0\u03b1sri 2GA (6) The local-frame compliance matrices of Eq. (1) are transferred to the hinge global frame Oxyz, which is placed at the end O in Fig. 4, by means of transformation matrices [Tr(i)] as: [ C(i) O ] = [ Tr(i) ]T [ C(i) Oi ][ Tr(i) ] with [ Tr(i) ] = [ R(i)] [ T(i) OOi ] (7) As discussed in [18], the translation matrix of Eq. (7) is calculated based on the offsets \u0394xi, \u0394yi, \u0394zi, which are distances from O to Oi along the global directions x, y and z, namely: [ T (i) OOi ] = \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 1 0 0 0 0 0 0 1 0 0 0 0 \u0394yi \u2212 \u0394xi 1 0 0 0 0 \u0394zi 0 1 0 \u2212 \u0394yi \u2212 \u0394zi 0 0 0 1 \u0394xi 0 0 0 0 0 1 \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 (8) The offsets of segment 12 of the ND configuration are all zero. The other segments offsets are expressed based on the dimensions shown in Fig. 4 as: \u0394x1 = r1; \u0394y1 = 0;\u0394z1 = 0;\u0394x2 = \u0394x1; \u0394y2 = 0;\u0394z2 = 0;\u0394x3 = \u0394x2 + l2; \u0394y3 = 0;\u0394z3 = 0; \u0394x4 = \u0394x3 + r3; \u0394y4 = r3;\u0394z4 = 0; \u0394x5 = \u0394x4; \u0394y5 = \u0394y4 + l4;\u0394z5 = 0; \u0394x6 = \u0394x67 = \u0394x4; \u0394y6 = \u0394y67 = \u0394y5 + r5;\u0394z6 = r5; \u0394x7 = \u0394x4; \u0394y7 = \u0394y6;\u0394z7 = \u0394z6 + l6 The rotation matrix of Eq. (7) is the product of three rotation matrices: [ R(i)] = [ R(i) \u03c8i ][ R(i) \u03b8i ][ R(i) \u03c6i ] (9) which correspond to a three-rotation sequence resulting in the local frame orientation at Oi from the global frame orientation at O. The first rotation of angle \u03c6i is around the global z axis, the second rotation of angle \u03b8i is performed around the resulting x axis, and the third rotation of angle \u03c8 i is around the new z axis, which results after the second rotation", " 3(a), the two filleted end segments were meshed with tapered beam elements while the other (5) segments were discretized with regular, constant cross-section beam elements. Several FE tests were performed to identify a mesh structure that is both computationally economical and accurate. The final mesh was formed of 15 elements for the circular-axis portions, 20 elements for the constant cross-section, straight-axis segments, and of 10 tapered beam elements for the shorter end filleted segments of the FD hinges. To determine the FE compliances of the 6 \u00d7 6 matrix [CO] of Eq. (11), unit loads (forces and moments) were applied, one at a time, at the free end O in Fig. 4, while the opposite end is fixed. For each individual load, 6 displacements (3 translations: ux, uy, uz and 3 rotations: \u03b8x, \u03b8y, \u03b8z) are obtained. These 6 displacements are actually the 6 compliances on the row of [CO] that corresponds to the specific force (f) or moment (m) in the load vector [f] = [fx, fy, mz, mx, my, fz]T. We have used the following material properties: Young\u2019s modulus E = 1.7 109 N/m2, Poisson\u2019s ratio \u03bc = 0.3, and shear modulus G = E/ [2(1 + \u03bc)]. Several sets of geometrical parameters were utilized to run both FE simulation and analytical compliance calculation for the ND and N", " The finite element results matched very closely the analytical model predictions for all the individual compliances in [CO] and Table 2 lists a sample of those results together with the relative errors of the two models \u2013 the errors were all less than 1%. Two tripod mechanisms, as the ones shown in the photographs of Fig. 6, were 3D printed from plastic material (a variety of Stratasys RDG131 Digital ABS). The dimensions defining the ND configuration of Fig. 6(b) and skeleton representation of Fig. 4(b) are: d = 0.003 m, l12 = l4 = l67 = 0.02 m, r3 = r5 = 0.004 m. The FD design pictured in Fig. 6(a) and whose geometry is shown in Fig. 4(a) is defined by d = 0.003 m, l2 = l4 = l6 = 0.02 m, r3 = r5 = 0.004 m, r1 = r7 = 0.006 m. For both designs the hinge root radius (see the mechanism of Fig. 5) is R = 0.03 m. The experimental setup, which is pictured in Fig. 7, was designed to measure the vertical (z-axis) displacement uz at the rigid platform center when a downward z-axis force fz is applied at the same point. With the aid of uz and fz, the corresponding experimental stiffness is calculated asKfz \u2212 uz = fz/uz. The force was generated by an electromagnetic actuator (Thorlabs VC625M) and the resulting translation was measured by a laser displacement sensor (\u03bcEpsilon NCDT ILR 1320-25)" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000489_s00170-016-9165-4-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000489_s00170-016-9165-4-Figure6-1.png", "caption": "Fig. 6 a Experimental setup of temperature rise caused by friction force and b the measured points and channel numbers of thermocouples on the LM guide system", "texts": [ " The measurement of the friction force was conducted using the experimental procedure presented in Section 2.3; the predicted friction force was calculated using Eq. (4). Figure 5a, b compares the predicted and experimental results of the friction force on the LMB, respectively. The ideal estimate line shows conditions in which the results predicted by the BBD model are in good agreement with the experimental results. The prediction equations show differences within 5 and 9 % for the initial BBD and additional experiments, respectively. Figure 6 shows the experimental apparatus used to measure temperature distributions in a LM guide system. Thermocouples and a data logger were added to the original experimental setup (Fig. 1) to acquire temperature data. Ten thermocouples were positioned along the LM guide system; six thermocouples were attached to the bearing block, and three were attached to the rail, as shown in Fig. 6. One thermocouple was used to measure ambient temperature. Four load/velocity conditions were considered for the temperature experiments: (i) 300 kgf and 100 mm/s, (ii) 300 kgf and 150 mm/s, (iii) 300 kgf and 200 mm/s, and (iv) 500 kgf and 100 mm/s. A heavy preload was applied in all cases. During the experiment, the LMB had a reciprocated stroke of 400 mm until the temperature reached a steady state. Figure 7a shows temperature variation over time for the third experimental case, which had the maximum temperature rise" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002766_j.procs.2021.06.071-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002766_j.procs.2021.06.071-Figure3-1.png", "caption": "Fig. 3. A spider robot leg kinematic scheme.", "texts": [ " This mount is a rotational pair in the plane of the \u201cspider\u201d body. The \ud835\udc42\ud835\udc421\ud835\udc67\ud835\udc67 axis is directed vertically up along the rotation axis of the actuator. The \ud835\udc42\ud835\udc421\ud835\udc65\ud835\udc65 axis is directed perpendicular to the \ud835\udc42\ud835\udc421\ud835\udc67\ud835\udc67 axis, and is directed to the center of mass of the spider's body; the \ud835\udc42\ud835\udc421\ud835\udc66\ud835\udc66 axis is perpendicular to the \ud835\udc42\ud835\udc421\ud835\udc67\ud835\udc67\ud835\udc65\ud835\udc65 plane and forms the right coordinate system. 606 S.Yu. Misyurin et al. / Procedia Computer Science 190 (2021) 604\u2013610 Author name / Procedia Computer Science 00 (2019) 000\u2013000 3 Let's go to the kinematic description of the robot's leg (Fig. 3). \ud835\udc42\ud835\udc421, \ud835\udc42\ud835\udc422, \ud835\udc42\ud835\udc423 are rotary kinematic pairs with angles of rotation. The \ud835\udc42\ud835\udc421\ud835\udc42\ud835\udc422 link rotates in the \ud835\udc42\ud835\udc421\ud835\udc65\ud835\udc65 plane, \ud835\udefc\ud835\udefc is the angle of the link deviation from the \ud835\udc42\ud835\udc421\ud835\udc65\ud835\udc65 axis. The axis of rotation of the kinematic pairs \ud835\udc42\ud835\udc422, \ud835\udc42\ud835\udc423 are parallel to the \ud835\udc42\ud835\udc421\ud835\udc65\ud835\udc65 plane. We introduce the following notation: |\ud835\udc42\ud835\udc421\ud835\udc42\ud835\udc422| = \ud835\udc59\ud835\udc591, |\ud835\udc42\ud835\udc422\ud835\udc42\ud835\udc423| = \ud835\udc59\ud835\udc592, |\ud835\udc42\ud835\udc423\ud835\udc42\ud835\udc424| = \ud835\udc59\ud835\udc593, |\ud835\udc42\ud835\udc421\ud835\udc34\ud835\udc341| = \ud835\udf0c\ud835\udf0c1, |\ud835\udc42\ud835\udc422\ud835\udc34\ud835\udc342| = \ud835\udf0c\ud835\udf0c2, |\ud835\udc42\ud835\udc423\ud835\udc34\ud835\udc343| = \ud835\udf0c\ud835\udf0c3; \ud835\udc3d\ud835\udc3d\ud835\udc34\ud835\udc341, \ud835\udc3d\ud835\udc3d\ud835\udc34\ud835\udc342, \ud835\udc3d\ud835\udc3d\ud835\udc34\ud835\udc343 are the moments of inertia of links 1, 2 and 3 relative to their centers of mass; \ud835\udc5a\ud835\udc5a1,2,3 are masses of these links; \ud835\udf0c\ud835\udf0c1,2 are the coordinates of their centers of mass relative to the kinematic pairs; \ud835\udc3d\ud835\udc3d\ud835\udc51\ud835\udc511, \ud835\udc3d\ud835\udc3d\ud835\udc51\ud835\udc512, \ud835\udc3d\ud835\udc3d\ud835\udc51\ud835\udc513 are the moments of inertia of the engine rotors, installed in the joints of the links (if the movement from the actuator to the link is transmitted through the reducer, then the moment of inertia of the is substituted into the equation given to the output shaft of the gearbox); \ud835\udefc\ud835\udefc, \ud835\udefd\ud835\udefd and \ud835\udefe\ud835\udefe are generalized coordinates \u2013 the angles of rotation of the links, calculated according to the scheme in figure 3. \ud835\udc40\ud835\udc40\ud835\udc51\ud835\udc511, \ud835\udc40\ud835\udc40\ud835\udc51\ud835\udc512, \ud835\udc40\ud835\udc40\ud835\udc51\ud835\udc513 are the moments of engines and \ud835\udc40\ud835\udc40\ud835\udc50\ud835\udc501, \ud835\udc40\ud835\udc40\ud835\udc50\ud835\udc502, \ud835\udc40\ud835\udc40\ud835\udc50\ud835\udc503 are the moments of resistances. We introduce the notation for the lengths |\ud835\udc42\ud835\udc421\ud835\udc34\ud835\udc342| = \ud835\udc3f\ud835\udc3f1, |\ud835\udc42\ud835\udc421\ud835\udc34\ud835\udc343| = \ud835\udc3f\ud835\udc3f2, |\ud835\udc42\ud835\udc422\ud835\udc34\ud835\udc343| = \ud835\udc3f\ud835\udc3f3. When solving problems of kinematics and dynamics of a robot with six legs, we rely on the fact that part of the legs moves in the air, while part of the legs are on the ground and supports the robot. It is necessary to find out how each leg moves separately, which movement will be the most economical or fast", " The speed of the point \ud835\udc34\ud835\udc342 is a function of the generalized coordinates \u03b1 and \ud835\udefd\ud835\udefd, thus \ud835\udc49\ud835\udc49\ud835\udc34\ud835\udc342 = \ud835\udc49\ud835\udc49\ud835\udc34\ud835\udc342(\ud835\udefc\ud835\udefc, \ud835\udefd\ud835\udefd), therefore: \ud835\udc49\ud835\udc49\ud835\udc34\ud835\udc342 2 = (\ud835\udc49\ud835\udc49\ud835\udc34\ud835\udc342\ud835\udefc\ud835\udefc + \ud835\udc49\ud835\udc49\ud835\udc34\ud835\udc342\ud835\udefd\ud835\udefd)2 = (\ud835\udc3f\ud835\udc3f1?\u0307?\ud835\udefc + \ud835\udf0c\ud835\udf0c2?\u0307?\ud835\udefd)2 = \ud835\udc3f\ud835\udc3f1 2?\u0307?\ud835\udefc2 + \ud835\udf0c\ud835\udf0c2 2?\u0307?\ud835\udefd2 + \ud835\udc3f\ud835\udc3f1?\u0307?\ud835\udefc \u2219 \ud835\udf0c\ud835\udf0c2?\u0307?\ud835\udefd \u2219 \ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc50(90) = \ud835\udc3f\ud835\udc3f1 2?\u0307?\ud835\udefc2 + \ud835\udf0c\ud835\udf0c2 2?\u0307?\ud835\udefd2 S.Yu. Misyurin et al. / Procedia Computer Science 190 (2021) 604\u2013610 607 Author name / Procedia Computer Science 00 (2019) 000\u2013000 3 Fig. 1. General view of the spider robot Fig. 2. Spider robot leg Let's go to the kinematic description of the robot's leg (Fig. 3). \ud835\udc42\ud835\udc421, \ud835\udc42\ud835\udc422, \ud835\udc42\ud835\udc423 are rotary kinematic pairs with angles of rotation. The \ud835\udc42\ud835\udc421\ud835\udc42\ud835\udc422 link rotates in the \ud835\udc42\ud835\udc421\ud835\udc65\ud835\udc65 plane, \ud835\udefc\ud835\udefc is the angle of the link deviation from the \ud835\udc42\ud835\udc421\ud835\udc65\ud835\udc65 axis. The axis of rotation of the kinematic pairs \ud835\udc42\ud835\udc422, \ud835\udc42\ud835\udc423 are parallel to the \ud835\udc42\ud835\udc421\ud835\udc65\ud835\udc65 plane. We introduce the following notation: |\ud835\udc42\ud835\udc421\ud835\udc42\ud835\udc422| = \ud835\udc59\ud835\udc591, |\ud835\udc42\ud835\udc422\ud835\udc42\ud835\udc423| = \ud835\udc59\ud835\udc592, |\ud835\udc42\ud835\udc423\ud835\udc42\ud835\udc424| = \ud835\udc59\ud835\udc593, |\ud835\udc42\ud835\udc421\ud835\udc34\ud835\udc341| = \ud835\udf0c\ud835\udf0c1, |\ud835\udc42\ud835\udc422\ud835\udc34\ud835\udc342| = \ud835\udf0c\ud835\udf0c2, |\ud835\udc42\ud835\udc423\ud835\udc34\ud835\udc343| = \ud835\udf0c\ud835\udf0c3; \ud835\udc3d\ud835\udc3d\ud835\udc34\ud835\udc341, \ud835\udc3d\ud835\udc3d\ud835\udc34\ud835\udc342, \ud835\udc3d\ud835\udc3d\ud835\udc34\ud835\udc343 are the moments of inertia of links 1, 2 and 3 relative to their centers of mass; \ud835\udc5a\ud835\udc5a1,2,3 are masses of these links; \ud835\udf0c\ud835\udf0c1,2 are the coordinates of their centers of mass relative to the kinematic pairs; \ud835\udc3d\ud835\udc3d\ud835\udc51\ud835\udc511, \ud835\udc3d\ud835\udc3d\ud835\udc51\ud835\udc512, \ud835\udc3d\ud835\udc3d\ud835\udc51\ud835\udc513 are the moments of inertia of the engine rotors, installed in the joints of the links (if the movement from the actuator to the link is transmitted through the reducer, then the moment of inertia of the is substituted into the equation given to the output shaft of the gearbox); \ud835\udefc\ud835\udefc, \ud835\udefd\ud835\udefd and \ud835\udefe\ud835\udefe are generalized coordinates \u2013 the angles of rotation of the links, calculated according to the scheme in figure 3. \ud835\udc40\ud835\udc40\ud835\udc51\ud835\udc511, \ud835\udc40\ud835\udc40\ud835\udc51\ud835\udc512, \ud835\udc40\ud835\udc40\ud835\udc51\ud835\udc513 are the moments of engines and \ud835\udc40\ud835\udc40\ud835\udc50\ud835\udc501, \ud835\udc40\ud835\udc40\ud835\udc50\ud835\udc502, \ud835\udc40\ud835\udc40\ud835\udc50\ud835\udc503 are the moments of resistances. We introduce the notation for the lengths |\ud835\udc42\ud835\udc421\ud835\udc34\ud835\udc342| = \ud835\udc3f\ud835\udc3f1, |\ud835\udc42\ud835\udc421\ud835\udc34\ud835\udc343| = \ud835\udc3f\ud835\udc3f2, |\ud835\udc42\ud835\udc422\ud835\udc34\ud835\udc343| = \ud835\udc3f\ud835\udc3f3. The leg has three degrees of freedom and moves with the help of kinematic pairs \ud835\udc42\ud835\udc421, \ud835\udc42\ud835\udc422, \ud835\udc42\ud835\udc423. Fig. 3. A spider robot leg kinematic scheme. When solving problems of kinematics and dynamics of a robot with six legs, we rely on the fact that part of the legs moves in the air, while part of the legs are on the ground and supports the robot. It is necessary to find out how each leg moves separately, which movement will be the most economical or fast. In other words, at the first stage, we assume that the robot's body is stationary. According to this assumption, we consider each leg as an open kinematic chain" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001853_tia.2020.2983903-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001853_tia.2020.2983903-Figure2-1.png", "caption": "Fig. 2. Conceptual 1/4 rotor structure of a 3-phase, 4-pole, direct-on-line synchronous reluctance motor by inserting conductors into the barriers.", "texts": [ "r r ds qsL L Therefore, at their corresponding currents of r dsi and ,r qsi from the steady-state electromagnetic torque (Te) of a P-pole SynRM with: (3 4 )( ) (3 4)( ) ,r r r r r r r r e ds qs qs ds ds qs qs dsT P i i P L L i i (1) it is obvious that the torque will also be reduced and hence the desired high-efficiency objective may not be satisfied if such structural arrangements for rotor conductors are adopted. Apparently, to equip line-start capability to the SynRM by inserting proper conductors into its rotor, the key point is not to affect the original optimally designed rotor flux paths. Consequently, to maintain the original designed structure of the punched rotor whereas certain amounts of conductors are desired to insert axially, it is evident that the only available spaces will be the rotor flux barriers. As conceptually illustrated in Fig. 2 [9] and compared with Fig. 1, it can be seen that it is flexible to adjust the conductor sizes along the existing flux barriers. Therefore, the coupled fluxes and resistance in the equivalent enclosed conductor loops can be adjusted, as well as the start-up and dynamic stabilization characteristics of the SynRM with conductors inserted into its rotor barriers can then be accordingly controlled. As the current-circulation loops in the qr- and dr-axes are respectively established by the two conductors that are symmetrically apart to their corresponding axes, the equivalent circuits that can characterize the operational behaviors of this DOL_SynRM at the orthogonal qr- and draxes are illustrated in Fig", " As the magnetic permeability of those possible conductor materials are closed to that of the air, the magnetic flux paths of the adopted DOL_SynRMs will be maintained at their respective original SynRM ones at the steady-state synchronized speeds. Therefore, by increasing the conductor volumes in the Authorized licensed use limited to: UNIVERSITY OF VICTORIA. Downloaded on April 06,2020 at 00:56:18 UTC from IEEE Xplore. Restrictions apply. 0093-9994 (c) 2020 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. directions as indicated in Fig. 2, it is clear that the equivalent resistances in those current-circulation loops will be reduced and larger stator winding currents in the corresponding axes will be generated during the motor start-up periods. Consequently, from (1), (4), and (5), it is clear that the evaluation references for the DOL_SynRM capabilities can thus be appropriately established by setting the most severe start-up load-speed characteristics (Load D) for the motordriven equipment, along with all the available rotor barrier spaces being filled with conductors", " In general, with the design objectives of high-efficiency operations for different motor-driven equipment in the metal industry, the above general constraint as listed in (7) might not lead to the most appropriate solution designs. Therefore, to design suitable DOL_SynRMs for these equipment applications, the dynamic start-up characteristics based on the specific load types and their corresponding equivalent shaft inertias must be thoroughly assessed. With the inherent features of robust structure and relatively higher operational efficiency, as conceptually illustrated in Fig. 2, a feasible scheme for equipping the SynRM with line-start capability is to fill its available rotor flux-barrier spaces with appropriate conductors and then short-circuit these conductors at both ends of the rotor. Based on such structural concepts, as depicted in Fig. 5, almost the same steady-state operational characteristics can be exhibited from the two SynRMs with identical conductor volumes being filled into the available flux-barrier spaces of these two rotor structures [11]. For further line-start capability confirmations at different load-speed and inertia conditions, the dynamic start-up characteristics of these DOL_SynRMs at two extreme cases were investigated and their corresponding speed patterns at these start-up periods are illustrated in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002163_is48319.2020.9199967-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002163_is48319.2020.9199967-Figure5-1.png", "caption": "Figure 5. Developed arm\u2019s lifting part.", "texts": [ " Thus, owing to the dynamics of the rubber, the arm contacts an object from the root to the tip by pulling the string, as shown in Fig. 4. This action of \u201ccontacting from the root to the tip\u201d is like that of an octopus [12-13], and by this action, the arm can cover and grasp an unknown object without measuring its shape. This complex grasping behavior is realized by the interaction between the arm and the object using the dynamics of the soft arm. Silicone block (30 A) Silicone (60 A) String cloth 10.8 [mm] 5.0[mm] PTFE tube String Fixed Object Pull string Loose string Figure 4. Movement of the arm\u2019s grasping part. Fig. 5 shows the lifting part. The lifting part is also made of silicone rubber 60A, and cloth is embedded on both sides of the arm for reinforcement. By pulling the string, the lifting part shrinks (Fig. 6), and the other part of the arm is lifted. Pull string Loose string Pull string Loose string Loose string Pull string Figure 6. Movement of the arm\u2019s lifting part. Fig. 7 shows the whole arm. The string is connected to the pulley and is pulled and loosened by the motor. Since the lifting part has a higher elasticity than the grasping part, the grasping part holds the object first; then, the lifting part lifts the body" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001885_tvt.2020.2986395-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001885_tvt.2020.2986395-Figure1-1.png", "caption": "Fig. 1. Cross-sectional drawing, winding arrangement of CP and MSCP PM machines. (a) CP PM machine. (b) MSCP PM machine. (c) MSCP-AW1 PM machine. (d) MSCP-AW2 PM machine. (e) MSCP-AT PM machine.", "texts": [ " The electromagnetic characteristics, including open-circuit field distributions and back-EMF waveforms, torque characteristics, unbalanced magnetic force (UMF), loss distribution characteristics as well as efficiency are compared and given in Section IV. In Section V, the modular prototype machine together with the AW is manufactured and tested to verify the theoretical and FE analysis. Conclusions are given in Section VI. II. MACHINE TOPOLOGIES The 18-slot/24-pole (18/24) fractional slot concentrated winding (FSCW) CP PM machine is composed by six 3- slot/4-pole (3/4) unit sub-machine, as shown in Fig. 1(a). The sub-machine exists even-order back-EMF harmonics due to the number of back-EMF phasor of one phase is one. In order to investigate the influence of modular stator (MS) on the electromagnetic performance of CP PM machine, the MS structure, MS with various auxiliary distributed winding as well as the MS with AT are employed. The section diagram together with the detailed winding connection of the MSCP PM machine, MSCP PM machine with long end AW (MSCP-AW1), MSCP PM machine with short end AW (MSCP-AW2) and MSCP PM machine with AT (MSCP-AT) are employed, as shown in Figs", " The influence of various AWs on the electromagnetic performance will be investigated under the same copper loss (Pcu=16.8W). Although the winding connection of MS can eliminate the even-order back-EMF harmonics effectively, however, its effective air-gap length is larger than that of the traditional non-modular stator, resulting in lower torque density. In order to reduce the effective air-gap improve the torque density, whilst realizes the physically isolated between the stator modular, the MS machine with auxiliary teeth is proposed, as presented in Fig. 1(e). Authorized licensed use limited to: University of Exeter. Downloaded on June 19,2020 at 02:47:19 UTC from IEEE Xplore. Restrictions apply. 0018-9545 (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. III. ELIMINATE PRINCIPLE OF EVEN-ORDER HARMONICS AND EFFECT OF AUXILIARY WINDING The even-order back-EMF harmonics have negative influence on the torque characteristics", " 1(c) and (d), respectively, and the AW back-EMF phasor of phase A together with the main winding of the two machines are shown in Figs. 7(a) and (b), respectively. Symbol AW21 and Authorized licensed use limited to: University of Exeter. Downloaded on June 19,2020 at 02:47:19 UTC from IEEE Xplore. Restrictions apply. 0018-9545 (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. AW22 denote the two distributed coils belong to one phase in the MSCP-AW2 PM machine, as shown in Fig. 1(d). It can be observed that the back-EMF of AW1 and AW2 can be in phase with the main winding. Fig. 8 gives the backEMF waveforms of MW and AW, which verifies the validity of the star of slots. The back-EMF waveforms of AW1 and AW2 together with modular-1 and 2 are shown in Figs. 8(a) and (b), respectively, and their harmonic contents are listed in Table. II. It can be noted that the even-order back-EMF harmonic can be eliminated effectively by the modular-1 and 2. In addition, the back-EMF waveform of the AW can be in phase with MW" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000489_s00170-016-9165-4-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000489_s00170-016-9165-4-Figure1-1.png", "caption": "Fig. 1 Plots of a the experimental equipment configuration and b the experimental setup for measuring friction force of the linear motion (LM) guide system", "texts": [ " To quantify the temperature rise from friction heat within a linear feed system, steady-state thermal analysis was performed using the finite element method [13]. The steady-state temperature distributions were also measured and compared with theoretical estimations. Our results indicated a maximum temperature difference of 13 %. A friction measurement system was created to measure the friction force of a LMB [14]. The one-axis feed drive system consisted of linear motion (LM) guide system, a LM actuator (THK SKR-46A), a servomotor, and a data acquisition (DAQ) device, as shown in Fig. 1. Two LM guides (THK HSR-30R) were installed facing each other, above and below. A jig attachment was used to apply a normal force to the LMB. The LM guide system used clamp\u2013clamp supports. The LMB traveled a reciprocating feed of 400 mm on a LM rail of 840 mm. The LM guide system was connected with a LM actuator by a kinematical ball joint, to reduce moment effects. The LM actuator, driven by the servomotor, generated a reference motion of reciprocating linear motion for the LM guide to follow. The servomotor driven axis was controlled by a Delta-Tau Turbo PMAC2, with 0", " Velocity was set at three levels: 100 mm/s (low), 150 mm/s (center), and 200 mm/s (high), based on the suitable operating range of an experimental LM guide system. Three normal loads were used: no load (low), 300 kgf (center), and 600 kgf (high). LMB manufacturers commonly provide the preload code to indicate preload level: common, light (C1), and heavy preload (C0). In this research, the preload levels were replaced by representative values: 0 (common preload), 1 (light preload), and 2 (heavy preload). The measurement of friction force was conducted using the experimental setup shown in Fig. 1. The friction force was directly measured using a horizontally coupled loadcell, while a full stroke of the LMB represented reciprocal movement. Figure 3a shows a typical Stribeck friction curve corresponding to the velocity. Across four well-defined lubrication regimes [5], this curve can be clearly divided into two parts: partial fluid lubrication and full-fluid lubrication; the division occurs at a velocity of \u223c20 mm/s. In this study, we only considered velocity under full-fluid lubrication conditions", " Figure 5a, b compares the predicted and experimental results of the friction force on the LMB, respectively. The ideal estimate line shows conditions in which the results predicted by the BBD model are in good agreement with the experimental results. The prediction equations show differences within 5 and 9 % for the initial BBD and additional experiments, respectively. Figure 6 shows the experimental apparatus used to measure temperature distributions in a LM guide system. Thermocouples and a data logger were added to the original experimental setup (Fig. 1) to acquire temperature data. Ten thermocouples were positioned along the LM guide system; six thermocouples were attached to the bearing block, and three were attached to the rail, as shown in Fig. 6. One thermocouple was used to measure ambient temperature. Four load/velocity conditions were considered for the temperature experiments: (i) 300 kgf and 100 mm/s, (ii) 300 kgf and 150 mm/s, (iii) 300 kgf and 200 mm/s, and (iv) 500 kgf and 100 mm/s. A heavy preload was applied in all cases. During the experiment, the LMB had a reciprocated stroke of 400 mm until the temperature reached a steady state" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002828_j.jii.2021.100265-Figure9-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002828_j.jii.2021.100265-Figure9-1.png", "caption": "Fig. 9. (a) The torch rotation around Rz, and (b) The simulation results with the adjusted torch angle.", "texts": [ " Correct use of the torch angles Rx and Ry are essential in achieving a high-quality weld. So experimental determination of suitable ranges for each of these variables is critical for practical applications. Note that as the tip of the welding gun is symmetric around the wire electrode, rotation around Rz L. Yuan et al. Journal of Industrial Information Integration xxx (xxxx) xxx L. Yuan et al. Journal of Industrial Information Integration xxx (xxxx) xxx will not affect the deposition process and the resultant geometry of the weld bead, as shown in Fig. 9(a). Thus, the value of Rz can range from -180\u25e6 to 180\u25e6. The other available parameter adjustment ranges are given based on our experimental tests and are listed in Table 4, below. Finally, the obtained information will be sent to a code translator to generate robot code for a selected industrial robot (we have made templates for many kinds of industrial robot in the code translator for user selection). For the example part outlined in Fig. 6, only one collision exists when we enter the torch pose adjustment stage. As shown in collision matrix C shown in Fig. 8, whilst performing the deposition process for layer 2-7, collision with layer 3-8 was detected. Torch angles were adjusted algorithmically, and the required robot motions were updated to avoid this collision. As simulation results in Fig. 9(b) indicate, the identified collision was removed by implementing a new deposition strategy. Note that the red layers in the figure are layers 2-8 to 2-10, which will be deposited after layer 2-7, meaning they can be ignored. To test the effectiveness of the proposed algorithm, a sample part with overhanging features and closed structures was fabricated. The CAD model of the workpiece is provided in Fig. 10(a). The geometries of the two stacked structures indicate that collision between welding torch and previously deposited layers may occur, particularly when depositing the overhanging features" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001060_j.jmapro.2019.11.004-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001060_j.jmapro.2019.11.004-Figure1-1.png", "caption": "Fig. 1. Welding plate geometry with different groove depths (H): (a) Stair-stepping plate (for H=4mm, 6mm, 8mm or 10mm, 12mm, 14mm or 16mm, 18mm, 20mm) (b) normal plate (for H=2mm) and (c) normal plate (for H=0).", "texts": [ " The influences of groove constraint space on the temperature, electron density and morphology of plasma and droplet transfer behavior were observed and discussed. This study should be helpful to understand the plasma behaviors influenced by the groove space and beneficial to optimize the groove geometry and process parameters during the laserarc hybrid welding of thick plate. In experiment, the Ti-6Al-4V plate was used as base metal and the standard Ti-6Al-4V welding wire was used with a diameter of 1.2mm. The chemical compositions of base metal and wire feedstock are shown in Table 1. Fig. 1 shows the welding plate geometry. The V-groove of welding plates was prepared with a groove angle of 20\u00b0 and different groove depth (varies from 0 to 20mm with an interval of 2mm). The corresponding weld plate size for each groove depth was 50mm\u00d750mm. For welding experiment with a groove depth of 4\u201320mm, stair-stepped shape plates with three steps were employed and each step corresponded to a groove depth. The normal plates were only used for the groove depth of 0 and 2mm. The bottom of the groove was designed as a plane with a width of 4mm, and the distance from the plane to the plate bottom was 5mm" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002689_ls.1554-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002689_ls.1554-Figure3-1.png", "caption": "FIGURE 3 Sessile drop contact angle of (a) water and (b) coconut oil on the steel surface [Color figure can be viewed at wileyonlinelibrary.com]", "texts": [ "94 revealing (001) plane of GO with an interlamellar spacing of 0.81 nm. The presence of oxygen functionalities in the basal plane of GO expanded the interlamellar spacing. The lack of organised structure and a limited number of lamellae in each GO signified a broader signature due to the (001) plane.53 The affinity of lubricating media towards the engineering surfaces governs the thin film formation ability, consequently enhancing the lubrication properties. In the present work, coconut oil is used as a lube base oil to formulate the grease. Figure 3 displays sessile drops of coconut oil and water on the polished sample of steel. The low contact angle ( 20 ) by coconut oil compared to water (75 ) implied that the coconut oil furnished excellent wettability to the steel surface, which could enhance the lubrication effect by coconut oil-based grease. The spherical silica nanoparticles, combined with 2D lamellar materials, that is, MoS2 and GO sheets in variable weight ratio, were thoroughly blended during the formulation of coconut grease. Two different combinations based on spherical particles and lamellar sheets are selected to explore their synergistic effects on the physicochemical and tribological properties of coconut grease", " Moreover, the shearthinning characteristics of grease increased the shear rate.41 The tribo-experiments were performed at 75 C and the heat generated at the contact zone because of frictional force further increase the temperature, which led to the softening of the grease. The high temperature and shear rate compromised the strengthening of the coconut grease's fibres network, which resulted in bleeding of the trapped coconut base oil. The thoroughly dispersed nanoadditives are also released and lubricate the tribo-interfaces. The excellent wettability of coconut oil (Figure 3(b)) has excellent potential to lubricate the tribosurfaces effectively by forming a tribo thin film. The morphology of nanoadditives plays a crucial role in the enhancement of tribological properties. The SiO2 nanoparticles (Figure 1(e),(f)) are spherical and impersonate as nano-bearing between the tribo-pairs, which converted the sliding friction into a combined effect of sliding and rolling friction, consequently lowering of the friction.24 The lubrication mechanism and the role of nanomaterials are governed by their shape, size, and composition" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001312_s00170-020-05434-3-Figure26-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001312_s00170-020-05434-3-Figure26-1.png", "caption": "Fig. 26 The designed bladeshaped part. a CADmodel. b The cross-sectional profile of the part with geometries (unit: mm)", "texts": [ " The reason for the error on the narrow side can be explained by the fact that the overlapping rate and the accumulated heat on the narrow side are large, which causes the molten pool to become large. Finally, the width of the fabricated parts becomes large. In order to reduce the error at the narrow side, the cooling time should be set between adjacent clad tracks. In this section, we place the mathematical model in the context of a case study to demonstrate its effectiveness in challenging practical applications. A typical variable width thin-walled structure, a blade-shaped part, is designed and fabricated. Figure 26a shows the CAD model of the designed part. Geometrical data of the part are presented in Fig. 26b. The maximum width of the thin- walled part is 6.89 mm and the minimum width is 2.4 mm. The number of tracks is calculated to be 3, and the overlapping rate is 0.06\u20131. The scanning speed range is 261.6\u2013600 mm/min. The dotted line shows the scan path. The scanning direction of adjacent layers should be reversed. The order of deposition in a single layer is deposited from the middle to the sides. The total height of the part was about 18 mm. The manufacturing parameters obtained by mathematical model calculation are shown in Table 6" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000173_01691864.2019.1680316-Figure13-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000173_01691864.2019.1680316-Figure13-1.png", "caption": "Figure 13. x\u0304 in the four cases depending on the contact point.", "texts": [ " These angles are calculated as \u03b11 = 2atan2 ( 2Lh \u2212 \u221a (d2 + h2)(4L2 \u2212 d2 \u2212 h2) d2 + 2Ld + h2 ) , (3) \u03b12 = 2atan2 ( 2Lh + \u221a (d2 + h2)(4L2 \u2212 d2 \u2212 h2) d2 + 2Ld + h2 ) , (4) \u03b21 = 2atan2 ( 2Llz \u2212\u221a (l2x + l2z)(4L2 \u2212 l2x \u2212 l2z) l2x + 2Llx + l2z ) , (5) \u03b22 = 2atan2 ( 2Llz +\u221a (l2x + l2z)(4L2 \u2212 l2x \u2212 l2z) l2x + 2Llx + l2z ) , (6) where lx = d + x\u0304 + l, (7) lz = 2h \u2212 L sin\u03b12, (8) x\u0304 is the position between the third wheel of the elongating part and the contact point along x axis, and l \u2265 0 is a design parameter to change the x position of the first wheel in the elongating part as Figure 12. The contact point between the robot and the stair is categorized as four cases according to the parameters as Figure 13. Let lb be the distance between the center and the bottom of the link. x\u0304 can be geometrically calculated considering Figure 13. (3) and (4) have a solution without imaginary parts because of (2). If (5) and (6) have a solution without imaginary parts, the following condition has to be satisfied: l2x + l2z \u2264 4L2. (9) Figure 14 shows the relationship between d and the maximum h which is numerically obtained in the case where l = 0.01m, L = 0.181m, r = 0.075m, and lb = 0.025m. Let the start time of the shift be t = 0 and the end time of the shift be t = tshift. If the shift from Figure 11(a,b) is carried out, we design the joint angles according to the following geometric relationships as Figure 15: \u03b8i+2 = (\u03b21 \u2212 \u03b12)t\u2032 + \u03b12, (10) \u03b8i+3 = (\u03b22 \u2212 \u03b11)t\u2032 + \u03b11, (11) \u03b8i = 2 atan2 ( 2L z\u2212 \u221a ( x2+ z2)(4L2\u2212 x2\u2212 z2) x2+ z2+2L x ) , (12) \u03b8i+1 = 2 atan2 ( 2L z+ \u221a ( x2+ z2)(4L2\u2212 x2\u2212 z2) x2+ z2+2L x ) , (13) where t is time, t\u2032 = 6t\u03035 \u2212 15t\u03034 + 10t\u03033 is the quintic curve, given as a cam curve, t\u0303 = t/tshift, 0 \u2264 t \u2264 tshift, and x = d + L(cos\u03b21 + cos\u03b22) \u2212 L(cos \u03b8i+2 + cos \u03b8i+3), (14) z = h + L(sin\u03b21 + sin\u03b22) \u2212 L(sin \u03b8i+2 + sin \u03b8i+3)" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002450_j.compstruct.2021.113608-Figure29-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002450_j.compstruct.2021.113608-Figure29-1.png", "caption": "Fig. 29. The analysis diagram of attitude accessibility.", "texts": [ " First, the three\u2010dimensional space of the part surface was obtained as shown in Fig. 26, which was calculated according to the model of the mould surface. The comparison diagram of the position space of the AFP machine and the three\u2010dimensional space of the part was shown in Fig. 27. Next, the mapping of the normal vector of the part surface on the Gaussian sphere was obtained as shown in Fig. 28. Comparative analysis of attitude reachable space mapping and normal vector distribution range mapping was shown in Fig. 29. According to the analysis results in Figs. 27 and 29, the position and attitude of the gantry AFP machine can be reached for this part. Based on the above analysis, appropriate roller specifications are CR64_20, CR64_40 and CR64_60. Considering the laying efficiency, select CR64_60 from the three specifications of the compaction roller. According to the compaction roller specification and prepreg tow width (1/4 in.), 1/4 in. prepreg tow was selected and the number of single course of tows was 8" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001981_s10846-020-01213-0-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001981_s10846-020-01213-0-Figure2-1.png", "caption": "Fig. 2 Overall system with main reference frames", "texts": [ " 1b) and the mobile frame to the rotor frame (pitch joint, see Fig. 1c). This aerial vehicle is able to move in all directions without changing the platform attitude, but only modifying the thrust direction by acting on the roll and pitch joints. In order to grasp and move the object, let us assume that a grasping tool (see Fig. 1d) rigidly connects the platform to the object and that a force-torque sensor is mounted on the tool, in order to measure the contact wrench between the object and the ODQuad. Let us define the following coordinate frames (see Fig. 2): \u2013 The inertial coordinate frame, {O, x, y, z}; \u2013 The coordinate frame k{Ok, xk, yk, zk}, attached to the kth robot platform (k = 1, 2, . . . , N); \u2013 The coordinate frame ek {Oek , xek , yek , zek }, attached to the tool and with origin at the contact point; \u2013 The coordinate frame o{Oo, xo, yo, zo} attached to the object center of mass. 3Modeling of the Aerial Vehicles The model of the ODQuad aerial vehicle has been derived in [17] and is briefly reviewed here. The configuration of the kth vehicle in the cooperative system (k = 1, " ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002002_012076-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002002_012076-Figure3-1.png", "caption": "Figure 3. Geometry obtained in ANSYS workbench", "texts": [ "585 since the relation would not be a linear one where: L = load limit at pressure P L 0 = maximum load limit at P 0 L/P n = const. V. V =air volume ICMSMT 2020 IOP Conf. Series: Materials Science and Engineering 872 (2020) 012076 IOP Publishing doi:10.1088/1757-899X/872/1/012076 Equating (1) and (2)... L 0 / P n 0 *P n =6.65xP 0.585 xS 1.702 x[(D R +S)/(19+S)] P0 can be calculated which is the required pressure. The succeeding approach was followed to solve the problem and the same was checked for both the materials during the analysis. In figure 3 its clearly being highlighted with inner portion of the tire completely filled with a polyethylene material. As we are not considering the non-linearity behavior of the tire, we are going forward with Static structural analysis as our major form of analysis. Figure 4 illustrate the project model in case of existing condition and polyethylene packaged. The tire of the wheel is in no-slip condition (bonded), the rim and the tire are in contact (no separation) as shown in figure 5. ICMSMT 2020 IOP Conf" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000557_aim.2016.7576989-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000557_aim.2016.7576989-Figure6-1.png", "caption": "Figure 6. Dimensions of the unit", "texts": [ " Using this continuous process, the bowel can transport boluses by application of a small force. We mimic the function of one circular muscle for one unit. The circular muscle is mimicked by a cylindrical tube. The straight-fiber-type artificial muscle mimics the longitudinal muscle. Each unit has circular contraction and relaxation and an axial contraction movement in order to imitate bowel peristalsis. The cross-sectional diagram of the peristaltic conveyor unit, the dimensions of the unit, and the unit specifications are shown in Fig. 5, Fig. 6, and Table I, respectively. We use a straight-fiber-type artificial muscle, a cylindrical tube, and flanges. The cylindrical tube is arranged inside the artificial muscle. The flange side has a vent for the supply and exhaust of the air with the vent connected to the chamber between the artificial muscle and the cylindrical tube. Upon application of air pressure, the artificial muscle expands on the outside, and the cylindrical tube expands on the inside. Fig. 7 shows the unit in the normal and pressurized states" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000859_1077546319856147-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000859_1077546319856147-Figure1-1.png", "caption": "Figure 1. Timoshenko beam element.", "texts": [ " As useful information for condition monitoring and fault diagnosis, the variation in the average values of bearing stiffness components is investigated as a function of rotational speed, internal radial clearance, and rotor shaft flexibility. The rotor shaft is considered as an elastic deformable free body which is subjected by external forces from ball bearings, gravity, unbalance forces, and gyroscopic moments. The FE model (Hashish and Sankar, 1984) for the flexible shaft is coupled with the bearing stiffness matrix at bearings\u2019 nodes. The equations of motion of the Timoshenko FE are assumed to be analogous in two orthogonal planes, to simplify the intricacies due to unified theory. Following the Timoshenko beam theory (see Figure 1) \"zz z,x\u00f0 \u00de \u00bc x @\u2019@z \u00f0a\u00de #x \u00bc @X @z \u2019 \u00f0b\u00de Vx \u00bc GsA#x \u00f0c\u00de My \u00bc EI @\u2019@z \u00f0d \u00de 9>>>>= >>>>; \u00f01\u00de Kinetic energy of beam element T \u00bc 1 2 Z 1 0 mz @X @t 2 \u00fe @Y @t 2 \" #( \u00fe JT @\u2019 @t 2 \u00fe @ @t 2 \" # \u00fe!rJz @\u2019 @t \u2019 @ @t dz \u00f02\u00de and potential energy of beam element P \u00bc 1 2 Z 1 0 EI @\u2019 @z 2 \u00fe @ @t 2 \" #( \u00fe AGs @X @z \u2019 2 \u00fe @Y @z 2 \" #) dz \u00f03\u00de Here mz, JT, and JZ represent mass, transverse, and polar mass moments of inertia per unit length, and !r is rotational speed of the rotor. The work done by distributed loads Wx \u00bc 1 2 R 1 0 fx z, t\u00f0 \u00de:x dz \u00f0a\u00de Wy \u00bc 1 2 R 1 0 fy z, t\u00f0 \u00de:y dz \u00f0b\u00de ) \u00f04\u00de The differential equations of motion for two orthogonal planes, that is, x-z and y-z are derived using Hamilton\u2019s principle, which are as follows @ @Z AGs @X @z \u2019 \u00fe fx z, t\u00f0 \u00de \u00bc mz @2X dt2 \u00f0a\u00de @ @Z AGs @X @z \u2019 \u00fe fx z, t\u00f0 \u00de \u00bc mz @2X dt2 \u00f0b\u00de @ @Z EI @\u2019 @Z \u00fe AGs @X @z \u2019 \u00bc JT @2\u2019 dt2 \u00fe " ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001934_j.mechmachtheory.2020.103895-Figure11-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001934_j.mechmachtheory.2020.103895-Figure11-1.png", "caption": "Fig. 11. Superposition of deformations for a vertical load in condition of axial symmetry.", "texts": [ " Employing the hypotheses of linearity due to the small displacements assumption, the symmetry of configuration and loads, and remembering that the membrane stiffness of a flat shell is much higher than the bending stiffness, a simple verification can be conducted. It can be observed that when the robot, positioned in a configuration of axial symmetry such as p = (0 , 0 , \u22120 . 5) , is subject to a vertical force F z , its elastic deformation should satisfy the superposition principle being that of two springs in series. Fig. 11 explains this concept and it can be observed that the total deformation of the endeffector along the z -axis ( Tables 6 and 7 ) is the sum of the deformation of the legs when M is rigid ( Tables 8 and 9 ) + the deformation of M when the legs are rigid and all joints are locked + the deformation due to the membrane compliance. However, the membrane stiffness is about 100 times stiffer than the bending stiffness, thereby neglecting the membrane deformation should produce only minimal errors. From Table 11 it can be observed that for the model A1 the residual displacement term deriving from the membrane effect is too high whereas it should be proportional to 10 \u22126 (m)" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001373_j.promfg.2020.05.106-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001373_j.promfg.2020.05.106-Figure1-1.png", "caption": "Fig. 1. ASTM E8 Tensile samples showing build geometry and layer orientation: a) 0 degree orientation; b) 15 degree orientation; c) 30 degree orientation; d) specimen dimensions in mm\u20143 mm thickness not shown.", "texts": [ " Five individual tensile specimens were then extracted from the block via wire EDM. The specimens were extracted in such a manner so that the effective build orientation with respect to the lengthwise axis of the samples were oriented at 0, 15, and 30 degrees. Samples were then polished using 1000 grit sandpaper to remove small surface defects and the recast layer left from the EDM process. The specimen cross-sections measured 3 mm x 6 mm with a 32 mm gage length. The geometry of the completed samples is shown in Figure 1. Using an Instron 5982 Universal Materials Testing System, the samples were loaded uniaxially until failure. Tests were conducted using stepped strain rate control, 0.05 mm/mm/min (1.6 mm/min crosshead speed) though the yield point and then 0.5 mm/mm/min (16 mm/min crosshead speed) until failure was detected. Displacement was measured using a contact extensometer affixed to the gage section; the extensometer measured 25.4 mm between arms and had to be removed at 10% strain to prevent damage to the device" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002332_j.matpr.2020.08.560-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002332_j.matpr.2020.08.560-Figure1-1.png", "caption": "Fig. 1. 3D Meshed model.", "texts": [ " It modified into introduced thru the organization that developed it, Parametric degree within the release of this suite of layout products that embody packages which incorporates assembly modeling, 2 dimensional orthographic views for technical drawing, finite detail analysis [19] and extra below (Table 1). In this present work, I\u2019m proposing these materials of aluminum alloy 7475 and aluminum 6061 as they high corrosion resistance, less machining time, lighter in weight, costing is also less when compared to other materials and these alloys are easily welded and joined (Fig. 1). Limited detailed appraisal is a method of comprehending, ordinarily about, the positive difficulties in designing and period. Hence it is utilized predominantly for inconveniences for which the no interesting arrangements, expressible in some scientific shapes, is been accessible. Accordingly, it\u2019s miles a numerical instead a logical technique. Strategies for this sort are wished because of the reality diagnostic methods can\u2019t govern the genuine, complex difficulties which might be met inward designing" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001251_j.measurement.2020.107825-Figure10-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001251_j.measurement.2020.107825-Figure10-1.png", "caption": "Fig. 10. Temperature distribution ( C) of a 1.5 mm line processing by laser beam.", "texts": [ " Accurate temperature value of melt pool was abbreviated as Tr5. Rcr \u00bc \u00f0Tcam=Tmn\u00de 1 10 \u00f018\u00de Tr5 \u00bc 14:16 Tcam\u00f0Tcam=Tmn\u00de 1 10 \u00f019\u00de Recorded temperature values of parts were put into the equation. All three manufacturing sets of samples examined at Fig. 7, Fig. 8 and Fig. 9 below. To find melt pool temperatures, transient thermal finite element analysis was performed on Ansys. As a result of finite element analysis, a sample temperature distribution that occurs after 1.5 mm long line laser treatment is shown in Fig. 10. manufacturing. anufacturing. As shown in Fig. 10, the temperature increases instantaneously at the point of contact of the laser, and then quickly reaches the temperature of the piece by immediately cooling down. Ti6Al4V material cooling rate is given as approximately 105\u2013106 K/s according to the literature [34,35] In the FEA method; it was found around 106 K/s in accordance with the literature. At the first moments, there was a rapid cooling because the temperature difference of the part and melt region was quite high. Thereafter, a cooling rate decreased gradually" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002002_012076-Figure10-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002002_012076-Figure10-1.png", "caption": "Figure 10. Total deformation on tire with air package", "texts": [], "surrounding_texts": [ "The calculated forces like bump force were applied by the road to the tire of the wheel with a force of 2500N and the analysis is carried out in the no-slip condition as shown in figure 7." ] }, { "image_filename": "designv11_5_0002039_00295639.2020.1777023-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002039_00295639.2020.1777023-Figure4-1.png", "caption": "Fig. 4. Interpretation adhesion area of the vacuum mobile platform.", "texts": [ " (6) by applying the following boundary condition: Pin \u00bc Pout 6Q\u03bc \u03c1\u03c0h ln Rout Rin ; \u00f06\u00de NUCLEAR SCIENCE AND ENGINEERING \u00b7 VOLUME 00 \u00b7 XXXX 2020 where Q is the leakage flow rate and h is the gap between the sealing pads and the wall. In addition, through the modified Bernoulli equation, the internal inlet airflow can be expressed as follows: P1 \u00bc P2 \u00fe \u03b3 V2 2 V1 2 2g \u00fe hL ; \u00f07\u00de where \u03b3 is the specific weight and hL is the head loss. To determine the suitability of flow modeling, we conducted a simulation evaluation. Flow analysis was conducted with the fluid dynamics software ANSYS CFX to check the adhesion pressure of the robot. As shown in Fig. 4, the boundary conditions are set as follows: The external pressure is the atmospheric pressure, the working fluid is air, and the wall is considered to be nonslip. The flow pressure inside the robot chamber was analyzed with an air gap of 0.05 mm as shown in Fig. 5. The pressure distribution at the air gap of 0.05 mm showed that the inside of the chamber had a negative pressure of approximately \u22124.2 kPa as the impeller rotated. Using Eq. (7), we can determine that the air gap must be smaller than 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002614_j.compositesa.2021.106449-Figure17-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002614_j.compositesa.2021.106449-Figure17-1.png", "caption": "Fig. 17. Representation of the gaps, overlaps, and tow drops using (a) shell element, (b) solid element (discrete tows are represented as different colours).", "texts": [ " Exporting geometry (as opposed to directly exporting mesh) has the advantage of utilising the powerful internal meshing algorithms available in commercial FE codes. In this work, the tow points were directly converted to geometry in MSC Marc Mentat using the available PyMentat module. Surface interpolation and meshing were conducted automatically within the Mentat GUI using PyMentat commands. Shell or solid mesh of the tow domain can be created depending on the desired modelling strategy. An example of the solid element and X. Li et al. Composites Part A 147 (2021) 106449 shell element representation is presented in Fig. 17, which uses the same geometric information as Fig. 16. In this case, a high order surface definition near the tow drop area, Bezier, is generated by regarding the tow points as a set of control points to smooth out the sharp edges due to the tow drop. The flat areas are modelled with linear surfaces. The ultimate mesh is generated by the surface to mesh convert function. For the solid element, it can be noted that slight penetration occurs due to the high order interpolation functions. The penetration can be removed before applying any mechanical loading in the FE package without introducing extra initial stress, such as the function \u201cadjust to remove overclosure\u201d in Abaqus [27]", " To tackle this issue, a higher-order surface definition, Bezier, is used to smooth out the sharp edges (from C0 to C1 continuity). The point cloud near the area of the defect (tow drop) is regarded as a set of control points, which are used to generate the Bezier surface. The Bezier surface can be converted to FE mesh in the FE package later. For the flat regions, a linear surface definition can be applied directly. A typical example comparing the geometry before and after smoothing is shown in Fig. 16 and Fig. 17. Although there will be slight penetrations induced by the higher-order interpolation, it can be removed entirely in the FE package without introducing extra initial stress before applying any mechanical loading, such as the function \u201cadjust to remove overclosure\u201d in Abaqus. One major benefit of the TWM is the generation of a structured or material aligned, quadrilateral mesh pattern as used by Falco\u0301 et al [28,29]. However, the triangular mesh is inevitable to be used near the trimming boundaries and the areas closed to the defects to avoid element warping or distortion" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001549_j.ymssp.2020.107328-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001549_j.ymssp.2020.107328-Figure3-1.png", "caption": "Fig. 3. Slider-crank mechanism with joints clearance.", "texts": [ " When the clearance of a joint changes due to wear, the contact force and sliding velocity of other joints will also be affected. Correspondingly, the wear of other joints will also change because the wear of twomating surfaces depends on the relative sliding distance and the contact force of contact interfaces [32]. The wear interactions between the joints may not be the same. The interaction among joints is transmitted by linkages between them. When there are more linkages between two joints, the interaction between these two joints will be weaker. Take a slider-crankmechanism as an example (see Fig. 3). The influence of revolute joint A on the motion of revolute joint C is smaller than that of revolute joint B because joint A and joint B are neighboring while joint A and joint C are connected by two linkages. 1) Correlation modeling based on vine-copula Copula has been widely used to deal with correlation problems in various filed, such as finance [33], hydrology [34] and mechanical engineering [36,36]. Someone parameter multivariate copulas such as multivariate Clayton copula and multivariate Gaussian copula [37,37], use one parameter to describe the correlation among variables", " Therefore, in this study, we only model the free flight and impact motions, and the distance between the journal center and bearing center is given as: ri \u00bc ci \u00fe yi \u00f012\u00de where ci and yi are the initial clearance and wear depth of ith revolute joint in a mechanism. In the free flight motion, the distance between the journal center and bearing center is smaller than the value in Eq. (12). When this value is adopted, it will give a conservative prediction of the RUL. Combining Eq. (12), the performance function in Eq. (11) can be further written as: g \u00bc G\u00f0L; r;u\u00de \u00f013\u00de Take the slider-crank mechanism in Fig. 3 as an example to illustrate how to obtain the performance function. In Fig. 3, the power drives the linkage AB to rotate clockwise, and h1 is defined as the input angle. The loop-closure equations of the slider-crank mechanism are given by: r1 cosu1 \u00fe L1 cosh1 \u00fe r2 cosu2 \u00fe L2 cosh2 r3 cosu3 G \u00bc 0 r1 sinu1 \u00fe L1 sinh1 \u00fe r2 sinu2 \u00fe L2 sinh2 r3 sinu3 L3 \u00bc 0 \u00f014\u00de Eliminating h2, we obtain the displacement G as: G \u00bc r1 cosu1 \u00fe L1 cosh1 \u00fe r2 cosu2 r3 cosu3 \u00fe ffiffiffi A p \u00f015\u00de where A \u00bc L22 \u00f0L3 \u00fe r3 sinu3 L1 sinh1 r1 sinu1 r2 sinu2\u00de2. From the performance function of the slider-crank mechanism, we can find that another influencing factor is the motion angles u(u1, u2,", " The posterior summaries of the parameters for the Gamma process and copulas are listed in Table 3. In the table, In the table, the symbols vA, vB, and vC represent shape parameters the Gamma process modeling the wear depth of Joints A, B, and C in the slider-crank mechanism, respectively. The symbols uA, uB, and uC denote the corresponding scale parameters, respectively. The numbers 2.50% and 97.50% represent the 95% posterior intervals of the parameters (based on the 2.5% and 97.5% posterior percentiles). The schematic diagram of the slider-crank mechanism is shown in Fig. 3. With the driving system, the angle h2 between linkage L1 and L2 is time-varying. Then, it drives the slider-crank mechanism to operate. Hence, h2 is defined as the input angle. Eliminating h1 in the loop-closure equations from Eq. (14), we obtain the displacement G as: G \u00bc r1cosu1 \u00fe L2cosh2 \u00fe r2cosu2 r3cosu3 \u00fe ffiffiffi A p \u00f024\u00de where A \u00bc L21 \u00f0L3 \u00fe r3sinu3 L2sinh2 r1sinu1 r2sinu2\u00de2. The position of the slider while h2 = 0 is investigated in this study. Based on Eq. (16), the motion angles um(um,1, um,2, um,3) leading to the largest deviation of the desired accuracy can be obtained as: In a mechanism, the size of the clearance is far less than the dimension of linkages" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001852_physreve.102.022610-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001852_physreve.102.022610-Figure3-1.png", "caption": "FIG. 3. Graphical representation of the correction to the linear terms in the equation of motion (4.1).", "texts": [ " We reorganize the resultant EOM so that it has the same form as (4.1) but with various coefficients renormalized. The calculation of the renormalization of the coefficients arising from the process of eliminating u> \u22a5(k\u0303) can be represented by graphs. The basic rules for the graphical representation are illustrated in Fig. 2. Following these rules and the prescription of Ref. [30], the renormalization of the linear terms and the noise to oneloop order are represented by the graphs in Figs. 3 and 4, respectively. For example, Fig. 3(a) represents a linear term in the EOM for u\u22a5 j given by \u22122D\u03bb2k\u22a5 u u\u22a5 c (k\u0303) (2\u03c0 )d+1 \u222b > q\u0303 (k\u22a5 i \u2212 q\u22a5 i )Ciu(q\u0303)Gjc(k\u0303 \u2212 q\u0303), (4.4) where \u222b > q\u0303 \u2261 \u222b +\u221e \u2212\u221e d \u222b +\u221e \u2212\u221e dqx \u222b e\u2212d (k\u0303) of the noise f , which gives a closed EOM for u< \u22a5(k\u0303). This step can be represented by graphs. The basic rules for the graphical representation are illustrated in Fig. 2. We will now evaluate these graphs, each of which can be interpreted as adding a term to the equation of motion for u< \u22a5(k\u0303). We begin with the graph in Fig. 3(a). 1. Renormalizations of the \u03bc\u2019s a. Graph in Fig. 3(a) The graph in Fig. 3(a) gives a contribution (\u2202t u< j )\u03bc,a to the EOM for u< j (k\u0303): (\u2202t u < j )\u03bc,a = \u22122D\u03bb2k\u22a5 u u\u22a5 c (k\u0303) (2\u03c0 )d+1 \u222b > q\u0303 (k\u22a5 i \u2212 q\u22a5 i )Ciu(q\u0303)Gjc(k\u0303 \u2212 q\u0303) \u2261 \u22122D\u03bb2k\u22a5 u u\u22a5 c (k\u0303) [( I\u03bc,a 1 ) c ju (k\u0303) \u2212 (I\u03bc,a 2 ) c ju (k\u0303) ] , (A4) where \u222b > q\u0303 \u2261 \u222b >|q\u22a5|> e\u2212d dd\u22121q\u22a5 \u222b \u221e \u2212\u221e d \u222b \u221e \u2212\u221e dqx, (A5) Ciu(q\u0303) \u2261 |GT (q\u0303)|2\u03b4\u22a5 iu = \u03b4\u22a5 iu \u03c92 + (\u03bc1q2 \u22a5 + \u03bcxq2 x )2 , (A6) ( I\u03bc,a 1 ) c ju(k\u0303) \u2261 k\u22a5 i (2\u03c0 )d+1 \u222b > q\u0303 Ciu(q\u0303)Gjc(k\u0303 \u2212 q\u0303), (A7) ( I\u03bc,a 2 ) c ju(k\u0303) \u2261 1 (2\u03c0 )d+1 \u222b > q\u0303 q\u22a5 i Ciu(q\u0303)Gjc(k\u0303 \u2212 q\u0303). (A8) Since we are interested in terms only to O(k2), because that is the order of the \u03bc terms in the EOM, and (A4) already has an explicit factor k\u22a5 u , we need only expand these integrals (I\u03bc,a 1,2 )c ju to linear order in k", " Inserting (A9) and (A10) into (A4), we find: (\u2202t u < j )\u03bc,a = \u2212 { 1 32 D\u03bb2\u221a \u03bcx\u03bc 3 1 Sd\u22121 (2\u03c0 )d\u22121 d\u22124d [ 4 \u2212 3 (d \u22121) ]} k2 \u22a5u< j = \u2212 [ 1 8 \u2212 3 32(d \u2212 1) ] (g1\u03bc1d )k2 \u22a5u< j , (A12) 022610-23 where in the second equality we have used our earlier definition (4.11) of g1. Since this contribution to \u2202t u< j has the same form as the \u03bc1 term already present, we can absorb it into a renormalization of \u03bc1: (\u03b4\u03bc1)\u03bc,a = 1 32 D\u03bb2\u221a \u03bcx\u03bc 3 1 Sd\u22121 (2\u03c0 )d\u22121 d\u22124d [ 4 \u2212 3 (d \u2212 1) ] = [ 1 8 \u2212 3 32(d \u2212 1) ] g1\u03bc1d . (A13) b. Graph in Fig. 3(b) The graph in Fig. 3(b) gives a contribution (\u2202t u< j )\u03bc,b to the EOM for u< j (k\u0303): (\u2202t u < j )\u03bc,b = \u22122\u03bb2Dk\u22a5 u u\u22a5 c (k\u0303) (2\u03c0 )d+1 \u222b > q\u0303 q\u22a5 i Cju(q\u0303)Gic(k\u0303 \u2212 q\u0303) = \u22122\u03bb2Dk\u22a5 u u\u22a5 c (k\u0303)(I\u03bc,b)c ju(k\u0303), (A14) where (I\u03bc,b)c ju(k\u0303) \u2261 1 (2\u03c0 )d+1 \u222b > q\u0303 q\u22a5 i Cju(q\u0303)Gic(k\u0303 \u2212 q\u0303). (A15) Inserting (A3) and (A6) into (A15) we get (I\u03bc,b)c ju(k\u0303) = \u03b4\u22a5 u j (2\u03c0 )d+1 \u222b > q\u0303 q\u22a5 c |GT (q\u0303)|2GT (k\u0303 \u2212 q\u0303) = \u03b4\u22a5 u j (2\u03c0 )d+1 \u222b > q\u0303 q\u22a5 c |GT (q\u0303)|2[G(\u2212q\u0303)]2 \u00d7 (2\u03bc1q\u22a5 \u00b7 k\u22a5 + 2\u03bcxqxkx ), = 3\u03b4\u22a5 u jk \u22a5 c 64(d \u2212 1) 1\u221a \u03bcx\u03bc 3 1 Sd\u22121 (2\u03c0 )d\u22121 d\u22124d ", " (A17) From the form of this correction (namely, the fact that it is proportional to k\u22a5 j k\u22a5 c u\u22a5 c ), we can identify this as a correction to \u03bc2 (which we remind the reader is the parameter whose bare value we took to be zero): (\u03b4\u03bc2)\u03bc,b = 3 32(d \u2212 1) D\u03bb2\u221a \u03bcx\u03bc 3 1 Sd\u22121 (2\u03c0 )d\u22121 d\u22124d . (A18) Thus, it would appear at this point that, even starting as we have with a model in which \u03bc2 = 0, we generate a nonzero \u03bc2 on renormalization. This proves to not be the case, at least to one-loop order. Instead, to this order, (A18) is exactly canceled by other graphs, as we will now see. c. Graph in Fig. 3(c) The graph in Fig. 3(c) gives a contribution (\u2202t u< j )\u03bc,c to the EOM for u< j (k\u0303): (\u2202t u < j )\u03bc,c = 2\u03bb2Du\u22a5 u (k\u0303) (2\u03c0 )d+1 \u222b > q\u0303 (k\u22a5 i \u2212 q\u22a5 i )q\u22a5 u Ci (q\u0303)Gj (k\u0303 \u2212 q\u0303) \u2261 2\u03bb2Du\u22a5 u (k\u0303) [( I\u03bc,c 1 ) ju(k\u0303) + (I\u03bc,c 2 ) ju(k\u0303) ] , (A19) where( I\u03bc,c 1 ) ju(k\u0303) \u2261 k\u22a5 i (2\u03c0 )d+1 \u222b > q\u0303 q\u22a5 u Ci (q\u0303)Gj (k\u0303 \u2212 q\u0303), (A20) ( I\u03bc,c 2 ) ju(k\u0303) \u2261 \u2212 1 (2\u03c0 )d+1 \u222b > q\u0303 q\u22a5 i q\u22a5 u Ci (q\u0303)Gj (k\u0303 \u2212 q\u0303). (A21) Inserting (A6) and (A3) into (A20) leads to ( I\u03bc,c 1 ) ju(k\u0303) = k\u22a5 j (2\u03c0 )d+1 \u222b > q\u0303 q\u22a5 u |GT (q\u0303)|2GT (k\u0303 \u2212 q\u0303) = k\u22a5 j (2\u03c0 )d+1 \u222b > q\u0303 q\u22a5 u |GT (q\u0303)|2[GT (\u2212q\u0303)]2 \u00d7 (2\u03bc1q\u22a5 \u00b7 k\u22a5 + 2\u03bcxqxkx ) = 3k\u22a5 j k\u22a5 u 64(d \u2212 1) 1\u221a \u03bcx\u03bc 3 1 Sd\u22121 (2\u03c0 )d\u22121 d\u22124d . (A22) We deliberately leave (I\u03bc,c 2 ) ju(k\u0303) untouched since will show later in next section that this piece is canceled out by that from Fig. 3(d). Inserting only this (I\u03bc,c 1 ) ju contribution (A22) into (A19) leads to a term in the equation of motion for u< j : (\u2202t u < j )\u03bc,c = [ 3 32(d \u2212 1) D\u03bb2\u221a \u03bcx\u03bc 3 1 Sd\u22121 (2\u03c0 )d\u22121 d\u22124d ] k\u22a5 j k\u22a5 u u\u22a5 u . (A23) which, as before, can be interpreted as a correction to \u03bc2: \u03b4\u03bc2 = \u2212 3 32(d \u2212 1) D\u03bb2\u221a \u03bcx\u03bc 3 1 Sd\u22121 (2\u03c0 )d\u22121 d\u22124d . (A24) Note that this exactly cancels the contribution to \u03bc2 from Fig. 3(b) that we just calculated. d. Graph in Fig. 3(d) The graph in Fig. 3(d) gives a contribution (\u2202t u< j )\u03bc,d to the EOM for u< j (k\u0303): (\u2202t u < j )\u03bc,d = 2\u03bb2Du\u22a5 u (k\u0303) (2\u03c0 )d+1 \u222b > q\u0303 q\u22a5 i q\u22a5 u Cj (q\u0303)Gi (k\u0303 \u2212 q\u0303) \u2261 2\u03bb2Du\u22a5 u (k\u0303)(I\u03bc,d )u j (k\u0303), (A25) where (I\u03bc,d )u j (k\u0303) \u2261 1 (2\u03c0 )d+1 \u222b > q\u0303 q\u22a5 i q\u22a5 u Cj (q\u0303)Gi (k\u0303 \u2212 q\u0303). (A26) 022610-24 This contribution cancels out the I\u03bc,c 2 contribution from Fig. 3(c) above exactly, leaving no correction to \u03bc2 at all, to one-loop order. Thus, to this order, \u03bc2 = 0 is a fixed point. 2. Noise renormalization The graphs in Fig. 4(a) and Fig. 4(b) represent the following two corrections to the noise correlator \u3008 f (k\u0303) fu(\u2212k\u0303)\u3009: \u3008 f (k\u0303) fu(\u2212k\u0303)\u3009D,a = 4\u03bb2D2 (2\u03c0 )d+1 \u222b > q\u0303 q\u22a5 i q\u22a5 mCim(k\u0303 \u2212 q\u0303)C u(q\u0303), (A27) \u3008 f (k\u0303) fu(\u2212k\u0303)\u3009D,b = 4\u03bb2D2 (2\u03c0 )d+1 \u222b > q\u0303 q\u22a5 i ( k\u22a5 m \u2212 q\u22a5 m ) Ciu(k\u0303 \u2212 q\u0303) \u00d7C m(q\u0303). (A28) Since the noise strength D is the value of this correlation at k = 0, we set k\u0303 = 0 in (A27) and (A28) to get \u3008 f (k\u0303) fu(\u2212k\u0303)\u3009D,a = 4\u03bb2D2 (2\u03c0 )d+1 \u222b > q\u0303 q\u22a5 i q\u22a5 mCim(\u2212q\u0303)C u(q\u0303), (A29) \u3008 f (k\u0303) fu(\u2212k\u0303)\u3009D,b = 4\u03bb2D2 (2\u03c0 )d+1 \u222b > q\u0303 q\u22a5 i (\u2212q\u22a5 m ) Ciu(\u2212q\u0303)C m(q\u0303)", " III, have, for the propagators: Gi j (k\u0303) \u2261 L\u22a5 i j (k\u22a5)GL(k\u0303) + P\u22a5 i j (k\u22a5)GT (k\u0303), (B1) with GL(k\u0303) = 1 \u2212i\u03c9 + \u03bcLk2 \u22a5 + \u03bcxk2 x , (B2) GT (k\u0303) = 1 \u2212i\u03c9 + \u03bc1k2 \u22a5 + \u03bcxk2 x , (B3) and the longitudinal and transverse projection operators L\u22a5 i j (k\u22a5) and P\u22a5 i j (k\u22a5), respectively, are defined as L\u22a5 i j (k\u22a5) \u2261 k\u22a5 i k\u22a5 j /k2 \u22a5, (B4) which projects any vector along k\u22a5, and P\u22a5 i j (k\u22a5) \u2261 \u03b4\u22a5 i j \u2212 k\u22a5 i k\u22a5 j /k2 \u22a5, (B5) which projects any vector onto the space orthogonal to both the mean direction of flock motion x\u0302 and k\u22a5. We now also have similar decompositions for the correlation functions: Ci j (k\u0303) \u2261 L\u22a5 i j (k)|GL(k\u0303)|2 + P\u22a5 i j (k)|GT (k\u0303)|2. (B6) With these in hand, we will now calculate the graphical corrections to the various parameters in the full model with \u03bc2 = 0. Note that all of the graphs are exactly the same as those we evaluated in the previous section; all that will change is their values, because we are now taking \u03bc2 = 0. 1. Renormalizations of \u03bc1,2,x a. Graph in Fig. 3(a) The graph in Fig. 3(a) gives a contribution (\u2202t u< j )\u03bc,a to the EOM for u< j (k\u0303): (\u2202t u < j )\u03bc,a = \u22122D\u03bb2k\u22a5 u u\u22a5 c (k\u0303) (2\u03c0 )d+1 \u222b > q\u0303 (k\u22a5 i \u2212 q\u22a5 i )Ciu(q\u0303)Gjc(k\u0303 \u2212 q\u0303) \u2261 \u22122D\u03bb2k\u22a5 u u\u22a5 c (k\u0303) {( I\u03bc,a 1 ) c ju(k\u0303) \u2212 (I\u03bc,a 1 ) c ju(k\u0303) } , (B7) where ( I\u03bc,a 1 ) c ju (k\u0303) \u2261 k\u22a5 i (2\u03c0 )d+1 \u222b > q\u0303 Ciu(q\u0303)Gjc(k\u0303 \u2212 q\u0303), (B8) ( I\u03bc,a 2 ) c ju(k\u0303) \u2261 1 (2\u03c0 )d+1 \u222b > q\u0303 q\u22a5 i Ciu(q\u0303)Gjc(k\u0303 \u2212 q\u0303). (B9) Note that the overall correction already has a common factor ku outside the loop integral. That means for both (I\u03bc,a 1 )c ju and (I\u03bc,a 2 )c ju we can set \u03c9 = 0 inside the loop integral since expanding the loop integral to O(\u03c9) or higher orders in \u03c9 only gives terms irrelevant compared to \u2212i\u03c9u\u22a5, which is already present in the EOM (4", " The overall correction to the equation of motion is thus: (\u2202t u < j )\u03bc,a = \u22122\u03bb2Dk\u22a5 u u\u22a5 c (k\u0303) [( I\u03bc,a 1 ) c ju(k\u0303) \u2212 (I\u03bc,a 2 ) c ju(k\u0303) ] = \u22122\u03bb2DKd d\u22124d k\u22a5 u u\u22a5 c (k\u0303) (d \u2212 1)(d + 1) \u00d7 \u23a1 \u23a3k\u22a5 u \u03b4\u22a5 jc \u239b \u239d\u2212 7 64 \u221a \u03bcx\u03bc 3 L + (d2 \u2212 2d \u2212 2) 16 \u221a \u03bcx\u03bc 3 1 + (d + 2)A(\u03bcL, \u03bc1) + dA(\u03bc1, \u03bcL ) \u2212 2d\u03bc1B(\u03bcL, \u03bc1) \u239e \u23a0 +k\u22a5 j \u03b4\u22a5 cu \u239b \u239d (4d \u2212 3) 32 \u221a \u03bcx\u03bc 3 L + 1 8 \u221a \u03bcx\u03bc 3 1 \u2212 2dA(\u03bcL, \u03bc1) \u2212 2A(\u03bc1, \u03bcL ) + 4\u03bc1B(\u03bcL, \u03bc1) \u239e \u23a0 \u23a4 \u23a6 022610-27 = \u22122\u03bb2DKd d\u22124d (d \u2212 1)(d + 1) \u00d7 \u23a1 \u23a3k2 \u22a5u\u22a5 j (k\u0303) \u239b \u239d\u2212 7 64 \u221a \u03bcx\u03bc 3 L + (d2\u22122d\u22122) 16 \u221a \u03bcx\u03bc 3 1 + (d+2)A(\u03bcL, \u03bc1)+dA(\u03bc1, \u03bcL )\u22122d\u03bc1B(\u03bcL, \u03bc1) \u239e \u23a0 +k\u22a5 j k\u22a5 u u\u22a5 u (k\u0303) \u239b \u239d (4d \u2212 3) 32 \u221a \u03bcx\u03bc 3 L + 1 8 \u221a \u03bcx\u03bc 3 1 \u2212 2dA(\u03bcL, \u03bc1) \u2212 2A(\u03bc1, \u03bcL ) + 4\u03bc1B(\u03bcL, \u03bc1) \u239e \u23a0 \u23a4 \u23a6 = \u2212 2\u03bc1g1d (d \u2212 1)(d + 1) { k2 \u22a5u\u22a5 j (k\u0303) [ \u2212 7 64 (1 + \u03b1)\u22123/2 + (d2 \u2212 2d \u2212 2) 16 + (d + 2) \u221a \u03bcx\u03bc 3 1A(\u03bcL, \u03bc1) + d \u221a \u03bcx\u03bc 3 1A(\u03bc1, \u03bcL ) \u2212 2d \u221a \u03bcx\u03bc 5 1B(\u03bcL, \u03bc1) ] + k\u22a5 j k\u22a5 u u\u22a5 u (k\u0303) [ (4d \u2212 3) 32 (1 + \u03b1)\u22123/2 + 1 8 \u22122d \u221a \u03bcx\u03bc 3 1A(\u03bcL, \u03bc1) \u2212 2 \u221a \u03bcx\u03bc 3 1A(\u03bc1, \u03bcL ) + 4 \u221a \u03bcx\u03bc 5 1B(\u03bcL, \u03bc1) ]} . (B12) b. Graph in Fig. 3(b) The graph in Fig. 3(b) gives a contribution (\u2202t u< j )\u03bc,b to the EOM for u< j (k\u0303): (\u2202t u < j )\u03bc,b = \u22122\u03bb2Dk\u22a5 u u\u22a5 c (k\u0303) (2\u03c0 )d+1 \u222b > q\u0303 q\u22a5 i Cju(q\u0303)Gic(k\u0303 \u2212 q\u0303) = \u22122\u03bb2Dk\u22a5 u u\u22a5 c (k\u0303)(I\u03bc,b)c ju(k\u0303), (B13) where (I\u03bc,b)c ju(k\u0303) \u2261 1 (2\u03c0 )d+1 \u222b > q\u0303 q\u22a5 i Cju(q\u0303)Gic(k\u0303 \u2212 q\u0303). (B14) Again, since there is already a factor k\u22a5 c outside the loop integral, we can set \u03c9 = 0 in the integrand. Also we only need to expand the integrand to O(k) and keep the O(k) part, since the integration of the zeroth-order part is odd in q\u22a5 i and hence vanishes", " (B15) The overall correction (\u2202t u< j )\u03bc,b to the EOM for u< j (k\u0303) is therefore: (\u2202t u < j )\u03bc,b = \u22122\u03bb2Dk\u22a5 u u\u22a5 c (k\u0303)(I\u03bc,b)c ju(k\u0303) = \u22122\u03bb2DKd d\u22124d k\u22a5 u u\u22a5 c (k\u0303) (d \u2212 1)(d + 1) { k\u22a5 u \u03b4\u22a5 jc [ 7 64 \u221a \u03bcx\u03bc 3 L + 1 16 \u221a \u03bcx\u03bc 3 1 \u2212 A(\u03bcL, \u03bc1) \u2212 A(\u03bc1, \u03bcL ) \u2212 2\u03bcLB(\u03bc1, \u03bcL ) ] + k\u22a5 c \u03b4\u22a5 ju [ (5 \u2212 2d ) 32 \u221a \u03bcx\u03bc 3 L + (d2 \u2212 2d \u2212 1) 16 \u221a \u03bcx\u03bc 3 1 \u2212 (d2 \u2212 2d \u2212 1)A(\u03bc1, \u03bcL )\u2212(1\u2212d )A(\u03bcL, \u03bc1) + 2(d \u2212 1)\u03bcLB(\u03bc1, \u03bcL ) ]} = \u2212 2\u03bc1g1d (d\u22121)(d+1) { k2 \u22a5u\u22a5 j (k\u0303) [ 7 64 (1+\u03b1)\u22123/2+ 1 16 \u2212 \u221a \u03bcx\u03bc 3 1A(\u03bcL, \u03bc1) \u2212 \u221a \u03bcx\u03bc 3 1A(\u03bc1, \u03bcL ) \u2212 2 \u221a \u03bcx\u03bc 3 1\u03bcLB(\u03bc1, \u03bcL ) ] + k\u22a5 j k\u22a5 c u\u22a5 c (k\u0303) [ (5 \u2212 2d ) 32 (1 + \u03b1)\u22123/2 + d2 \u2212 2d \u2212 1 16 \u2212 (d2 \u2212 2d \u2212 1) \u221a \u03bcx\u03bc 3 1A(\u03bc1, \u03bcL ) \u2212(1 \u2212 d ) \u221a \u03bcx\u03bc 3 1A(\u03bcL, \u03bc1) + 2(d \u2212 1) \u221a \u03bcx\u03bc 3 1\u03bcLB(\u03bc1, \u03bcL ) ]} . (B16) c. Graph in Fig. 3(c) The graph in Fig. 3(c) gives a contribution (\u2202t u< j )\u03bc,c to the EOM for u< j (k\u0303): (\u2202t u < j )\u03bc,c = 2\u03bb2Du\u22a5 u (k\u0303) (2\u03c0 )d+1 \u222b > q\u0303 ( k\u22a5 i \u2212 q\u22a5 i ) q\u22a5 u Ci (q\u0303)Gj (k\u0303 \u2212 q\u0303) \u2261 2\u03bb2Du\u22a5 u (k\u0303) [( I\u03bc,c 1 ) ju (k\u0303) + (I\u03bc,c 2 ) ju (k\u0303) ] , (B17) where ( I\u03bc,c 1 ) ju(k\u0303) \u2261 k\u22a5 i (2\u03c0 )d+1 \u222b > q\u0303 q\u22a5 u Ci (q\u0303)Gj (k\u0303 \u2212 q\u0303), (B18) ( I\u03bc,c 2 ) ju(k\u0303) \u2261 \u2212 1 (2\u03c0 )d+1 \u222b > q\u0303 q\u22a5 i q\u22a5 u Ci (q\u0303)Gj (k\u0303 \u2212 q\u0303). (B19) Let us calculate (I\u03bc,c 1 ) ju first. Again, since there is already a factor k\u22a5 i outside the integral, we can set \u03c9 = 0 in the integrand", " Therefore, we have ( I\u03bc,c 1 ) ju(k\u0303) = k\u22a5 i (2\u03c0 )d+1 \u222b > q\u0303 q\u22a5 u [|GL(q\u0303)|2L\u22a5 i (q) + |GT (q\u0303)|2P\u22a5 i (q) ]{ GL(\u2212q\u0303) (2L\u22a5 j (q)q\u22a5 mk\u22a5 m \u2212 k\u22a5 j q\u22a5 \u2212 k\u22a5 q\u22a5 j ) q2 \u22a5 + GL(\u2212q\u0303)2L\u22a5 j (q) ( 2\u03bcLq\u22a5 mk\u22a5 m + 2\u03bcxqxkx )\u2212 GT (\u2212q\u0303) (2L\u22a5 j (q)q\u22a5 mk\u22a5 m \u2212 k\u22a5 j q\u22a5 \u2212 k\u22a5 q\u22a5 j ) q2 \u22a5 + GT (\u2212q\u0303)2P\u22a5 j (q) ( 2\u03bc1q\u22a5 mk\u22a5 m + 2\u03bcxqxkx )} 022610-29 = k\u22a5 i (2\u03c0 )d+1 \u222b > q\u0303 q\u22a5 u { |GL(q\u0303)|2 [ GL(\u2212q\u0303) (2L\u22a5 i j (q)q\u22a5 mk\u22a5 m \u2212 k\u22a5 j q\u22a5 i \u2212 L\u22a5 i (q)k\u22a5 q\u22a5 j ) q2 \u22a5 + GL(\u2212q\u0303)2L\u22a5 i j (q) ( 2\u03bcLq\u22a5 mk\u22a5 m )\u2212 GT (\u2212q\u0303) (2L\u22a5 i j (q)q\u22a5 mk\u22a5 m \u2212 k\u22a5 j q\u22a5 i \u2212 k\u22a5 q\u22a5 j L\u22a5 i (q)) q2 \u22a5 ] + |GT (q\u0303)|2) [ \u2212 GL(\u2212q\u0303) k\u22a5 q\u22a5 j P\u22a5 i (q) q2 \u22a5 + GT (\u2212q\u0303) k\u22a5 q\u22a5 j P\u22a5 i (q) q2 \u22a5 + GT (\u2212q\u0303)2P\u22a5 i j (q) ( 2\u03bc1q\u22a5 mk\u22a5 m )]} = k\u22a5 i (2\u03c0 )d\u22121 \u222b > q\u22a5 q\u22a5 u q5 \u22a5 \u23a7\u23a8 \u23a9 (2L\u22a5 i j (q)q\u22a5 mk\u22a5 m \u2212 k\u22a5 j q\u22a5 i \u2212 L\u22a5 i (q)k\u22a5 q\u22a5 j ) 16 \u221a \u03bcx\u03bc 3 L + 3 ( \u03bcLq\u22a5 mk\u22a5 m ) L\u22a5 i j (q) 64 \u221a \u03bcx\u03bc 5 L \u2212 [2L\u22a5 i j (q)q\u22a5 mk\u22a5 m \u2212 k\u22a5 j q\u22a5 i \u2212 k\u22a5 q\u22a5 j L\u22a5 i (q) ] A(\u03bcL, \u03bc1) \u2212 k\u22a5 q\u22a5 j P\u22a5 i (q)A(\u03bc1, \u03bcL ) + k\u22a5 q\u22a5 j P\u22a5 i (q) 16 \u221a \u03bcx\u03bc 3 1 + 3P\u22a5 i j (q) ( \u03bc1q\u22a5 mk\u22a5 m ) 64 \u221a \u03bcx\u03bc 5 1 \u23ab\u23ac \u23ad = Kd d\u22124d k\u22a5 i (d \u2212 1)(d + 1) \u23a7\u23a8 \u23a9 ( \u22a5 mi juk\u22a5 m \u2212 (d + 1)\u03b4iuk\u22a5 j ) 16 \u221a \u03bcx\u03bc 3 L + 3\u03bcL \u22a5 mi juk\u22a5 m 64 \u221a \u03bcx\u03bc 5 L \u2212 [ \u22a5 mi juk\u22a5 m \u2212 (d + 1)\u03b4iuk\u22a5 j ] A(\u03bcL, \u03bc1) + (\u2212 \u22a5 mi juk\u22a5 m + (d + 1)\u03b4\u22a5 juk\u22a5 i ) \u23a1 \u23a3\u2212A(\u03bc1, \u03bcL ) + 1 16 \u221a \u03bcx\u03bc 3 1 \u23a4 \u23a6+ (\u2212 \u22a5 mi juk\u22a5 m + (d + 1)\u03b4\u22a5 i j k \u22a5 u ) 3 64 \u221a \u03bcx\u03bc 3 1 \u23ab\u23ac \u23ad = Kd d\u22124d (d \u2212 1)(d + 1) \u23a7\u23a8 \u23a9k2 \u22a5\u03b4\u22a5 ju \u23a1 \u23a3 7 64 \u221a \u03bcx\u03bc 3 L \u2212 A(\u03bcL, \u03bc1) \u2212 dA(\u03bc1, \u03bcL ) + (4d \u2212 3) 64 \u221a \u03bcx\u03bc 3 1 \u23a4 \u23a6 + k\u22a5 j k\u22a5 u \u23a1 \u23a3 (5 \u2212 2d ) 32 \u221a \u03bcx\u03bc 3 L + (d \u2212 1)A(\u03bcL, \u03bc1) + 2A(\u03bc1, \u03bcL ) + (3d \u2212 11) 64 \u221a \u03bcx\u03bc 3 1 \u23a4 \u23a6 \u23ab\u23ac \u23ad. (B20) Now we turn to (I\u03bc,c 2 ) ju(k\u0303). In Appendix B 1 d we show that the sum of (I\u03bc,c 2 ) ju(k\u0303) and (I\u03bc,d ) ju(k\u0303), which is introduced in evaluating Fig. 3(d), is at most of O(k2). This means we can set \u03c9 = 0 and focus on the O(k2) part when evaluating (I\u03bc,c 2 ) ju(k\u0303) and (I\u03bc,d ) ju(k\u0303), since their lower-order parts [i.e., O(1) and O(k)] all cancel out. We will use this knowledge in the following calculations. ( I\u03bc,c 2 ) ju(k\u0303) = \u2212 1 (2\u03c0 )d+1 \u222b > q\u0303 q\u22a5 u q\u22a5 GL(q\u0303)GL(\u2212q\u0303) { \u2212(\u03bcLk2 \u22a5 + \u03bcxk2 x ) GL(\u2212q\u0303)2L\u22a5 j (q) + (4\u03bc2 L(q\u22a5 \u00b7 k\u22a5)2 + 4\u03bc2 xq2 x k2 x ) \u00d7 GL(\u2212q\u0303)3L\u22a5 j (q) + { L\u22a5 j (q) [ \u2212 k2 \u22a5 q2 \u22a5 + 4(q\u22a5 \u00b7 k\u22a5)2 q4 \u22a5 ] + k\u22a5 j k\u22a5 q2 \u22a5 \u2212 2q\u22a5 \u00b7 k\u22a5 ( k\u22a5 j q\u22a5 + k\u22a5 q\u22a5 j ) q4 \u22a5 } GL(\u2212q\u0303) + 2\u03bcLq\u22a5 \u00b7 k\u22a5GL(\u2212q\u0303)2 ( 2L\u22a5 j (q)q\u22a5 \u00b7 k\u22a5 \u2212 k\u22a5 j q\u22a5 \u2212 k\u22a5 q\u22a5 j q2 \u22a5 ) \u2212 [ L\u22a5 j (q) ( \u2212 k2 \u22a5 q2 \u22a5 + 4(q\u22a5 \u00b7 k\u22a5)2 q4 \u22a5 ) + k\u22a5 j k\u22a5 q2 \u22a5 \u2212 2q\u22a5 \u00b7 k\u22a5 ( k\u22a5 j q\u22a5 + k\u22a5 q\u22a5 j ) q4 \u22a5 ] GT (\u2212q\u0303) \u22122\u03bc1q\u22a5 \u00b7 k\u22a5GT (\u2212q\u0303)2 ( 2L\u22a5 j (q)q\u22a5 \u00b7 k\u22a5 \u2212 k\u22a5 j q\u22a5 \u2212 k\u22a5 q\u22a5 j q2 \u22a5 )} , (B21) 022610-30 where we have discarded terms linear in qx since the integration of these terms gives 0", " Performing the and qx integrals in (B21), as described in detail in Appendix (C), we have ( I\u03bc,c 2 ) ju(k\u0303) = \u2212 1 (2\u03c0 )d\u22121 \u222b > q\u22a5 q\u22a5 u q\u22a5 q5 \u22a5 { \u2212 3\u03b4 j ( \u03bcLk2 \u22a5 + \u03bcxk2 x ) 128 \u221a \u03bcx\u03bc 5 L + ( 5\u03bc2 Lq\u22a5 mq\u22a5 n k\u22a5 m k\u22a5 n + \u03bcx\u03bcLk2 x q2 \u22a5 ) \u03b4 j 128 \u221a \u03bcx\u03bc 7 Lq2 \u22a5 + [ \u03b4 j ( \u2212k2 \u22a5 + 4(q\u22a5 \u00b7 k\u22a5)2 q2 \u22a5 ) + k\u22a5 j k\u22a5 \u2212 2q\u22a5 \u00b7 k\u22a5 ( k\u22a5 j q\u22a5 + k\u22a5 q\u22a5 j ) q2 \u22a5 ] 1 16 \u221a \u03bcx\u03bc 3 L + 3\u03bcLq\u22a5 \u00b7 k\u22a5 64 \u221a \u03bcx\u03bc 5 L ( 2\u03b4 j q\u22a5 \u00b7 k\u22a5 \u2212 k\u22a5 j q\u22a5 \u2212 k\u22a5 q\u22a5 j q2 \u22a5 ) \u2212 [ \u03b4 j ( \u2212k2 \u22a5 + 4(q\u22a5 \u00b7 k\u22a5)2 q2 \u22a5 ) + k\u22a5 j k\u22a5 \u2212 2q\u22a5 \u00b7 k\u22a5 ( k\u22a5 j q\u22a5 + k\u22a5 q\u22a5 j ) q2 \u22a5 ] A(\u03bcL, \u03bc1) \u2212 2\u03bc1q\u22a5 \u00b7 k\u22a5B(\u03bcL, \u03bc1) ( 2\u03b4 j q\u22a5 \u00b7 k\u22a5 \u2212 k\u22a5 j q\u22a5 \u2212 k\u22a5 q\u22a5 j q2 \u22a5 )} = \u2212 Kd d\u22124d (d \u2212 1)(d + 1) \u23a7\u23a8 \u23a9\u22123(d + 1)\u03b4\u22a5 ju ( \u03bcLk2 \u22a5 + \u03bcxk2 x ) 128 \u221a \u03bcx\u03bc 5 L + 5\u03bc2 L \u22a5 mn juk\u22a5 m k\u22a5 n + (d + 1)\u03b4\u22a5 ju\u03bcx\u03bcLk2 x 128 \u221a \u03bcx\u03bc 7 L + (\u2212(d + 1)\u03b4\u22a5 juk2 \u22a5 + 2 \u22a5 mn juk\u22a5 m k\u22a5 n \u2212 (d + 1)k\u22a5 j k\u22a5 u )\u239b\u239d 1 16 \u221a \u03bcx\u03bc 3 L \u2212 A(\u03bcL, \u03bc1) \u239e \u23a0 + ( \u22a5 mn juk\u22a5 m k\u22a5 n \u2212 (d + 1)k\u22a5 j k\u22a5 u )\u239b\u239d 3\u03bcL 64 \u221a \u03bcx\u03bc 5 L \u2212 2\u03bc1B(\u03bcL, \u03bc1) \u239e \u23a0 \u23ab\u23ac \u23ad = \u2212 Kd d\u22124d (d \u2212 1)(d + 1) \u23a7\u23a8 \u23a9\u2212k2 x \u03b4 \u22a5 ju (d + 1) \u221a \u03bcx 64 \u221a \u03bc5 L + k2 \u22a5\u03b4\u22a5 ju \u23a1 \u23a3 (16 \u2212 11d ) 128 \u221a \u03bcx\u03bc 3 L + (d \u2212 1)A(\u03bcL, \u03bc1) \u2212 2\u03bc1B(\u03bcL, \u03bc1) \u23a4 \u23a6 + k\u22a5 j k\u22a5 u \u23a1 \u23a3 (20 \u2212 7d ) 64 \u221a \u03bcx\u03bc 3 L + (d \u2212 3)A(\u03bcL, \u03bc1) + 2(d \u2212 1)\u03bc1B(\u03bcL, \u03bc1) \u23a4 \u23a6 \u23ab\u23ac \u23ad. (B22) The total contribution (\u2202t u< j )\u03bc,c of the graph Fig. 3(c) to the equation of motion is therefore: (\u2202t u < j )\u03bc,c = 2\u03bb2Du\u22a5 u (k\u0303) [( I\u03bc,c 1 ) ju(k\u0303) + (I\u03bc,c 2 ) ju(k\u0303) ] = 2\u03bb2DKd d\u22124u\u22a5 u (k\u0303)d (d \u2212 1)(d + 1) \u23a7\u23a8 \u23a9k2 x \u03b4 \u22a5 ju (d + 1) \u221a \u03bcx 64 \u221a \u03bc5 L + k2 \u22a5\u03b4\u22a5 ju \u23a1 \u23a3 (11d \u2212 2) 128 \u221a \u03bcx\u03bc 3 L \u2212 dA(\u03bcL, \u03bc1) + 2\u03bc1B(\u03bcL, \u03bc1) \u2212 dA(\u03bc1, \u03bcL ) + (4d \u2212 3) 64 \u221a \u03bcx\u03bc 3 1 \u23a4 \u23a6+ k\u22a5 j k\u22a5 u \u23a1 \u23a3 (3d \u2212 10) 64 \u221a \u03bcx\u03bc 3 L + 2A(\u03bcL, \u03bc1) + 2(1 \u2212 d )\u03bc1B(\u03bcL, \u03bc1)2A(\u03bc1, \u03bcL ) + (3d \u2212 11) 64 \u221a \u03bcx\u03bc 3 1 \u23a4 \u23a6 \u23ab\u23ac \u23ad = 2\u03bc1g1d (d \u2212 1)(d + 1) \u23a7\u23a8 \u23a9k2 x u\u22a5 j (k\u0303) (d + 1) \u221a \u03bcx 64 \u221a \u03bc5 L + k2 \u22a5u\u22a5 j (k\u0303) [ (11d \u2212 2) 128 (1 + \u03b1)\u22123/2 \u2212 d \u221a \u03bcx\u03bc 3 1A(\u03bcL, \u03bc1) +2 \u221a \u03bcx\u03bc 5 1B(\u03bcL, \u03bc1) \u2212 d \u221a \u03bcx\u03bc 3 1A(\u03bc1, \u03bcL ) + (4d \u2212 3) 64 ] + k\u22a5 j k\u22a5 u u\u22a5 u (k\u0303) [ (3d \u2212 10) 64 (1 + \u03b1)\u22123/2 + 2 \u221a \u03bcx\u03bc 3 1A(\u03bcL, \u03bc1) + 2(1 \u2212 d ) \u221a \u03bcx\u03bc 5 1B(\u03bcL, \u03bc1) + 2 \u221a \u03bcx\u03bc 3 1A(\u03bc1, \u03bcL ) + (3d \u2212 11) 64 ]\u23ab\u23ac \u23ad. (B23) 022610-31 d. Graph in Fig. 3(d) The graph in Fig. 3(d) gives a contribution (\u2202t u< j )\u03bc,d to the EOM for u< j (k\u0303): (\u2202t u < j )\u03bc,d = 2\u03bb2Du\u22a5 u (k\u0303) (2\u03c0 )d+1 \u222b > q\u0303 q\u22a5 i q\u22a5 u Cj (q\u0303)Gi (k\u0303 \u2212 q\u0303) \u2261 2\u03bb2Du\u22a5 u (k\u0303)(I\u03bc,d ) ju(k\u0303), (B24) where (I\u03bc,d ) ju(k\u0303) \u2261 1 (2\u03c0 )d+1 \u222b > q\u0303 q\u22a5 i q\u22a5 u Cj (q\u0303)Gi (k\u0303 \u2212 q\u0303). (B25) First let us show that the sum of (I\u03bc,c 2 ) ju(k\u0303) and (I\u03bc,d ) ju(k\u0303) is at most of O(k2). The integrand of (I\u03bc,c 2 ) ju(k\u0303) can be rewritten as \u2212q\u22a5 i q\u22a5 u Ci (q\u0303)Gj (k\u0303 \u2212 q\u0303) = \u2212q\u22a5 u |GL(q\u0303)|2q\u22a5 Gj (k\u0303 \u2212 q\u0303) = \u2212q\u22a5 u |GL(q\u0303)|2(q\u22a5 \u2212 k\u22a5 ) Gj (k\u0303 \u2212 q\u0303) \u2212 q\u22a5 u k\u22a5 |GL(q\u0303)|2Gj (k\u0303 \u2212 q\u0303) = \u2212q\u22a5 u |GL(q\u0303)|2(q\u22a5 j \u2212 k\u22a5 j ) GL(k\u0303 \u2212 q\u0303) \u2212 q\u22a5 u k\u22a5 |GL(q\u0303)|2Gj (k\u0303 \u2212 q\u0303) = \u2212q\u22a5 u q\u22a5 j |GL(q\u0303)|2GL(k\u0303 \u2212 q\u0303) + k\u22a5 j q\u22a5 u |GL(q\u0303)|2GL(k\u0303 \u2212 q\u0303) \u2212 k\u22a5 q\u22a5 u |GL(q\u0303)|2Gj (k\u0303 \u2212 q\u0303)", " Again we will omit terms linear in qx since the integration of these terms gives 0, (I\u03bc,d ) ju(k\u0303) = 1 (2\u03c0 )d+1 \u222b > q\u0303 q\u22a5 i q\u22a5 u { |GL(q\u0303)|2L\u22a5 j (q) [ \u2212 (\u03bcLk2 \u22a5 + \u03bcxk2 x ) GL(\u2212q\u0303)2L\u22a5 i (q) + (4\u03bc2 L(q\u22a5 \u00b7 k\u22a5)2 + 4\u03bc2 xqxk2 x ) \u00d7GL(\u2212q\u0303)3L\u22a5 i (q) + [ L\u22a5 i (q) [ \u2212 k2 \u22a5 q2 \u22a5 + 4(q\u22a5 \u00b7 k\u22a5)2 q4 \u22a5 ] + k\u22a5 i k\u22a5 q2 \u22a5 \u2212 2q\u22a5 \u00b7 k\u22a5 ( k\u22a5 i q\u22a5 + k\u22a5 q\u22a5 i ) q4 \u22a5 ] \u00d7 [GL(\u2212q\u0303) \u2212 GT (\u2212q\u0303)] + 2\u03bcLq\u22a5 \u00b7 k\u22a5GL(\u2212q\u0303)2 [ 2L\u22a5 i (q)q\u22a5 \u00b7 k\u22a5 \u2212 k\u22a5 i q\u22a5 \u2212 k\u22a5 q\u22a5 i q2 \u22a5 ] \u22122\u03bc1q\u22a5 \u00b7 k\u22a5GT (\u2212q\u0303)2 [ 2L\u22a5 i (q)q\u22a5 \u00b7 k\u22a5 \u2212 k\u22a5 i q\u22a5 \u2212 k\u22a5 q\u22a5 i q2 \u22a5 ]] + |GT (q\u0303)|2P\u22a5 j (q) \u00d7 [ \u2212 (\u03bc1k2 \u22a5 + \u03bcxk2 x ) GT (\u2212q\u0303)2P\u22a5 i (q) \u2212 [(4\u03bc2 1(q\u22a5 \u00b7 k\u22a5)2 + 4\u03bc2 xq2 x k2 x + 4\u03bcx\u03bc1(q\u22a5 \u00b7 k\u22a5)qxkx ] \u00d7 GT (\u2212q\u0303)3P\u22a5 i (q) + ( k\u22a5 i k\u22a5 q2 \u22a5 \u2212 2q\u22a5 \u00b7 k\u22a5k\u22a5 q\u22a5 i q4 \u22a5 ) (GL(\u2212q\u0303) \u2212 GT (\u2212q\u0303)) \u22122\u03bcLq\u22a5 \u00b7 k\u22a5GL(\u2212q\u0303)2 ( k\u22a5 q\u22a5 i q2 \u22a5 ) + 2\u03bc1q\u22a5 \u00b7 k\u22a5GT (\u2212q\u0303)2 ( k\u22a5 q\u22a5 i q2 \u22a5 )]} = 1 (2\u03c0 )d\u22121 \u222b > q\u22a5 q\u22a5 i q\u22a5 u q5 \u22a5 \u23a7\u23a8 \u23a9\u22123L\u22a5 i j (q) ( \u03bcLk2 \u22a5 + \u03bcxk2 x ) 128 \u221a \u03bcx\u03bc 5 L + ( 5\u03bc2 Lq\u22a5 mq\u22a5 n k\u22a5 m k\u22a5 n + \u03bcx\u03bcLk2 x q2 \u22a5 ) L\u22a5 i j (q) 128 \u221a \u03bcx\u03bc 7 Lq2 \u22a5 + [ L\u22a5 i j (q) [ \u2212k2 \u22a5 + 4(q\u22a5 \u00b7 k\u22a5)2 q2 \u22a5 ] + k\u22a5 i k\u22a5 L\u22a5 j (q) \u2212 2q\u22a5 \u00b7 k\u22a5 [ k\u22a5 i q\u22a5 j + k\u22a5 q\u22a5 i L\u22a5 j (q) ] q2 \u22a5 ] \u00d7 \u23a1 \u23a3 1 16 \u221a \u03bcx\u03bc 3 L \u2212 A(\u03bcL, \u03bc1) \u23a4 \u23a6+ 2q\u22a5 \u00b7 k\u22a5 \u23a1 \u23a3 3\u03bcL 128 \u221a \u03bcx\u03bc 5 L \u2212 \u03bc1B(\u03bcL, \u03bc1) \u23a4 \u23a6 022610-32 \u00d7 [ 2L\u22a5 i j (q)q\u22a5 \u00b7 k\u22a5 \u2212 k\u22a5 i q\u22a5 j \u2212 k\u22a5 q\u22a5 i L\u22a5 j (q) q2 \u22a5 ]] + P\u22a5 j (q) ( k\u22a5 i k\u22a5 \u2212 2q\u22a5 \u00b7 k\u22a5k\u22a5 q\u22a5 i q2 \u22a5 ) \u00d7 \u23a1 \u23a3A(\u03bc1, \u03bcL ) \u2212 1 16 \u221a \u03bcx\u03bc 3 1 \u23a4 \u23a6\u2212 2q\u22a5 \u00b7 k\u22a5 \u23a1 \u23a3\u03bcLB(\u03bc1, \u03bcL ) \u2212 3\u03bc1 128 \u221a \u03bcx\u03bc 5 1 \u23a4 \u23a6(k\u22a5 q\u22a5 i q2 \u22a5 ) P\u22a5 j (q) \u23ab\u23ac \u23ad = Kd d\u22124d (d \u2212 1)(d + 1) \u23a7\u23a8 \u23a9\u22123\u03b4\u22a5 ju(d + 1) ( \u03bcLk2 \u22a5 + \u03bcxk2 x ) 128 \u221a \u03bcx\u03bc 5 L + 5\u03bc2 L \u22a5 mn juk\u22a5 m k\u22a5 n + \u03bcx\u03b4 \u22a5 ju(d + 1)\u03bcLk2 x 128 \u221a \u03bcx\u03bc 7 L + (\u2212\u03b4\u22a5 ju(d + 1)k2 \u22a5 + \u22a5 mn juk\u22a5 m k\u22a5 n )\u23a1\u23a3 1 16 \u221a \u03bcx\u03bc 3 L \u2212 A(\u03bcL, \u03bc1) \u23a4 \u23a6+ (\u2212(d + 1)k\u22a5 u k\u22a5 j + \u22a5 mn juk\u22a5 m k\u22a5 n ) \u00d7 \u23a1 \u23a3A(\u03bc1, \u03bcL ) \u2212 1 16 \u221a \u03bcx\u03bc 3 1 \u23a4 \u23a6\u2212 2 ( (d + 1)k\u22a5 u k\u22a5 j \u2212 \u22a5 mn juk\u22a5 m k\u22a5 n )\u23a1\u23a3\u03bcLB(\u03bc1, \u03bcL ) \u2212 3 128 \u221a \u03bcx\u03bc 3 1 \u23a4 \u23a6 \u23ab\u23ac \u23ad. (B28) The total contribution (\u2202t u< j )\u03bc,d of the graph Fig. 3(d) to the equation of motion is therefore: (\u2202t u < j )\u03bc,d = 2\u03bb2Du\u22a5 u (k\u0303)(I\u03bc,d ) ju(k\u0303) = 2\u03bb2Du\u22a5 u (k\u0303)Kd d\u22124d (d \u2212 1)(d + 1) \u23a7\u23a8 \u23a9\u2212\u03b4\u22a5 juk2 x (d + 1) \u221a \u03bcx 64 \u221a \u03bc5 L + \u03b4\u22a5 juk2 \u22a5 \u23a1 \u23a3 (2 \u2212 11d ) 128 \u221a \u03bcx\u03bc 3 L + dA(\u03bcL, \u03bc1) + A(\u03bc1, \u03bcL ) \u2212 7 64 \u221a \u03bcx\u03bc 3 1 + 2\u03bcLB(\u03bc1, \u03bcL ) \u23a4 \u23a6+ k\u22a5 u k\u22a5 j \u23a1 \u23a3 13 64 \u221a \u03bcx\u03bc 3 L \u2212 2A(\u03bcL, \u03bc1) + (1 \u2212 d )A(\u03bc1, \u03bcL ) + (7d \u2212 7) 64 \u221a \u03bcx\u03bc 3 1 + 2(1 \u2212 d )\u03bcLB(\u03bc1, \u03bcL ) \u23a4 \u23a6 \u23ab\u23ac \u23ad = 2\u03bc1g1d (d \u2212 1)(d + 1) \u23a7\u23a8 \u23a9\u2212k2 x u\u22a5 j (k\u0303) (d + 1) \u221a \u03bcx 64 \u221a \u03bc5 L + k2 \u22a5u\u22a5 j (k\u0303) [ (2 \u2212 11d ) 128 (1 + \u03b1)\u22123/2 + d \u221a \u03bcx\u03bc 3 1A(\u03bcL, \u03bc1) + \u221a \u03bcx\u03bc 3 1A(\u03bc1, \u03bcL ) \u2212 7 64 + 2 \u221a \u03bcx\u03bc 3 1\u03bcLB(\u03bc1, \u03bcL ) ] + k\u22a5 j k\u22a5 u u\u22a5 u (k\u0303) [ 13 64 (1 + \u03b1)\u22123/2 \u2212 2 \u221a \u03bcx\u03bc 3 1A(\u03bcL, \u03bc1) + (1 \u2212 d ) \u221a \u03bcx\u03bc 3 1A(\u03bc1, \u03bcL ) + 7(d \u2212 1) 64 + 2(1 \u2212 d ) \u221a \u03bcx\u03bc 3 1\u03bcLB(\u03bc1, \u03bcL ) ]\u23ab\u23ac \u23ad. (B29) 2. Overall propagator renormalization Summing over all the contributions from the one-loop diagrams in Fig. 3, we find two types of terms: k2 \u22a5u\u22a5 j (k\u0303) and k\u22a5 j k\u22a5 u u\u22a5 u (k\u0303). The sum of the coefficients of the former gives the correction to \u2212\u03bc1; that of the latter gives the correction to \u2212\u03bc2. There is no correction to \u03bcx to one-loop order, since no terms proportional to k2 x u j survive to this order. Thus the graphical correction \u03b4\u03bc1 to \u03bc1 is \u03b4\u03bc1 = 2\u03bc1g1d (d \u2212 1)(d + 1) {( \u2212 7 64 (1 + \u03b1)\u22123/2 + (d2 \u2212 2d \u2212 2) 16 + (d + 2) \u221a \u03bcx\u03bc 3 1A(\u03bcL, \u03bc1) + d \u221a \u03bcx\u03bc 3 1A(\u03bc1, \u03bcL ) \u22122d \u221a \u03bcx\u03bc 5 1B(\u03bcL, \u03bc1) ) + [ 7 64 (1 + \u03b1)\u22123/2 + 1 16 \u2212 \u221a \u03bcx\u03bc 3 1A(\u03bcL, \u03bc1) \u2212 \u221a \u03bcx\u03bc 3 1A(\u03bc1, \u03bcL ) \u2212 2 \u221a \u03bcx\u03bc 3 1\u03bcLB(\u03bc1, \u03bcL ) ] \u2212 [ (11d \u2212 2) 128 (1 + \u03b1)\u22123/2 \u2212 d \u221a \u03bcx\u03bc 3 1A(\u03bcL, \u03bc1) + 2 \u221a \u03bcx\u03bc 5 1B(\u03bcL, \u03bc1) \u2212 d \u221a \u03bcx\u03bc 3 1A(\u03bc1, \u03bcL ) + (4d \u2212 3) 64 ] 022610-33 \u2212 [ (2 \u2212 11d ) 128 (1 + \u03b1)\u22123/2 + d \u221a \u03bcx\u03bc 3 1A(\u03bcL, \u03bc1) + \u221a \u03bcx\u03bc 3 1A(\u03bc1, \u03bcL ) \u2212 7 64 + 2 \u221a \u03bcx\u03bc 3 1\u03bcLB(\u03bc1, \u03bcL ) ]} = 2\u03bc1g1d (d \u2212 1)(d + 1) { (2d2 \u2212 6d + 3) 32 + (1 + d ) \u221a \u03bcx\u03bc 3 1A(\u03bcL, \u03bc1) + 2(d \u2212 1) \u221a \u03bcx\u03bc 3 1A(\u03bc1, \u03bcL ) \u22122(d + 1) \u221a \u03bcx\u03bc 5 1B(\u03bcL, \u03bc1) \u2212 4 \u221a \u03bcx\u03bc 3 1\u03bcLB(\u03bc1, \u03bcL ) } , (B30) where \u221a \u03bcx\u03bc 3 1A(\u03bcL, \u03bc1) = 1 4\u03b1 ( \u221a 2\u221a 2 + \u03b1 \u2212 1\u221a 1 + \u03b1 ) , (B31) \u221a \u03bcx\u03bc 3 1A(\u03bc1, \u03bcL ) = 1 4\u03b1 ( 1 \u2212 \u221a 2\u221a 2 + \u03b1 ) , (B32) \u221a \u03bcx\u03bc 5 1B(\u03bcL, \u03bc1) = 2(2 + \u03b1)3/2 \u2212 \u221a 2(1 + \u03b1)(4 + \u03b1) 8 \u221a 1 + \u03b1(2 + \u03b1)3/2\u03b12 , (B33) \u221a \u03bcx\u03bc 3 1\u03bcLB(\u03bc1, \u03bcL ) = (1 + \u03b1) 2(2 + \u03b1)3/2 \u2212 \u221a 2(4 + 3\u03b1) 8(2 + \u03b1)3/2\u03b12 " ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001895_tcst.2020.2989691-Figure9-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001895_tcst.2020.2989691-Figure9-1.png", "caption": "Fig. 9. Motor dq voltage inputs with the proposed BIC for the vector control of a PMSM. (a) In the time domain. (b) Moving on a controller sphere.", "texts": [ " Initially, the electromagnetic torque reference is set to 1 N \u00b7m by passing through a first-order low-pass filter. The electromagnetic torque of the PMSM can track its reference well, as shown in Fig. 8(a), where the mean absolute error (MAE) of the electromagnetic torque tracking during 2.9 \u223c 3 s is about 0.0019 N \u00b7 m. The motor speed converges to its steady state within about 1.5 s, as shown in Fig. 8(b). The motor currents are shown in Fig. 8(c), wherein id is well regulated to 0, and iq is regulated to a positive number to generate the positive torque. The motor voltages are shown in Fig. 9(a) with the negative ud and the positive uq . At t = 3 s, the torque reference is changed to 1.5 N \u00b7 m, and (u2 d +u2 q) 1/2 increases to the maximum value quickly. Though the electromagnetic torque cannot be regulated because of the voltage limit, the BIC still regulates (u2 d + u2 q) 1/2 within the limit, as shown in Fig. 9(a), and the closed-loop system still keeps stable operation. When the torque reference is set to 0.5 N \u00b7 m at t = 6 s, both electromagnetic torque and motor currents are well regulated again, due to the decrease in (u2 d + u2 q) 1/2, where the MAE of the electromagnetic torque tracking during 8.8 \u223c 9 s is about 0.0036 N \u00b7m. The motor speed decreases with a lower torque. The BIC can always keep (u2 d + u2 q) 1/2 \u2264 1 at different stages and guarantee system stability. The controller states ud , uq , and u0 are illustrated in Fig. 9(b), where the controller states remain on a controller sphere Sr with the proposed BIC (6). The small neighborhoods of the equilibrium points E1, E2, and E3 represent three steady states of the system, during 2 \u223c 3 s, Authorized licensed use limited to: University of Exeter. Downloaded on June 10,2020 at 11:51:34 UTC from IEEE Xplore. Restrictions apply. 5 \u223c 6 s, and 8 \u223c 9 s, respectively. Therefore, the BIC can achieve good reference tracking and handle limited control power with time-invariant input weights in this experimental study" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0003137_1.1752269-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0003137_1.1752269-Figure3-1.png", "caption": "FIG. 3. Curve of motion of the pendulum.", "texts": [ " By starting from the equation of oscillation the method and its results are particularly easy to represent. The second section gives a rigorous treatment of the problem. When there is no viscous friction we have: end of a quarter period, brought about by friction damping. That is to say: 2ao is the shift of the zero line of the sinusoidal oscillation of the pendulum at the end of a half-period. Observing unilateral elongations, i.e., full periods, and de noting the first amplitude with A o, this decrease of the amplitude at the end of a full period is 4ao. Fig. 3 demonstrates the process of decrease, starting at a time, where the amplitude is Xo=Ao-ao. The first sinusoidal half-period is constructed on the line 1 as a basis in the distance ao above the zero line. At Xl the second half period starts with the amplitude X1=Xo-2ao, constructed on the basis line 2 at the distance ao below the zero line, etc. Observing full periods, the values of consecu tive amplitudes are: 0.: Ao; 1.: Ao-4ao; 2.: Ao-8ao; 3.: Ao-12ao ... n.: Ao-4nao, where n is the number of full periods" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002035_j.measurement.2020.108224-Figure8-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002035_j.measurement.2020.108224-Figure8-1.png", "caption": "Fig. 8. Stress map of centrifugal fan.", "texts": [ " The upper surface of the buckle teeth at the lower end of the blade is fixed in contacting with the lower surface of the rear disc. The circumferential displacement of the inner hole surface of the hub is restrained. When the blade rotates, it is mainly affected by centrifugal force and aerodynamic force. Reference shows that the centrifugal force is much greater than the aerodynamic force [19]. Therefore, this paper mainly focuses on the influence of centrifugal force of the blade. The stress results can be seen from following figure. As shown in Fig. 8, the maximum stress is 170.96 MPa that appears at the root of the outer corner of the blade. Because the maximum stress is close to two thirds of the yield limit and its area is very small, it doesn\u2019t cause damage in this situation. It can be preliminarily inferred that when the fan blade is fatigued, the most likely position to initiate cracks is the root of blade. The static strength analysis of centrifugal fan blades provides basic assurance for blade failure analysis later. According to the simulation analysis results, the location of measuring point should be the area with large stress and deformation", " Before the formal experiment, preloading collection has been down by 1\u20133 times. The sampling mode is continuous sampling and the rate is set to 8192 Hz. The speed is 1440 rpm. The test condition is the same as that of the simulation. There are 1 min for each working condition that repeat 3 times. The average stress experimental results at the rated speed are shown in Table 6 according to Eq. (1). The values are between 50 MPa and 65 MPa. There are some reasons for differences between simulation and experiment as compared in Fig. 8 with Table 6, such as inconsistency between the finite element model and real structure, the difference between the actual and analytical loads, and measurement and calculation errors. Especially when the mesh size is further decreased, the original maximum calculated stress will further increase. This result proves that the exact position of strain gauge Table 5 Nameplate parameters of centrifugal fan. Motor power (kW) Static pressure (Pa) Rated speed (r/min) Flow (m3\u2215h) Medium 4.0 960 1440 3000 Air is in the light green position of Fig. 8b. The measured value is closer to the real level. The following analysis is based on the test results. In the same process, we obtain stresses as Table 7 at different and increased speeds. The stress increases with speed obviously. The maximum speed is limited to 2200 rpm for the excessive noise and the ability of the transmission shaft. At the same time, it also meets the requirements of the test in the next section. According to the Table 7, the average stress is 55.4169 MPa at the rated speed" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002426_j.engfailanal.2020.105211-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002426_j.engfailanal.2020.105211-Figure4-1.png", "caption": "Fig. 4. Wheel-hub-bearing force diagram.", "texts": [ " The load history of the original load spectrum data collected by the data acquisition instrument after preprocessing, such as separation and extraction, zero drift removal, singular point elimination, filtering and trend term removal, is shown in Fig. 3. Mx is the driving torque of the wheel. The hub unit in this paper is located in the front wheel of the vehicle, so there is no driving moment. My is offset by the brake torque. Mz is the aligning torque of the wheel, its value and frequency are low, so it can be ignored. The radial load Fr is the resultant force in the direction of Fx and Fz, and the axial load Fa is the load in the direction of Fy. The actual force analysis of wheel hub bearing is shown in Fig. 4. Due to the hub unit structure, the load center is located to the left of inner bearing 1. The radial load of wheel hub bearing can be derived from the principle of force balance { Fr1 = Fra/L Fr2 = Fr + Fra/L (1) S1 and S2 are axial derived forces generated by radial forces Fr1 and Fr2 respectively { S1 = Fr1/2Y1 S2 = Fr2/2Y2 (2) The values of Y1, Y2 in Equation (2) should be taken when Fa/Fr > e. e is the judgment coefficient. When Fa + S2 > S1, then L.-H. Zhao et al. Engineering Failure Analysis 122 (2021) 105211 L" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000379_1.4033100-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000379_1.4033100-Figure1-1.png", "caption": "Fig. 1 Scheme size of the rollers and application of the sanding: (a) scheme size of the rollers and (b) diagrammatic sketch of the sanding", "texts": [ " The geometric sizes of rollers are determined by means of Hertzian rule, as is shown in the following equations: \u00f0q0\u00delab \u00bc \u00f0q0\u00defield (1) \u00f0a=b\u00delab \u00bc \u00f0a=b\u00defield (2) where (q0)lab and (q0)field are the maximum contact stresses in the laboratory and in the field, respectively. (a/b)lab and (a/b)field are the ratios of semimajor axis of the contact ellipses between the wheel and rail in the laboratory and field, respectively. The scheme of geometric size calculated by the above equation is shown in Fig. 1(a). The normal force of 120 N is applied on the interface between wheel and rail rollers. The rolling speed of rail roller is 400 rpm and the slippage ratio of rollers is 0.91%. The duration of each test is 120 mins. The wheel and rail rollers are made of the wheel and rail steels applied in the field (wheel: CL60, rail: U71Mn). Their chemical compositions in weight percentage are given in Table 1. The adhesion experiments are performed under the water condition and water is continuously added into the contact surfaces of wheel/rail at the flow rate of about 1.5 ml/min. This is found to 011401-2 / Vol. 139, JANUARY 2017 Transactions of the ASME Downloaded From: http://tribology.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jotre9/935372/ on 04/13/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use cause just sufficient water flow such that the disks are always wetted. Sand particles are added into the contacts of wheel/rail powered by self-gravity, as shown in Fig. 1(b). Silica sand is used in this experiment. The main composition of the sand is SiO2 and its hardness is about 7 Mohs. Sands can be divided into different particle diameters by sieves with different size. Four diameters of sand particles are used in this study: 0.05\u20130.2 mm (S sand), 0.2\u20130.3 mm (M sand), 0.3\u20130.45 mm (L sand), and 0.45\u20130.9 mm (XL sand). A photograph of sand particles is given in Fig. 2. Three different feed rates are chosen, which range from about 5 g/min to 10 g/min, 15 g/min controlled by a sanding valve" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002212_tie.2020.3031535-Figure18-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002212_tie.2020.3031535-Figure18-1.png", "caption": "Fig. 18. Comparison of flux density distributions when Id=Iq=0 and Idc=Irated.", "texts": [ "4 Torque ripple T o rq u e ri p p le ( % ) 0 3 6 9 12 15 0.0 0.2 0.4 0.6 0.8 Torque T o rq u e (N m ) w 1 (deg.) 0.0 0.1 0.2 0.3 0.4 Torque ripple T o rq u e ri p p le ( % ) Authorized licensed use limited to: University of Gothenburg. Downloaded on November 17,2020 at 09:38:58 UTC from IEEE Xplore. Restrictions apply. 0278-0046 (c) 2020 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. Fig. 18 shows the flux density distributions of the ST2, SPMM, and VFRM. The ST2 has very few flux lines in the rotor yoke compared with the other two machines. Fig. 19 compares the back EMFs of at 600rpm. The SPMM exhibits the highest fundamental harmonic, although its 5th harmonic is very obvious. When Idc=Irated, the VFRM shows the second highest back EMF. It is worth noting that the back EMF of ST2 is only around one tenth of the SPMM, due to that most of the PM flux circulates within the stator, as shown in Fig. 18(a). Fig. 20 shows that the SPMM has the highest cogging torque while the ST2 and VFRM have negligible cogging torques. On one hand, the SPMM has the highest airgap flux density. On the other hand, for the SPMM the least common multiple (LCM) of the slot number and pole number is only 12. Fig. 21 compares the torques of the three machines. The SPMM delivers the highest torque 2.36Nm and the highest torque ripple 32.7% mainly due to the highest back EMF and highest cogging torque, respectively. The ST2 produces the second highest torque, which is around half of the SPMM" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002212_tie.2020.3031535-Figure17-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002212_tie.2020.3031535-Figure17-1.png", "caption": "Fig. 17. Cross sections of optimized machines.", "texts": [ " All these three machines have the same stator slot number, outer diameter, stack length, airgap length, as well as concentrated windings. The optimized parameters of the ST2 are stated in Section III. For the SPMM and VFRM, the optimized parameters are stator inner radius r1, stator yoke width yk, stator tooth width tw, PM height hpm, tooth tip height th, slot opening so, rotor tooth width tw2, rotor slot depth h2. They are globally optimized for maximum torque under 20W copper loss, and the cross sections are shown in Fig. 17. The parameters are listed in TABLE II. It should be noted that the 6S11P ST2 shows 80% higher torque than the 12S22P ST2. This is because the armature MMF per tooth increases by 84%, and hence more PM flux can be forced into the rotor side at rated condition, which can be observed from the PM height listed in TABLEs I and II. The PM height of the 6S11P ST2 is 70% higher than that of the 12S22P ST2. 0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6 0.0 0.2 0.4 0.6 0.8 Torque T o rq u e (N m ) Remnant flux density (T) 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000718_978-1-4471-7278-9-Figure12.8-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000718_978-1-4471-7278-9-Figure12.8-1.png", "caption": "Fig. 12.8 The artificial life Strandbeest by Dutch artist Theo Jansen", "texts": [ "\u201d Combining art and science, the Dutch artist Theo Jansen created an artificial life called Strandbeests, which means \u201cbeach creatures\u201d in English. Strandbeests are wind-powered robotic sculptures made of stiff PVC pipes, water bottles and cloth. Jansan called those plastic pipes \u201cprotein\u201d because they are building blocks of the beach creature. Jansen wanted to create a new kind of life from scratch and totally different from existing creatures. He thought this might help humans to know the real form of life. Figure 12.8 is a sketch of one of Strandbeests. Strandbeest turned wind into energy. The winds sweep the air and store the energy in bottles, which can be used to provide energy for walking when the wind is gone. When the tide comes up, the robot will have the energy to walk away from Instinctive Computing 247 the tide. If there is too much wind, the robot might be blown away. So the nose is directed into the wind. The hammer will put stakes in the ground. Strandbeests have taken 20 years to evolve with the help from their inventor Jansen, and computational design tools to optimize the structure" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000539_s00170-016-9337-2-Figure28-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000539_s00170-016-9337-2-Figure28-1.png", "caption": "Fig. 28 Stock of platform surface roughing", "texts": [ " The cutter for platform surface roughing and blade surface roughing is a D32R5 torus-shaped milling cutter (shown in Fig. 27); the cutter for platform surface finishing is a R4 ball-end milling cutter. We make the post process using UGPost (NX8.0). We use platform surface roughing to verify the tool path generation algorithm and use platform surface finishing to verify the hybrid method. As described in Sect. 3.2, the platform surface is a 3D surface; it is part of the rotor/ stator surface, which is a surface of revolution. The stock for platform surface roughing is shown in Fig. 28; the platform surface (machining surface) is shown in Fig. 29. As there are both inner and outer boundaries in the machining region, and the distribution of the distance between the two boundaries is uneven, we use the longest gradient splitting method provided in Sect. 5.1 to control the step-over for platform roughing. The details of the parameters for platform surface roughing are shown in Table 2. The cutting and noncutting tool paths are shown in Fig. 30, and the machining results are shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002828_j.jii.2021.100265-Figure13-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002828_j.jii.2021.100265-Figure13-1.png", "caption": "Fig. 13. One more layer is recommended to be deposited on parent sub-volume and the torch angle should be adjusted to bond two layers.", "texts": [ " A new concept, collision matrix, is used to evaluate deposition sequences. (iii) Owing to the use of the advanced searching algorithm, the optimal deposition sequences and torch angles were employed instead of the conventional vertical-up method. The potential collisions were eliminated by the calculated deposition strategy when fabricating the part with closed structures. For practical applications, some further considerations have been summarised as follows: (i) Some parent sub-volumes, such as the one highlighted in Fig. 13, need to be extended so that the child sub-volume, which extends from it, can be supported appropriately. At this stage, these modifications are implemented manually in most of WAAM process. In the proposed system, an automated approach to algorithmically identify these scenarios Fig. 12. The manufacturing process. L. Yuan et al. Journal of Industrial Information Integration xxx (xxxx) xxx is developed based on the build direction of each sub-volume and dependency matrix D. then, the robot code can be adjusted accordingly. (ii) When reaching a junction between two (or more) sub-volumes, as shown in Fig. 13, some additional consideration is required. Experimental testing found that the connection between the two sub-volumes is physically weak. To strengthen the connection, the torch angle should be adjusted to 45\u25e6, as shown in Fig. 13, to allow for simultaneous bonding of both sub-volumes. In the developed algorithm, this junction is identified by building directions of each sub-volume, and the torch angle is set to bisect these directions. In summary, the experimental work demonstrated the successful development of a novel path planning algorithm for the multidirectional WAAM. When the software was integrated with a common WAAM hardware, the overhanging features and the closed structure, which are challenging for the conventional WAAM method, were deposited with high quality" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002281_j.jmrt.2020.11.080-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002281_j.jmrt.2020.11.080-Figure1-1.png", "caption": "Fig. 1 e Formation of a porous Ti structure on the surface of an DED printing method and (b) the diffusional bonding, direct bon", "texts": [ " However, since the elastic modulus of Ti alloys is considerably lower than that of CoeCr-based alloys [50], formation of a porous layer of Ti alloys on the surface of CoeCr-based alloys is more advantageous than formation of a porous CoeCr-based alloys on the surface of CoeCr-based alloys from the standpoint of reducing the stress-shielding effect around an implantebone interface where bone tissue he middle part of the CoeCr alloy plate. (b) The three-point specimens for the XRD measurement. and implants are connected. Accordingly, the idea of forming a porous structure using Ti on the surface of an implant made of CoeCr-based alloys using the 3D printing DED method has been proposed. The application of the DED method for forming a Ti porous structure on the CoeCr femoral component is schematically shown in Fig. 1a. Chemical bonding [51], direct bonding [52] and diffusional bonding [53] may be also considered as potential methods of making a porous Ti structure on the surface of the implant body (Fig. 1b). In these methods, first of all, a thin, porous Ti structure should be prepared. This porous structure may be easily manufactured by the PBF printing method, but there is a difficulty in separating the printed porous layer from the build plate. In addition, in order to bond the Ti porous structure to the femoral component, it is necessary to place the printed porous Ti strip to fit the surface curvature of the implant body and to find suitable surface and heat treatment conditions for high interfacial bonding", " As the scan speed increases, the external porosity increases, but the internal porosity decreases, such that the effect of scan speed on the total porosity is small. e tests observed by optical microscopy at powers of (a) 40 (b) at powers of (e) 40 (f) 50 (g) 70 and (h) 90 W at a given scan It is worthy to note that at the scan speeds of 500, 1000 and 1500 min/min with the powder feed rate of 2.3 g/min, the times taken to print the two layers of Ti powder coating on the surface of a femoral implant (with surface area of 60 mm 65 mm) shown in Fig. 1a are calculated to be 20, 10 and 7 min, respectively. Fig. 12a shows a typical microstructure from the built sample observed by optical microscopy. At the interface between the injected Ti powder and the CoeCr substrate, where the a-Ti solid and CoeCr melt pool liquid are in contact, columnar grains that are oriented parallel to the building direction are Fig. 22 e The EBSD phase maps for the samples built at a laser treatment at (a) 973, (b) 1073 K and (c) 1173 K for 3 h. observed (Fig. 12b). In the melt pool zone, microcracks aligned parallel to the building direction are occasionally found" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000200_s00202-019-00882-x-Figure18-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000200_s00202-019-00882-x-Figure18-1.png", "caption": "Fig. 18 Eddy current volumetric loss distribution, nin = 550\u00a0rpm/nout = 100\u00a0rpm", "texts": [], "surrounding_texts": [ "Distributions of the eddy current losses are presented in Figs.\u00a015, 16, 17, 18 and 19 for the NdFeB magnets [27, 28]. These losses are limited to electrically conductive parts only. Results correspond very well with these for electrical conductivities of permanent magnets, shown in Table\u00a02, and total losses and efficiency in Table\u00a03. Eddy currents are estimated at five rotational speeds, for inner rotor 150, 550, 1100, 1650 and 2200\u00a0rpm and for outer rotor rotational speeds 27, 100, 200, 300 and 400\u00a0rpm, respectively. Output rotor torque reduction due to eddy current demagnetization effect for different rotational speeds is shown in Fig.\u00a020. Eddy currents, neglecting the temperature changes of resistivity, are linearly related to rotational speed [29, 30]. Eddy current losses rise with the square of rotational speed. Also, they are influenced by the huge number of magnetic pole pairs. At rotational speeds greater than 500\u00a0rpm for the magnetic gear construction under consideration, the heat effect of eddy current losses could not be neglected. At speed 10,000\u00a0rpm, only the eddy current losses are overcoming 4% of the transmitted power. Eddy current volumetric loss distributions show that most of the losses are located in the low-speed rotor magnets. They are induced by high-speed rotor magnet movement and are amplified by modulating segments at high harmonics of the magnetic flux. Frequency separation of eddy current losses is important for loss analysis. 1 3 8 Magnetic field harmonic distortion of\u00a0CMG In many existing researches, eddy current losses are often ignored in steady state because of direct analogy with single-rotor electrical machines with permanent magnets. In MGs, these losses still appear because of relative movement of two rotors and modulating segments. Harmonic flux distortion increases eddy current effects. The main rotational frequencies of the magnetic field of the magnetic gear construction for the outer and inner rotors are 55\u00a0Hz and 10\u00a0Hz. The radial components of the flux density in the air gap between the inner rotor and the steel segments of the magnetic gear are shown in Fig.\u00a021. The radial components of the flux density in the air gap between the steel segments and the outer rotor of the magnetic gear are shown in Fig.\u00a022. The fast Fourier transform (FFT) analysis of the radial components of the flux density in the air gap between the inner rotor and the steel segments is shown in Fig.\u00a023. The FFT analysis of the radial components of the flux density in the air gap between the inner rotor and the steel segments is shown in Fig.\u00a024. 1 3 FFT analysis shows significant harmonic distortion. All frequency harmonics different from main 55/10\u00a0Hz are creating torques with no proper frequency of rotation, limiting the output CMG torque. According to these estimations, more than 20% of the total magnetic flux is engaged in these undesired frequencies [31]. This distortion effect does not influence the power efficiency, but it limits the CMG output torque and does not use the available flux from the permanent magnets. The harmonic distortion is not related to direct power losses because we are not using electrical supply for these fluxes, but it limits the output torque value of the CMG design. Some words on hysteresis loss are necessary, considering frequency-dependent hysteresis models, such as Steinmetz equation variations; they are suitable for loss superposition over magnetic flux amplitudes and frequency harmonics, like shown in Figs.\u00a023 and 24. However, a closer look reveals that hysteresis loss is not substantial for MG operation, as a power loss, covered by electrical excitation as it is in rotational machines. In MG, magnetic hysteresis loop causes time-dependent flux non-linearity, decreasing this way slightly the dynamic magnetic torque interaction between rotors. The summarized results for torque reduction and losses according to rotational velocity are presented in Table\u00a04. Losses are estimated for 150\u00a0rpm, 2200\u00a0rpm and 10,000\u00a0rpm of high-speed rotor. Results are showing significant dynamic torque reduction in high rotational speeds and rise of eddy current losses. According to estimated losses, efficiency mapping at MG overload, at torques above 320\u00a0Nm, is shown in Fig.\u00a022. At low-speed overload, efficiency is influenced by rotor slipping, while in high speeds it is influenced by eddy currents (Fig.\u00a025)." ] }, { "image_filename": "designv11_5_0002356_978-3-030-50460-1-Figure15.3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002356_978-3-030-50460-1-Figure15.3-1.png", "caption": "Fig. 15.3 Shapes and amplitude-periodic-length ratios. The simplified relation for the maximum-amplitude shape (d) ignores sheet thickness t", "texts": [ "2 indicates the reference coordinates in which the corrugated laminate is described and the curvilinear coordinates along the laminated sheet mid-plane. The relation between circular-segments radius R, corrugation amplitude c, and periodic length of the corrugation pattern P is given with (Kress and Winkler, 2010): R = 16c2 + P 2 32c . (15.1) This equation is rearranged to calculate the periodic length of one corrugationpattern unit cell for given reference-shape curvature R and the ratio c\u0304 that is the corrugation amplitude c normalized with respect to the periodic length P : P = R 32c\u0304 16c\u03042 + 1 ; c\u0304 = c P , (15.2) where Fig. 15.3 illustrates that the normalized amplitude c\u0304 is a shape characteristic. It appears in Figure 15.3, and Fig. 15.4(a) illustrates it more systematically, that the highest normalized shape curvature \u03ba\u0304 = P/R appears at intermediate corrugation amplitude, specifically at the semi-circular shape. Note that extreme morphing stretches, where the corrugated laminate tends to become flat, will create less bending curvatures for high-amplitude corrugations than for corrugation shapes composed of semi-circles (c\u0304 = 0.25). The condition, that the laminate with sheet thickness t must not penetrate itself, gives the upper bound of the corrugation amplitude c (Filipovic and Kress, 2018): 238 Kress, Filipovic 0 \u2264 c \u2264 P \u2212 t+ \u221a (P \u2212 t) 2 \u2212 P 2 4 2 ", "5) If the corrugated-shape unit cell with undeformed length P is stretched to flatness, its length will increase from P to Ls. This thought allows the estimate of maximum morphing capacity in terms of stretch \u03bb (Kress and Filipovic, 2019): \u03bblim = Ls P , (15.6) where Fig. 15.4(b) illustrates that the morphing stretch limit increases progressively 15 Manufacturing and Morphing Behavior of High-Amplitude Corrugated Laminates 239 with normalized corrugation amplitude c\u0304. For a maximum-corrugation amplitude shown in Fig. 15.3 (d) where c\u0304 \u2248 0.933, the stretch limit reaches a value of more than five. Note that the stretch limit in (15.6) is based on a pure geometric consideration and that material strength may dictate lower values, depending on material strain limits and the ratio of laminate thickness to reference-shape curvature. The proposed manufacturing method relies on temperature-induced curvature where the ratio of laminate thickness t to bending radius R remains rather small so that the classical theory of laminated plates (CTLP), see for instance in Jones (1975), is applicable" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001286_lifetech48969.2020.1570620257-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001286_lifetech48969.2020.1570620257-Figure1-1.png", "caption": "Fig. 1. IoT hardware.", "texts": [ " Because the machine is based IoT, it can use for monitoring the amount of feeding food of the pet for each meal in order to be precaution to reduce the risk of sickness. A. IoT Hardware 1) Arduino IDE: It is an open-source arduino software for developing IoT projects because they contain a large set of common input and output functions. The program is written in C and C++ language. It is used to write and upload program to Arduino compatible boards and other vendor development boards such as ESP32. This software is very widely used among programmers because of its open source. 2) ESP32: As shown in Fig. 1IoT hardwarefigure.1(a), it is a microcontroller IC that supports WiFi and Bluetooth 4.2 BLE made by Espressif. It is suitable for a use in IoT applications in terms of properties and price as shown in [1], [2]. They used ESP32 for a supervisory control and data acquisition and a smart surveillance system. 978-1-7281-7063-3/20/$31.00 \u00a92020 IEEE 132 20 20 IE EE 2 nd G lo ba l C on fe re nc e on L ife S ci en ce s a nd T ec hn ol og ie s ( Li fe Te ch ) 9 78 -1 -7 28 1- 70 63 -3 /2 0/ $3 1. 00 \u00a9 20 20 IE EE 1 0. 11 09 /L ife Te ch 48 96 9", " Date & Time Feeding food amount per meal (grams) Date Time Desired food amount Left food in tray Current feeding amount 1 16/11/2019 18:00 50.00 0.00 53.15 2 16/11/2019 19:00 50.00 32.86 22.14 3 17/11/2019 13:00 100.00 55.23 54.20 4 17/11/2019 15:00 100.00 24.36 84.69 5 18/11/2019 12:30 80.00 18.00 66.27 6 19/11/2019 16:00 60.00 0.00 65.98 7 19/11/2019 12:00 60.00 28.32 38.29 8 19/11/2019 18:00 60.00 14.43 55.78 9 20/11/2019 14:30 30.00 0.00 37.63 10 20/11/2019 15:00 30.00 0.00 35.32 B. IoT sensors, motors, and others MG995 High-speed standard servo as shown in Fig. 1IoT hardwarefigure.1(b) is used as a motor to release food from a container to a tray. It can rotate approximately 120 degrees (60 in each direction). It is suitable for this work because it can use any servo code, hardware or library to control these servos as addressed in [3], [4], and it can fit in small places. A load cell, called weight sensor, is used with HX711 ADC [5] to amplify the signal received from the load cell. This amplifier enables the Arduino to detect changes in resistance from the load cell. They as shown in Fig. 1IoT hardwarefigure.1 (c)-(d) are used as equipment for a scale to determine the weight of things. Fig. 3. Work flow diagram 133 Protoboard breadboard 400 holes as shown in Fig. 1IoT hardwarefigure.1 (e) is a solderless breadboard with dual bus strips on both sides. Because it does not require soldering, it is reusable and popular for with students and in technological education. This makes it easy to use for creating temporary prototypes and experimenting with circuit design. C. IoT Software Blynk is a platform for connecting to the development board. It could connect with the Arduino, ESP32, and other microcontroller boards vial Ethernet, Wi-Fi, Bluetooth Cellular or Serial. Blynk provides a server for collecting the IoTfeeding data, an application builder for creating an interface, and libraries for developing the user program" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000451_j.robot.2016.05.012-Figure10-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000451_j.robot.2016.05.012-Figure10-1.png", "caption": "Fig. 10. The foot sensor and foot sensor housing.", "texts": [ " The simulation comparing the loaded and non-loaded scenarios validates the presented criterion. As expected, the MFFSM is sensitive to height, foot placement, and top-heaviness and both the FFSM and MFFSM indicate instability at the same time. To physically validate the FFSM and MFFSM, the Lynxmotion hexapod robot, shown in Fig. 6, was equippedwith Force-Sensitive Resistors (FSR-402) similar to [38]. The sensors were calibrated after they were embedded into rapid prototyped sensor housings that attach to the robot legs as shown in Fig. 10. The calibration results, shown in Fig. 11, were used to fit a curve for each sensor allowing a direct correlation between the output voltage and the applied force. The hexapod robot properties and the properties of the cylindrical rod are given in Table 1. Fig. 12 provides a side profile sketch of the hexapod robot indicating the lengths provided in Table 1. The following experiments are done with the hexapod robot under the support/stance phase of a tripod gait with no changes in the support configuration" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002210_j.robot.2020.103670-Figure20-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002210_j.robot.2020.103670-Figure20-1.png", "caption": "Fig. 20. Bottom of the tread of the robot\u2019s front wheels.", "texts": [ " \u2206x is the orizontal movement of coordination as the robot\u2019s angle inclines the distance between 0\u03a3B and \u03a3B, Fig. 19). The wheelchair emains stopped until the system finishes lifting the robot\u2019s front heels. p4 does not change. Thus, x = \u2212(X4 \u2212 0X4) (4) ubstituting (2) and (3) for (4), x = l4C (cos 0\u03c6123 \u2212 cos\u03c6123) + l2(cos 0\u03c612 \u2212 cos\u03c612) +lLB(1 \u2212 cos\u03c61) + hLB sin\u03c61 (5) The position vector of \u03a3B in 0\u03a3B is described below (Fig. 19). p\u2206x = [\u2206x 0]T (6) 4.2. Robot front wheel height needed to climb a step The angle of the robot necessary to place the front wheels on a step is clarified in Stage 3. In Fig. 20, q1 is the position of the robot\u2019s middle wheel axes (Here, q1 = p1, Fig. 18), q2 is the position of the robot\u2019s front wheel axes, and q3 is the tread of the robot\u2019s front wheels. \u03a3I and \u03a3II are the local coordinate system. (Here, \u03a3I = \u03a31, Fig. 18.) \u03c11 and \u03c12 are the angles of \u03a3I and \u03a3II formed by \u03a3B and \u03a3I , respectively. From Figs. 18 and 20, \u03c11 = \u03c61 (7) The position vectors for the joints are expressed as Bqj = [xj yj]T (j = 1 \u2212 3). When \u03a3 and \u03a3 parallel \u03a3 , Bq = [0 R ] T , I II B 1 B 1q2 = [WBf \u2212 RB + rBf ]T and 2q3 = [0 rBf ]T (Figs", " 12 = \u03c1123 = \u03c11 + \u03c12 because \u03c13 = 0. Thus, the homogeneous ransformation matrices Bt2 and Bt3 are as given below. t2 : [cos \u03c112 \u2212 sin \u03c112 x2 sin \u03c112 cos \u03c112 y2 0 0 1 ] (8) Bt3 : [cos \u03c112 \u2212 sin \u03c112 x3 sin \u03c112 cos \u03c112 y3 0 0 1 ] (9) Here, Bq2 : [ x2 y2 ] = [ cos \u03c11 \u2212 sin \u03c11 sin \u03c11 cos \u03c11 ][ WBf \u2212(RB \u2212 rBf ) ] + [ 0 RB ] (10) Bq3 : [ x3 y3 ] = [ cos \u03c11 \u2212 sin \u03c11 sin \u03c11 cos \u03c11 ][ WBf \u2212(RB \u2212 rBf ) ] +rBf [ cos \u03c112 sin \u03c112 ] + [ 0 RB ] (11) When the tread of the robot\u2019s front wheels, q3, is at the bottom of the wheel, \u03c112 = \u221290 deg (Fig. 20). In this case, from (7) and (11), y3 = WBf sin\u03c61 \u2212 (RB \u2212 rB) cos\u03c61 + RB \u2212 rBf (12) ere, in\u03c61 = \u221a 1 \u2212 cos2 \u03c61 (13) From (7), (12), and (13), cos\u03c61 = e1(e1 \u2212 y3) + WBf \u221a 2 \u00b7 e1 \u00b7 y3 + WBf 2 \u2212 y32 e12 + WBf 2 (14) Here, e = R \u2212 r . 1 B Bf s d 4 b r 0 i t s B 0 F v \u2225 r f 4 i h r v B 0 w c \u2225 \u221a h t i 5 e \u00b5 a u 5 A r t v w b a o s b When the robot climbs a step the height of which is h, substitute y3 = h for (14), cos\u03c61 = e1(e1 \u2212 h) + WBf \u221a 2 \u00b7 e1 \u00b7 h + WBf 2 \u2212 h2 e12 + WBf 2 (15) Using (15), the system can tilt the robot and control its angle ufficiently to put the wheels on the step when the vehicles etect the step height" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000654_smc.2016.7844562-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000654_smc.2016.7844562-Figure2-1.png", "caption": "Fig. 2. Body and ground fixed frames", "texts": [ " The spacing between two adjacent arms is 45 degrees. Rotors depicted with blue color (2, 4, 6 and 8) rotate clockwise, while the ones depicted with red color (1, 3, 5 and 7) rotate counterclockwise. Therefore, the octorotor is inherently balanced with regards to the drag moment generated by the rotors. All other system components (on a real system) should be placed in a protective case which is mounted to the support plate. In order to derive the mathematical model of the octorotor we define two frames of reference as shown in Fig. 2. The first frame of reference {o} is a body fixed frame attached to the center of mass of the octorotor, while the second frame {g} is a ground fixed frame using the ENU convention meaning that the XB , YB and ZB axes are pointing north, east and up, respectively. The XY Z and XBYBZB coordinate systems are both right handed Cartesian coordinate systems in three dimensions and we refer to them as the local and the global coordinate system, respectively. Rotations around the axes of the local coordinate system in terms of minimal representation can be described by the Euler angles \u03c6, \u03b8 and \u03c8 which are also known as roll, pitch and yaw, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002541_j.surfcoat.2021.127028-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002541_j.surfcoat.2021.127028-Figure1-1.png", "caption": "Fig. 1. shows the experimental devices (a) JECM device (b) JED device.", "texts": [ " Since nickel coating and nickel-based composite coatings are important decorative and protective coatings in the industry [21,22], the authors tried to deposit a nickel coating on the post-JECM using JED. In this article, we will study the influence of different JECM current densities on the surface morphology, bonding force, hardness, and corrosion resistance of the post-JED substrate. In the experiment, the jet range, scanning speed, processing time, flow rate, and bath temperature demonstrate that they can be controlled by the self-designed JECM system and JED system. This ensures the accuracy and stability of the electrolysis and deposition process. The experimental equipment is shown in Fig. 1. JECM is a process of material removal by anodic dissolution. In the process of JECM, the SLM substrate is used as the workpiece to be connected to the positive electrode of the DC power supply and fixed on the worktable. The jet device is connected to the negative pole and reciprocates with a feed rate of 100 mm/min. Using a 10 wt% NaCl solution at 25 \u25e6C, the filtered NaCl electrolyte is pumped into the electrolyte spray device through a corrosion-resistant pump, and the electrolyte flow rate is 350 L/h (vJECM = 4.86 m/s). The nozzle structure of JECM is similar to that of JED, as shown in Fig. 1b. The nozzle hole area of the JECM is 2 \u00d7 10\u2212 6 m2, and the nozzle hole area of the JED is 4 \u00d7 10\u2212 6 m2. After being fully \u201cnegative\u201d in the cavity of the spray device, an electrolyte jet with a stable broken length is formed through the nozzle, which shoots to the surface of the SLMed substrate, and electrochemical anodic dissolution occurs at the spray point. Through the reciprocating jet electrolysis nozzle, the surface of the SLMed substrate is continuously dissolved to remove surface defects and improve surface roughness. Meanwhile, the micropore structure formed after JECM can improve the bonding strength between the SLM substrate and coating, as shown in Fig. 1a. During electrochemical machining, there are many factors that affect the surface quality, such as machining time, electrolyte temperature, electrolyte type, current density, etc. [23,24]. The processing parameters of JECM are set as follows: the initial machining gap is 1 mm and the processing time is 10 min. In the JECM process, the current densities are variable, at 30 A/cm2, 60 A/cm2, 90 A/cm2, and 120 A/cm2, respectively. It is necessary to use an average to describe the current density. The current density of JECM was marked as iJECM", " Surface & Coatings Technology 412 (2021) 127028 During JED, the SLM substrate is used as the workpiece to connect to the negative electrode of the DC power supply, and the spray device is connected to the negative electrode. Watts solution is pumped into the spraying device by the water pump, then flowed through the nickel electrode in the spraying device, and sprayed on the surface of the SLM substrate through the nozzle. At the spray point, nickel ions are reduced and deposited to form a nickel-plating layer, as shown in Fig. 1b. The compositions of the JED solution are shown in Table 1. Its compositions were 260 g/L NiSO4\u22c56H2O, 40 g/L NiCl2\u22c56H2O, and 40 g/L H3BO4. The JED process is shown in Fig. 1b. Here, the cathode workpiece was fixed, the anode nozzle was reciprocated, and the feed rate was 200 mm/min. The electrodeposition solution temperature was controlled at 50 \u00b1 1 \u25e6C, the pH was controlled within a range of 4.0 \u00b1 0.2, and the flow rate was200 L/h. The current density of the prepared nickel coating was 100 A/dm2, and the current densities mentioned in this paper are \u2018average current density,\u2019 which is obtained by dividing the output current by the outlet area of the nozzle. The current density of JECM was marked as iJED" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002578_j.ymssp.2021.107786-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002578_j.ymssp.2021.107786-Figure3-1.png", "caption": "Fig. 3. (a) The UofA planetary gear experimental test rig [30] and the accelerometer locations in planetary gearboxes (b) and fixed-axis gearboxes (c).", "texts": [ " In this paper, to validate the explanation discussed above, the coarse IAS estimation was obtained using the second harmonic as well as the third PGM harmonic for comparison. The magnitude of the third harmonic of the PGM (around 75 Hz), shown in Fig. 2, is weaker than the second PGM harmonic, while still being separable from the background. However, due to the requirement that the speed variation be less than 20%, existing phase demodulation based techniques tested on this data all used the second PGM for method validation [27,32]. The UofA test-rig, shown in Fig. 3, consists of a transmission with five gear stages, with the drive motor connected to a bevel gearbox (speed reducer), followed by the first planetary gear stage (sequential meshing arrangement, speed reducer), the second planetary stage (sequential meshing arrangement, speed reducer), and two stages of speed-up fixed-axis gearboxes. Two variable frequency drives (VFD) are used to control the speed of the drive motor and the load of the driven motor. The two stages of planetary gearboxes were lubricated by a recirculating oil system while the bevel gearbox and two speedup fixed-axis gearboxes were lubricated by splash lubrication" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002175_icra40945.2020.9197401-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002175_icra40945.2020.9197401-Figure6-1.png", "caption": "Figure 6. r1 and r2 represent FRAs. r3, r4, and r5 represent the pressure chambers from the base to the end effector. g: gravity force.", "texts": [ " 5d), the SCS (Fig. 5g), and the encapsulation layer (Fig. 5i) were assembled (Fig. 5j). As the final step, the prepared DragonSkin 30 was poured into a capping mold, then both sides of the assembled manipulator were dipped into the mold one by one. The pressure pipes for the FRA actuation were inserted and fixed by silicone glue (Fig. 5k-l). The manipulator was characterized in terms of compliance, stiffness capability and workspace. All tests were performed in the x-y plane, with reference to Fig. 6, by compensating the gravity. For clarification, the pressurized regions are named as ri (i = 1, 2, 3, 4, 5). While the r1 and r2 represent the FRAs, the r3, r4 and r5 represent the joints of the SCS from proximal to distal part. To actuate the manipulator, only one (r1) of the two FRAs is pressurized at a time. On the other hand, the joints of the SCS are pressurized together with the same pressure to demonstrate the maximum stiffness contribution of the SCS to the manipulator. For the compliance, SCS stiffness and controllable stiffness tests, force data were collected" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000327_j.mechatronics.2015.06.014-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000327_j.mechatronics.2015.06.014-Figure3-1.png", "caption": "Fig. 3. Force analysis of link rod for driving limb.", "texts": [ " Then the force screw applied at the end of actuated limb i produced by force f i can be achieved by $f i \u00bc f i; rBi f i\u00bd T \u00bc S\u00fei F is0; r\u0302Bi S\u00fei F is0 T S\u00fei G T JAi; r\u0302Bi S\u00fei G T JAi T si; \u00f09\u00de where r\u0302 is defined as cross-product factor that makes r f \u00bc r\u0302 f , and rBi represents the position vector of joint Bi. According to literature [17], the force screw of $fi could be also called actuated force screw. Combining all actuated limbs and using matrix expression yields X6 i\u00bc1 $fi \u00bc Fs0 G f s h i s1 s2 s6\u00bd T ; \u00f010\u00de where Fs0 \u00bc P6 i\u00bc1 S\u00fei F is0; r\u0302Bi S\u00fei F is0 T is determined by motion states of actuated limbs and G f s h i \u00bc G f s1 G f s2 G f s6 h i , of which G f si \u00bc S\u00fei G T JAi; r\u0302Bi S\u00fei G T JAi T . Force analysis of link rod of actuated limb i is shown in Fig. 3. For this link rod, the micrometric displacement of the end point Bi is produced synergistically by the tension/compression deformation along its axial direction and bending deformation along its radial direction. According to paper [22], the bending deformation which is caused by the gravity and inertia moment is relatively smaller than the tension/compression deformation. Moreover the maximum bending deformation will exist near the middle of the link rod and, therefore, influence on the micrometric displacement of the point Bi is very little" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001397_1350650120942329-Figure13-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001397_1350650120942329-Figure13-1.png", "caption": "Figure 13. Velocity streamline with the azimuth angle of: (a) 260 ; (b) 270 ; (c) 273 ; (d) 280 ; (e) 290 .", "texts": [ " Furthermore, a relatively better choice for jet elevation angle value that can generate better lubrication. Similarly, the jet nozzle spatial position is arranged at the point (210, 100, 40). The jet elevation angle is set to a recommended value of 97.5 previously, while the jet azimuth angle increases from 260 to 290 (see Table 2). Figure 12 illustrates the contour plots of the oil volume fractions. Obviously, Figure 12(c) exhibits superior oil volume fraction on the pinion surface with the jet azimuth angle of 273 . Figure 13 illustrates the velocity streamlines of the oil. Obviously, the streamline distribution in Figure 13(c) is better than others so it lubricates best consistent with the above oil volume fraction results. Figure 13 illustrates the behavior of the oil volume fractions and oil pressure. It can also be seen that both the oil volume fractions and oil pressure reach maximum values under the jet azimuth angle of 273 . By applying the mathematical model with different jet azimuth angles, the corresponding impingement depths can be calculated as shown in Figure 14(c). A maximum impingement depth is achieved with jet azimuth angle of 273 ; besides, the azimuth angle should within the range of 257 to 284 . Exceeding that range there will be no impingement depth on the gear surface, resulting in rather poor lubrication" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000619_asjc.1433-Figure10-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000619_asjc.1433-Figure10-1.png", "caption": "Fig. 10. Single link robot with flexible joint.", "texts": [ " In term of Settling time (ts), the proposed controller has 0.90s compared to the 1.93 s of the conventional FSMC. We can also note that the simulation results demonstrate the superiority of proposed controller regarding Table II and Figs 8 and 9. 4.2 Example 2 In this example, we have taken two experimental cases (with and without model uncertainty) in order to analyze the performance of the proposed controller. Then they are compared with other methods reported in the literature. The single link robot with flexible joint [27], shown in Fig. 10, is investigated as a case study for the proposed approach. The state space equations for the system are _x1 t\u00f0 \u00de \u00bc x3 t\u00f0 \u00de _x2 t\u00f0 \u00de \u00bc x4 t\u00f0 \u00de _x3 t\u00f0 \u00de \u00bc 1=I k x2 t\u00f0 \u00de x1 t\u00f0 \u00de\u00f0 \u00de mgL sin x1 t\u00f0 \u00de\u00f0 \u00de\u00f0 \u00de _x4 t\u00f0 \u00de \u00bc 1 J u t\u00f0 \u00de k x2 t\u00f0 \u00de x1 t\u00f0 \u00de\u00f0 \u00de\u00f0 \u00de= 8><>: where k = 100 Nm/rad, g = 9.8 m/s2, and the rest of parameters are assumed unity. In order for the sector nonlinearity condition satisfied, the nonlinear expression z = sin (x1(t)) for x1(t)\u2208 [ \u03c0,\u03c0] is expressed as [28]: \u00a9 2016 Chinese Automatic Control Society and John Wiley & Sons Australia, Ltd z \u00bc sin x1 t\u00f0 \u00de\u00f0 \u00de \u00bc \u2211 2 i\u00bc1 Mi z\u00f0 \u00debi x1 t\u00f0 \u00de where, b1 = 1,b2 = 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001839_012015-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001839_012015-Figure6-1.png", "caption": "Figure 6. Sets of 3D printed tribology specimens", "texts": [ " These parameters are expected to influence the results of the subsequent tests (Tensile, tribology, bending, impact, and surface roughness). Table 2 summarises all the examined parameters of the current study. CEURSIS IOP Conf. Series: Materials Science and Engineering 749 (2020) 012015 IOP Publishing doi:10.1088/1757-899X/749/1/012015 The 3D printed specimens for tribology test have a diameter of 8 mm and a length of 15.2 mm. Some sets of tribology test pieces printed in various orientations using FDM and DLP are exhibited in figure 6. The printed specimens were examined by three tests which are tribology, surface structure (optical microscope) and surface roughness. The tribological measurement was performed at tribology laboratory in Szent Istv\u00e1n University under the laboratory conditions. The first step in the measurement process is to connect the measuring circuit, which consists of a computer, Spider 8 measuring converter, a tribotester, and an inverter. Spider 8 is a strain gauge measurement device for measuring load, force, and wear" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000867_0305215x.2019.1634702-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000867_0305215x.2019.1634702-Figure1-1.png", "caption": "Figure 1. Geometry of a single-stage spur gear (D\u00f6rterler, \u015eahin, and G\u00f6k\u00e7e 2019).", "texts": [ " In that study, the problem defined by Yokota, Taguchi, and Gen (1998) was solved first and then a modified version of the problem was created and solved. The results were compared with the GA results in Yokota, Taguchi, and Gen (1998) and it was shown that more successful results were obtained. D\u00f6rterler, \u015eahin, and G\u00f6k\u00e7e (2019) applied the GWO to both cases of the problem in the studies by Savsani, Rao, and Vakharia (2010) and Yokota, Taguchi, and Gen (1998), and obtained better scores on both cases with GWO than the scores of the previous studies. In this study, the problem of the design of the spur gear shown in Figure 1 with respect to the minimum weight is handled in two different situations. The problem formulated in Yokota, Taguchi, and Gen (1998) with five design variables and five constraints is discussed as Case 1; and the version presented in Savsani, Rao, and Vakharia (2010) as per the standard American Gear Manufacturers Association (AGMA) equations, to include six design variables and eight constraints, is discussed as Case 2. Two different cases of the problem were applied to the design variable ranges proposed in the previous studies (Savsani, Rao, and Vakharia 2010; Yokota, Taguchi, and Gen 1998) and thus the problem and its cases were conducted on four different scenarios" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002420_s12540-020-00900-9-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002420_s12540-020-00900-9-Figure4-1.png", "caption": "Fig. 4 Geometric and heat source model. a Geometric model; b Gaussian heat source model", "texts": [ " A three-dimensional steady-state model was established, while the heat transfer and fluid flow in the molten pool were calculated by solving the equations for the conservation of energy, momentum and mass. Since the coating is symmetrical about the centerline only half of the workpiece is considered in the calculations. The molten metal is assumed to be an incompressible, laminar, and Newtonian fluid. The substrate with dimensions of 50 \u00d7 10 \u00d7 8\u00a0mm was established, while the morphology of the coating was consistent with the experimental results. The finite element mesh was exhibited in Fig.\u00a04a. Since the temperature of the materials in the laser cladding process change rapidly, it is necessary to define the thermal physical properties of the substrate and the FeCoCrNi HEAs powder at different temperatures to obtain accurate simulation results. In this paper, the specific heat, thermal conductivity, viscosity and surface tension is varied with temperature, while density is defined as constant. Furthermore, the effect of the latent heat of fusion is considered. The initial condition refers to ambient temperature of the workpiece, which can be set to 20\u00a0\u00b0C since the actual laser cladding process is conducted at room temperature", " The convective heat transfer coefficient is set to be 40 [J/(m2\u00a0K)] since air is the cooling medium in this research [33], while the boundary convection condition is applied to all external surfaces of the finite element model. In this paper, due to the low laser power and fast scanning speed, the keyhole effect will not appear in the laser experiment, and the ratio of depth to width is small. Therefore, the Gaussian surface heat source model is used to characterize the thermal effect in laser cladding, as displayed in Fig.\u00a04b. Figure\u00a05 illustrated the cross-sectional profiles of the samples, which can be divided into three zones: substrate, heataffected zone (HAZ), and coating. The arched surface of the coating is located above the substrate and the fused line can be observed between HAZ and the coating, demonstrating a good metallurgical bonding is formed. Dilution rate \u03b7 can be calculated as follows [34]: where S1 is the fusion area of the substrate and S2 is the area of the coating above the substrate surface. The cross-sectional morphology of the coating can be considered as a combination of two ideal arcs with different radii" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002525_taes.2021.3061795-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002525_taes.2021.3061795-Figure1-1.png", "caption": "Fig. 1. Frame definitions: red - fixed north-east-down frame, blue: body frame fixed on the body of the UAV, green: frame fixed on the IMU frame used for its measurements", "texts": [ " In this section, we present the mathematical model of the six degree of freedom kinematics considered in this paper, and the measurements from IMU and GPS. Let the b-frame be the body-fixed frame of the quadrotor UAV with the origin located at its center of gravity. The local north-east-down (NED) frame is defined as the f -frame, with the origin at the base GPS sensor. We distinguish the frame fixed to the IMU, in which the IMU measurement readings are resolved, from the b-frame, and it is denoted by the i-frame. These frames are illustrated in Figure 1. The orientation of the IMU with respect to the body-fixed frame is described by a fixed rotation matrix Rbi \u2208 SO(3), where the threedimensional special orthogonal group is given by SO(3) = {R \u2208 R3\u00d73 |RTR = I3\u00d73, det[R] = +1}. Let x \u2208 R3 be the position of the origin of the b-frame, or the center of gravity of the UAV. The corresponding velocity and acceleration are denoted by v \u2208 R3 and a \u2208 R3, respectively. All of x, v, and a are resolved in the f -frame. The rotation matrix from the b-frame to the f -frame is denoted by R , Rfb \u2208 SO(3)" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001844_s12541-020-00333-9-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001844_s12541-020-00333-9-Figure7-1.png", "caption": "Fig. 7 Finite element model of flexspline and wave generator", "texts": [ " Three-dimensional elements are used for finite-element modelling. The structure is divided into two parts while generating mesh, the gear and the cup body. Mesh generation of the gear part and wave generator are accomplished by sweeping with the mesh type as hexahedron element (Solid 185). The mesh density is gradually refined in the tooth area, which represents the potential contact region. Tetrahedral element (Solid 187) is used in the cup body part. The finite model consists of 56,266 elements and is shown in Fig.\u00a07. The material settings used in the finite element model are shown as Table\u00a04. The the bottom of the flexspline and inner hole of the wave generator are fully restrained. The contact type of flexspline, wave generator and circular spline are represented by CONTAC185 in ANSYS. The applied torque between the circular spline and flexspline is 25 Nm. The results indicate that the stress is concentrated at both ends of the teeth, especially at the back end of the teeth (Fig.\u00a08). It can be seen, that, when aided by the wave generator, the stress distribution at the root of the gear tooth is unevenly distributed in the direction of tooth thickness", " Taking \u03c3a as the average stress for a single tooth surface and \u03c3s as the standard deviation for the stress of the sampling points on the tooth surface, the bias load index P is defined as: where i is the order of different teeth in the meshing area, and j is the stress extraction point for the tooth root. In order to study the partial axial load index of different flexspline materials, the material 35CrMnSiA is taken as the control group. The detailed steps are as follows: Step 1: The control group (Material 35CrMnSiA) is firstly studied. In the simulation model (Fig.\u00a07), we take materials of the flexspline, circular spline and wave generator as 35CrMnSiA, 42CrMo and 45#, respectively. Their material properties are listed in Tab.4, including tensile modulus, poisson ratio, shear modulus, and density. Step 2: The simulation analysis is performed based on definitions of the torque and boundary conditions. Then, the stress for the root of each meshing tooth in the meshing area is extracted. (22) P = 1 m m\ufffd i=1 si ai = 1 m m\ufffd i=1 \ufffd 1 n n\u2211 j=1 \ufffd i,j \u2212 ai \ufffd2 ai Step 3: The partial axial load index of the material is obtained through substituting the extracted stress value into Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000129_tia.2019.2946525-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000129_tia.2019.2946525-Figure2-1.png", "caption": "Fig. 2. Rewound six-phase 24S/22P FSCW-SPM machine.", "texts": [ " This paper is arranged as follows: Section II investigates fault responses, highlighting the influence of the third harmonic back EMF and variable loads; Section III shows the development of online fault detection based on high-frequency injection for high-frequency input impedance extraction; the postfault control strategy is developed in Section IV demonstrating that the designed machine can continuously run following a STSC fault. Finally, conclusions are drawn in Section V. 0093-9994 (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. II. FAULT RESPONSE CONSIDERING THIRD HARMONIC BACK EMF The focus of this study is a six-phase 24-slot 22-pole alternate-teeth-wound FSCW-SPM machine as shown in Fig. 2. Phase F winding has been rewound for ITSC fault inception. Based on simplified magnetic circuit analyses, a closed-form method has been carried out for calculating the parameters of the T-type equivalent circuit as shown in Fig. 3. In the simplified model, the winding turns at slot level is modeled as horizontally distributed layers while the number of strands for each turn is assumed to be one. Detailed expressions are given in Appendix. Since a short circuit fault generally reduces saturation level, the saturation effect has been neglected for this case study" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001800_978-3-030-40410-9-Figure3.7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001800_978-3-030-40410-9-Figure3.7-1.png", "caption": "Fig. 3.7 Roll, pitch, and yaw angles", "texts": [ "23) In general, R1 2 is called RZY Z and denoted as ZYZ Euler angle transformation RZY Z = \u23a1 \u23a3C\u03c6C\u03b8C\u03c8 \u2212 S\u03c6S\u03c8 \u2212C\u03c6C\u03b8 S\u03c8 \u2212 S\u03c6C\u03c8 C\u03c6S\u03b8 S\u03c6C\u03b8C\u03c8 + C\u03c6S\u03c8 \u2212S\u03c6C\u03b8 S\u03c8 + C\u03c6C\u03c8 S\u03c6S\u03b8 \u2212S\u03b8C\u03c8 S\u03b8 S\u03c8 C\u03b8 \u23a4 \u23a6. (3.24) where C and S denote cos( ) and sin( ) functions, respectively. Other successive rotations about the principal axes of frame o1x1y1z1 called roll, pitch, and yaw angles\u2014denoted as (\u03c6, \u03b8, \u03c8), respectively\u2014can be used to specify the rotation matrix. We will describe the three successive rotations as follows. The first rotation is roll rotation about the z1-axis through \u03c6 angle, then pitch rotation about the y1-axis by an angle \u03b8 , and yaw about the x1-axis by an angle \u03c8 , as revealed in Fig. 3.7. Since the rotations are made relative to frame o1x1y1z1, as illustrated in Fig. 3.8, the transformation matrix will be RXY Z = \u23a1 \u23a3C\u03c6 \u2212S\u03c6 0 S\u03c6 C\u03c6 0 0 0 1 \u23a4 \u23a6 \u23a1 \u23a3 C\u03b8 0 S\u03b8 0 1 0 \u2212S\u03b8 0 C\u03b8 \u23a4 \u23a6 \u23a1 \u23a3 1 0 0 0 C\u03c8 \u2212S\u03c8 0 S\u03c8 C\u03c8 \u23a4 \u23a6 = \u23a1 \u23a3C\u03c6C\u03b8 \u2212S\u03c6C\u03c8 + C\u03c6S\u03b8 S\u03c8 S\u03c6S\u03c8 + C\u03c6S\u03b8C\u03c8 S\u03c6C\u03b8 C\u03c6C\u03c8 + S\u03c6S\u03b8 S\u03c8 \u2212C\u03c6S\u03c8 + S\u03c6S\u03b8C\u03c8 S\u03b8 C\u03b8 S\u03c8 C\u03b8C\u03c8 \u23a4 \u23a6. (3.25) Example 3.2 In general, we can write a program using MATLAB to calculate the Euler transformation matrix from the basic three angles. We have considered that the rotation will be in degrees, so we have used the function cosd( ) and sind( )" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002529_1350650121998519-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002529_1350650121998519-Figure4-1.png", "caption": "Figure 4. Schematic diagram of the test rig: (a) test rig of grooved water-lubricated bearing, (b) arrangements of pressure sensors and external load, and (c) schematic diagram of the test bearing.", "texts": [ " As the rotary speed increases to 4000 r/min, the numerical predictions agree well with the experimental results with a difference of only about 4.2%. Considering the modeling errors and manufacturing errors, the present numerical results and methods can be accepted. Table 1. Results of mesh sensitive analyses. Cases Mesh Load capacity (N) Difference compared to case 1% Circumferential Radial Axial Eccentricity ratio Groove width Single groove A film land between two grooves Overall smooth Groove Film land 20 0.1 0.7 10 15 387 10 10 200 181.867 0.89 As illustrated in Figure 4(a), the shaft is driven by a highspeed motorized spindle at one end. The shaft and the spindle are connected by a coupling. Two identical cylindrical roller bearings are used to support the shaft. A test grooved water-lubricated journal bearing is arranged at the other end of the rotor. The shaft is made of stainless steel and the test bearing is made of brass. Water is supplied into the test bearing through a pipe with a water pump, and flows out of the bearing along its axial sides and directly flows into the water tank. A rope is tied around the outer surface of the bearing and a weight hangs below the test bearing. Four pressure sensors are installed into the holes of the bearing along the circumferential direction with an interval of 30\u00b0 from 150\u00b0 to test the film pressure. The installation positions of the pressure sensors and the applied weight are shown in Figure 4(b). Figure 4(c) shows the schematic diagram of the test bearing. As shown in the figure, there are 12 grooves arranged along the circumferential direction with an interval of 30\u00b0. The width and depth of a groove are 200 and 500 \u03bcm, respectively. The diameter of the test bearing is 50 mm and the radial clearance is 50 \u03bcm. The rotary speeds are 2500 and 3000 r/min, and the applied force is 100 N. Figure 5 shows the numerical and experimental results of the grooved bearing. As shown in the figure, generally, the calculated varied trend for the pressure at different angle matches well with experimental results" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001584_s40815-020-00940-8-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001584_s40815-020-00940-8-Figure2-1.png", "caption": "Fig. 2 Ball and beam system, along with the position of the ball relative to the center of the beam at several sample points", "texts": [ " The output of the IT2-FLS is defined as [44]: yo \u00bc y\u00fe y =2; \u00f04\u00de where y and y are defined as follows [45]: y \u00bc PL i\u00bc1 f iYi \u00fe PN i\u00bcL\u00fe1 f iYi PL i\u00bc1 f i \u00fe PN i\u00bcL\u00fe1 f i ; \u00f05\u00de y \u00bc PR i\u00bc1 f iYi \u00fe PN i\u00bcR\u00fe1 f iYi PR i\u00bc1 f i \u00fe PN i\u00bcR\u00fe1 f i : \u00f06\u00de Here, L and R are the switching points which are extracted from Karnik and Mendel (KM) algorithms [42], f i and f i denote upper and lower bounds of firing strengths, respectively; and are defined as: f i \u00bc li 1 l ~Ai k ; \u00f07\u00de f i \u00bc l ~Ai 1 l ~Ai k : \u00f08\u00de Here, is the algebraic product \u00f0 \u00de operator. Consider the ball and beam system in Fig. 2a. In this figure, the distance between the ball and the beam center is x, and the beam makes an angle of / with the horizontal axis. A simple model of this system can be expressed as [46]: \u20acx \u00bc 9:81 sin kv\u00f0 \u00de: \u00f09\u00de In Fig. 2a, if the ball sits in the middle of the beam (x \u00bc 0), the beam angle generated by the motor will be / \u00bc 0. Motor voltage (v) constitutes the system input, and the angle the beam makes with the horizon is assumed to be proportional to motor voltage, i.e., / \u00bc kv; k is a positive constant. The control objective is to tune the motor voltage so that the reference ball tracks r \u00bc 0. In this case, the tracking error will be e \u00bc r x \u00bc x. The rules of FLS can be obtained in different ways. For example, with the opinions of an expert or based on data [47]. Figure 2a, b, and c shows examples of the bullet position on the bar. Depending on the position of the ball relative to the center of the bar and its velocity, fuzzy rules can be derived. Other shapes can be drawn for the position of the ball in different places of the bar, and a complete shape is presented in [46]. In FCMs, opinions of experts about the system determine concepts. For ball and beam, looking at Eq. 9 and defining the error shows that the error and the velocity error have a principal role in the controller", "; 9 Ri2 2 \u00bc IF _e is ~F i2 1 and v is ~F i2 2 Then ~w2;3 \u00bc Yi2 2 ; i2 \u00bc 1; . . .; 9 \u00f010\u00de ~F i1 j 2 N; Z;Pf g; j \u00bc 1; 2 and ~F i2 j 2 N; Z;Pf g; j \u00bc 1; 2 denote the membership functions related to the antecedent section of IT2-FCM rules. Also, l~F k j ; k \u00bc i1; i2; j \u00bc 1; 2 and l~F k j ; k \u00bc i1; i2; j \u00bc 1; 2 are the upper and the lower bounds of membership functions, respectively [42]. We obtain the IT2-FCM rules like the fuzzy rules. The fuzzy rules are presented in Table 1. FCM rules are obtained from Fig. 2. Using Table 1, the fuzzy rules of IT2-FCM are extracted. Example 1 If e is N, v can be either N or Z. So, if e is N and v is N, then w1;3 will be Z. If e is N and v is Z, then w1;3 can be Z or P (one expert may choose Z and another expert P). The fuzzy rules of IT2-FCM are eventually obtained in this way. After completing and modifying the rules, it is time to extract the output. Since the concepts can take on negative values here, we consider function F in Eq. 1 as a tangent hyperbolic function" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000366_j.mechmachtheory.2016.02.007-Figure9-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000366_j.mechmachtheory.2016.02.007-Figure9-1.png", "caption": "Fig. 9. Experimental setup.", "texts": [ " Experiments are performed on a rotor bearing system equipped with rolling element bearings and a rigid disk mounted at the center of the flexible shaft. The physical details of the rotor bearing system are as shown in Table 2. The rotor is supported on two rolling element bearings (Model SKF-RLS 5) with 45 microns clearance. The proposed polynomial type of nonlinear damping methodology as discussed in Section 3 is tested using experiments. The experimental setup of the rotor bearing system considered is shown in Fig. 9. The inherent nonlinearities of damping in the rotor bearing system are estimated using wavelets. Impulse response of the rotor is obtained using an impact hammer and the response is as shown in Fig. 10 (a). Impulse response is obtained from the shaft specimen, near the bearing since bearings are a major source of damping. The Fast Fourier Transform of the impulse response reveals the first natural frequency of the rotor at 48.6 Hz as shown in Fig. 10 (b), while the scalogram of the impulse response is presented in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001949_tec.2020.2998793-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001949_tec.2020.2998793-Figure1-1.png", "caption": "Fig. 1. The structure of a 12-8 pole DSEG.", "texts": [ " The optimization in dynamic response and adaptability to various operation points are verified by FEA. In Section 4, a 9 kW 12-8 pole DSEG prototype is tested with the proposed method and the VC method. The influences of the capacitor degradation are analyzed. The advantages of the proposed method are further verified by experiments. In this section, the structure of the DSEG is analyzed and the DSEG system configuration is illustrated. The advantages of the DSEG system are highlighted. The structure of the 12-8 pole DSEG is illustrated in Fig. 1. The rotor is made of lamination stacks. On the stator, field windings are around the phase windings to provide the magnetic field. The magnetic field, as well as the output voltage is controlled by regulating the field current. The simple structure of the DSEG leads to obvious advantages. The rotor is robust and suitable for high-speed applications in harsh environment. The design without permanent magnet makes the magnetic field easy to be controlled. The output voltage can be controlled by regulating the field current and the magnetic field can be rapidly restrained when a fault occurs" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001371_j.jmapro.2020.05.053-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001371_j.jmapro.2020.05.053-Figure3-1.png", "caption": "Fig. 3. The schematic of the layer positions in each sample for investigating the effect of deformation degree on the microstructural evolution.", "texts": [ " CW was produced in a similar way to AD, but after deposition of each layer and cooling down under 150\u2103, the deposited layer was cold worked by pressing in 350 bar pressure. The subsequent layer would deposit on the cold worked layer, and this procedure continued till CW would be built entirely, see Fig. 2. For investigating the effect of deformation degree on the microstructural evolution and hardness variations during cold working a layer and after the deposition of the subsequent layer, four samples were manufactured. As shown in Fig. 3 for preparing these samples, a layer with 200mm length was deposited, and cold worked with a constant force (F). Then, a sample was extracted with a 25mm length called XC. After that, a new layer was deposited on the remaining cold worked layer, and a sample with 25mm length was extracted that called XR. This procedure was continued to four steps (A, B, C, D) and after each step, the layer length reduced 50mm by extracting two samples(XC, XR). The letter \u201cX\u201d represents the step names and includes A, B, C, and D. Step by step, the pressure of cold working was getting higher because of constant force and reduction in layer top surface area. As shown in Fig. 3, every sample contains previous layers as well. The characteristics of each sample summarized in Table 2. For microstructural analysis, samples were cross sectioned perpendicular to the layer direction. After that, samples were polished and etched in an electrolyte solution containing 60mL HNO3 and 40mL H2O used at 1.1 V for 120 s. the microstructure of samples was characterized by using a light optical microscope and a scanning electron microscope (SEM). Energy dispersive spectrometry (EDS) were employed to measure micro-area composition", " [33], a fracture toughness can be determined as an area under the stress-strain curve resulted from the tensile test. The mean fracture toughness of CW sample decreased by 32 % in comparison to that of AD sample. The UTS, YS, and hardness of AD and CW samples were summarized in Table 3. As described in Table 2, AC and DC samples had the lowest and highest degree of the cold work, respectively. Fig. 12a represents the microstructure of the 1st layer in the AC sample after cold working with Pc pressure. By the deposition of the 2nd layer in AR sample (see Fig. 3), some regions of the 1st layer microstructure near the top layer boundary have been recrystallized that is shown in Fig. 12b. The 4th layer was cold worked by 4Pc pressure in DC sample. The microstructure of the cold worked layer in DC sample is shown in Fig. 13a. As seen in Fig. 13b, after depositing the 5th layer in sample DR, the cold worked microstructure has been recrystallized and the grain size has significantly become refined. As seen in Fig. 14, histograms represent the grain diameter distributions of AR, BR, CR and DR samples by classifying in diameter with 2 \u03bcm interval" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002212_tie.2020.3031535-Figure28-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002212_tie.2020.3031535-Figure28-1.png", "caption": "Fig. 28. Demagnetization ratio contour under three-phase short-circuit at 120\u00b0C.", "texts": [ "org/publications_standards/publications/rights/index.html for more information. defined to represent the decrease in remnant magnetic flux density. 0 (1 )*100(%)r r B B \u2212 (7) where Br0 is the pre-demagnetization remnant flux density and Br is the post-demagnetization value. Fig. 27 shows the currents under three-phase short-circuit condition at 120\u00b0C. The current peak value of the ST2 is only 1/30 of the SPMM, which means the PMs of ST2 are less likely to be demagnetized when three-phase short-circuit fault occurs. As shown in Fig. 28, both the machines have no risk of demagnetization under three-phase short-circuit condition. B. Under Pure D-axis Current at 120\u00b0C Armature current aligned with the negative d-axis causes more severe demagnetization than that aligned with the positive q-axis [35]. Hence, the q-axis current is set to 0 to investigate the irreversible demagnetization, and the demagnetizing currents last one electrical period. The critical demagnetizing MMF can be identified by increasing the stator current by 1pu per step" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000951_tro.2019.2936302-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000951_tro.2019.2936302-Figure7-1.png", "caption": "Fig. 7. Diagram of the mass\u2013spring\u2013damper system. The parameter m is equivalent to the total mass of the human body and the air floating device, k is the spring stiffness coefficient, and c is the damping coefficient.", "texts": [ " The elastic potential energy of the spring and the kinetic energy of the body are transformed into the internal energy through the viscous effect [23] of damping until the body velocity becomes zero and thus reaches a steady state. During the parking process, the human arms exhibit some elasticity as well as damping characteristics due to the contact force. The astronaut parking process can be interpreted as a dynamic coupling between a spring, damper, mass, and external force [24], [25]. The spring is responsible for buffering the collision impact force. The damper is used to prolong the collision time, which quickly consumes the kinetic energy of the mass prior to the collision. The mass\u2013spring\u2013damper system (see Fig. 7) is applied to model the dynamics of the human body parking process. The dynamic equation of the system [26] is as follows: mx\u0308+ cx\u0307+ kx = fcontact (3) where fcontact is the reaction force exerted on the mass by the handrail. Assume that the spring\u2013damper system has an initial state xt=0 = 0, x\u0307t=0 = 0, x\u0308t=0 = 0. At this point, the momentum of the entire system is mx\u0307 = 0. When the system is in contact with the handrail, according to the momentum theorem, we have the following: fcontact\u0394t = \u2211 (miv t+\u0394t i \u2212miv t i) (4) where \u0394t is the sampling interval" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000151_s10999-019-09479-5-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000151_s10999-019-09479-5-Figure2-1.png", "caption": "Fig. 2 A lubricated revolute joint in a planar system", "texts": [ " Under the proper initial conditions obtained from the kinematic simulation of the 3-PRR robot without lubricated joints, the vectors of acceleration \u20acq and Lagrange multiplier k can be solved directly from the above Eq. (3). Moreover, the vectors of position q and velocity _q can also be calculated by integrating of the acceleration vectors \u20acq at each integration time step. For an ideal revolute joint, the journal and bearing have a defined interaction of movement because of the coincident centres at all time steps. However, with regard to the lubricated revolute joint in Fig. 2, the centre of the journal is almost free to move relative to the centre of the bearing in a circle with the radius as the clearance size. A force constraint replaces the kinematic constraint in lubricated revolute joint. The interaction forces depend on not only the kinematic interaction of the journal in the bearing but also the hydrodynamic behaviours of the lubricant oil. The eccentricity vector that indicates the relative position of the centres of a journal and bearing in Fig. 3, is evaluated as e \u00bc rPj rPi \u00bc \u00f0rj \u00fe Tjr 0P j \u00de \u00f0ri \u00fe Tir 0P i \u00de \u00f04\u00de where rn (n \u00bc i; j) is the global coordinate vector of the centre of mass in XOY, while r 0P n is the local coordinate vector of the centre Pk of the lubricated revolute joint in okxkyk" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001750_j.triboint.2020.106186-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001750_j.triboint.2020.106186-Figure5-1.png", "caption": "Fig. 5. Overview of gas flow and dynamics of piston-ring assembly.", "texts": [ " There is a boundary condition that the gap equation (3) needs to satisfy: if h(x,y,t) equals to zero, the p\u00f0\u03be; \u03c2\u00de should be greater than zero, and if h(x,y,t) is greater than zero, the p\u00f0\u03be; \u03c2\u00de should be zero. Besides, the load balance is also required: W\u00f0t\u00de\u00bc ZZ \u03a9 p\u00f0x; y; t\u00dedxdy (5) The total contact load W(t) is predicted by ring pack simulation package developed by Liu [18,19]. Ring structural response simulation is based on the curved beam finite element method developed by Baelden [20], gas pressure distribution and oil transportation are also captured in this model (Fig. 5). The forces between the top two rings and liner include dry contact force and gas pressure. The hydrodynamic force is neglected because oil film thickness is assumed zero above TDC location of OCR [21]. Around combustion TDC, the top ring sits on the bottom of the groove due to high cylinder pressure. Thus the high cylinder pressure penetrates to inside the groove and pushes the top ring outwards to the liner elevating the contact stress between the top ring and liner (Fig. 6). The temperatures of the liner and ring sliding surface vary within an engine cycle at two different scales" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001397_1350650120942329-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001397_1350650120942329-Figure6-1.png", "caption": "Figure 6. Illustration of jet impingement depth on the gear (t0\u00bc 0).", "texts": [ "p ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 cos cos sin sin cos\u2019\u00f0 \u00de 2 q dI 2 z0 y sin\u2019 tan Ag \u00fe dI 2u tan g sin sin arc tan sin tan sin\u2019 \u00f020\u00de The derivation of the jet impingement depth dg on the gear is similar on the pinion. 1. At the initial moment (t0\u00bc 0): the position parameters of the spiral bevel gears and the jet streamline are as illustrated in Figure 6. To calculate efficiently, a new pinion coordinate system has been built and adopted, which is transformed from the original gear coordinate system by rotating around the y-axis and then translating along vector O0Op, where O0Op \u00bc Ap sin , 0,Ag Ap cos . The homogeneous coordinate of the new coordinate system is E \u00bc Trans Ap sin , 0,Ag Ap cos Rot y, 90 U \u00f021\u00de That is x y z 1 2 6664 3 7775 \u00bc cos 2 \u00f0 \u00de 0 sin 2 \u00f0 \u00de Ap sin 0 1 0 0 sin 2 \u00f0 \u00de 0 cos 2 \u00f0 \u00de Ag Ap cos 0 0 0 1 2 6664 3 7775 X Y Z 1 2 6664 3 7775 \u00bc cos 0 sin Ap sin 0 1 0 0 sin 0 cos Ag Ap cos 0 0 0 1 2 6664 3 7775 X Y Z 1 2 6664 3 7775 \u00f022\u00de The outside radius rIIa of the gear at the impingement point is rIIa 2 \u00bc X0 \u00fe Z0 ZII tan sin l 2 \u00fe Y0 Z0 ZII tan cos l 2 \u00f023\u00de where XII \u00bc X0 \u00fe Z0 ZII tan sin l YII \u00bc Y0 Z0 ZII tan cos l and XII,YII,ZII represent the impingement point position in the pinion coordinates" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001228_j.scitotenv.2020.138053-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001228_j.scitotenv.2020.138053-Figure6-1.png", "caption": "Fig. 6. The schematic of IEM-free and IEM", "texts": [ " Therefore, the improvement of electrochemical activity is conducive to the rapid usage of the substrate and the increase of system current, to provide enough electrons to the cathode for pollutant reduction. This ensures that DClNB can be reduced quickly in the system, so the pre-acclimated IEM-free MEC has a better pollutant removal performance. In a word, the mechanism of preacclimation strategy is mainly embodied in the increase of the proportion of living cells and the bioanode electrochemical activity due to the potential secretion of EPS. The traditional MEC is divided into anode and cathode chambers by IEM (Fig. 6). Although IEM could prevent the diffusion of pollutants to the bioanode to inhibit the bioanode, the resistance of the MEC would increase in the presence of IEM. Moreover, IEM led to a pH gradient in the MEC, i.e., an increase of pH in the cathode chamber and a decrease in the anode chamber. Low pH led to significant shocks to the electroactive microbes in the anode, which would increase the resistance of the systemanddecrease pollutant removal efficiency. Additionally,membrane foul is another possible problem in long-termoperation, which could further deteriorate reactor performance (Kim et al" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002264_s40430-020-02659-x-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002264_s40430-020-02659-x-Figure4-1.png", "caption": "Fig. 4 Relative positions of the j-th ball center and groove centers", "texts": [ " It can also be expressed as: where R j [39]: Force vector F of the inner raceway is added to each ball, and external load vector of the inner raceway of bearing is balanced; the balance equation of the bearing is obtained: or: (8)Fij = ij Mgj Db (9)Fej = ej Mgj Db (10)Q = [ Qr,Qz, T ]T (11)uj = [ urj, uzj, j ]T (12)uj = R j j (13)R j = \u23a1\u23a2\u23a2\u23a3 cos j sin j 0 \u2212zpj sin j zpj cos j 0 0 0 rpj sin j \u2212rpj cos j 0 0 0 \u2212 sin j cos j \u23a4\u23a5\u23a5\u23a6 T (14)F + n\u2211 j=1 fj = 0 The geometric relationship between the center of curvature of the raceway and the center of the ball is shown in Fig.\u00a04 when the bearing is stationary and after movement. According to the geometric relationship in Fig.\u00a04, loij and loej can be calculated: When the bearing is subjected to external load and running at high speed, the actual contact angles between ball and inner and outer raceways will no longer be equal, and the two contact angles can be obtained by the following formula. (15)F + n\u2211 j=1 RT j \u22c5 Qj = 0 (16)loij = rij \u2212 Db 2 \u2212 \u0394ij (17)loej = rej \u2212 Db 2 \u2212 \u0394ej (18)tan ij = loij sin 0 + uzj \u2212 vzj loij cos 0 + urj \u2212 vrj (19)tan ej = loej sin 0 + vzj loej cos 0 + vrj Journal of the Brazilian Society of Mechanical Sciences and Engineering (2020) 42:578 1 3 Page 5 of 24 578 where ij and ej represent the contact angles between the j-th ball and inner and outer raceway, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000113_j.finel.2019.103319-Figure8-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000113_j.finel.2019.103319-Figure8-1.png", "caption": "Fig. 8. Cook's skew beam test (L\u00bc 48.0; H\u00bc 44.0; S\u00bc 16.0; thickness\u00bc 1.0; E\u00bc 1.0; \u03bd\u00bc 1/3; F\u00bc 1.0).", "texts": [ " With respect to the in-plane load condition, TRIA\u00fe elements converged nicely to the reference solution whereas the others simulation packages failed to reach the satisfactory convergence. Moreover, the MITC4 exhibited the effect of the membrane locking when working in the in-plane load condition. Both for in-plane and out-of-plane conditions, QUAD\u00fe, S4 and Shell63 provided very efficient results since the convergence level was reached after just one refinement (\u201cQUAD 12 1\u201d in Figs. 6 and 7). As shown in Fig. 8, a skew cantilever beam is subjected to a shear load ults - in-plane load condition. applied at the free edge. Four different mesh refinements were tested as follows: elements were gradually increased from 2 to 16 on each side of the beam. Displacements were calculated at point P (Fig. 8). Fig. 9 represents the normalised displacements for QUAD and TRIA elements for varying the mesh size. The MITC4 slowly converged compared to others and QUAD\u00fe performed superior to other. Similar results are observed for TRIA\u00fe. Fig. 9. Cook's skew beam - normali The test represented a barrel vault roof under self-weight (Fig. 10). Complex states of membrane strain aroused in this test. Furthermore, the membrane locking phenomenon could not arise as the boundaries were not constrained along the z axis" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001289_tmech.2020.2992711-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001289_tmech.2020.2992711-Figure1-1.png", "caption": "Fig. 1. FLAM based active mooring line control for TLP-FOWT.", "texts": [ " The remainder of this paper is organized as follows. In Section II, the design of FLAM actuator is presented, along with its thermal analysis, dynamic modeling and a closed-loop force controller. Section III describes the AMLFC method with FLAM and FOWT modeling for LQR controller design. In Section IV, six load cases are simulated with a high-fidelity wind turbine model in FAST and the simulation result is analyzed. The paper is concluded in Section V. The FLAM actuation based AMLFC is illustrated with a schematic of FOWT with TLP platform in Fig. 1. The TLP platform features a cross-spoke design, in which a pair of tensioned mooring lines (legs) are assumed to be deployed at the tendon of each spoke, indexed as Lines 1 through 8 in the figure. The FLAM actuators placed at the junction between the mooring lines and the spoke enable active control of mooring line forces. By utilizing the spoke length as moment arm, the pitch and roll motions of the platform can thus be controlled by the FLAM actuator. It is noteworthy that deployment of FLAM actuators to all four tensioned legs allows platform stabilization control against wind and wave of virtually any directions", " The force outputs of individual FLAM actuator pairs i outF are controlled independently, which means the mooring line force controller works in a decentralized fashion from the wind turbine controller, with the feedback of the platform motion only. In other words, for this proof-of-concept study, the dynamic behavior of the wind turbine rotor and tower act as unmodeled disturbance for the AMLFC based platform control. Indeed, as a natural next step, centralized or distributed control strategy is desirable, which considers the wind turbine control and AMLFC in a coordinated fashion. It is noteworthy that there are 8 FLAM actuators in Fig. 1, which are grouped into 4 actuator pairs. As depicted in Fig. 8, the platform pitch and roll directions each has two actuator pairs which are manipulated by the respective AMLFC controller in the lumped fashion shown in Fig. 9. Therefore, the control-oriented modeling later will consider the actuation moment references (Mr, Mp) for the pitch and roll directions. To achieve balanced control effort and energy consumption of the actuators, an LQR is applied in the AMLFC. Previous research work [43-46] develops the platform motion equations from the first principles, which is complicated and timeconsuming considering nonlinear dynamics of wind turbine and complexity in hydraulic and wind forces" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001867_s00348-020-02946-2-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001867_s00348-020-02946-2-Figure1-1.png", "caption": "Fig. 1 Schematic of the air cannon with flex tip", "texts": [ " The balls are pitched into the still air using an air cannon that is capable of throwing the ball with desired speed and spin rate without damaging the texture of the ball. The air cannon uses a pneumatic linear accelerator to accelerate the ball at the desired speed. A flexible tip attached to the shaft of the cannon with high and low friction material on the opposite sides provides the rotation to the ball (Kensurd 2018). The rate of rotation is controlled by changing the compressive force on the ball or the length of the high friction material that the ball touches. Figure\u00a01 shows a schematic diagram of the air cannon with flex tip. The velocity and spin rate of the ball are measured from high-speed video over a length of 24\u00a0in. at 2000 frames per second. The standard deviation of the velocity and spin rate measurement is 0.4\u00a0m/s and 16\u00a0rpm, respectively. The PIV measurement location is 63\u00a0in. away from the ball release point; therefore, the velocity and spin rate do not change significantly at the measurement location after the ball is released. Two-component, planar particle image velocimetry (2CPIV) measurement is taken on the ball as it passes through the still air" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000906_j.robot.2019.103257-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000906_j.robot.2019.103257-Figure5-1.png", "caption": "Figure 5: The simulator set-up includes a pushing structure with two sliding rams (Ram 1 and 2) to exert arbitrary disturbances at random locations on the humanoid WALK-MAN for data collection.", "texts": [ " This data can be generated by applying different magnitude of disturbances at varied locations, until the robot falls, and storing the non-fall and fall data separately for further analysis. Figure 3a shows the steps involved in the determination of parameters. The following subsections first introduce the experimental setup, data processing, and finally, the determination of parameters from the processed data. Since it is highly risky to carry out both the fall and non-fall experiments with the real robot, and also it is extremely timeconsuming to perform a lot of experiments, a simulator setup is used. The setup considered is similar to the one used in our previous work.12 Figure 5 shows the setup, which includes a pushing structure and a humanoid, used to extract the raw data of the variables considered for each feature. The pushing structure exerts impact to the robot in the x-direction with Sliding Ram 1, and one can also change the height of the impact point in the z-direction with Sliding Ram 2, where their motion range can be determined by considering the robot\u2019s size. Impact exertion at various heights is necessary to observe the dynamic effects, such as the force required for toppling (lower force with increased height and vice versa), and feet slipping, which are in general caused by the non-uniform distribution of robot\u2019s mass in the feature variables" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001397_1350650120942329-Figure10-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001397_1350650120942329-Figure10-1.png", "caption": "Figure 10. Velocity streamlines under different jet elevation angles: (a) 95 ; (b) 97.5 ; (c) 100 ; (d) 110 ; (e) 120 .", "texts": [ " Likewise, the jet impingement depth on the pinion is the major concern in this paper. According to the initial layout and the constraints in the simulation, the nozzle spatial position is arranged at the initial point (210, 100, 40), and the jet azimuth angle is set to 270 . The jet elevation angle increases from 95 to 120 (see Table 2). Figure 9 illustrates the contour plots of the oil volume fractions. Obviously, Figure 9(b) exhibits a superior oil volume fraction on the pinion surface with the jet elevation angle of 97.5 . Figure 10 illustrates the velocity streamlines of the oil around the pinion, with which the flow field behavior can be directly observed. The rotation directions of the flow and the pinion are consistent. Obviously, the streamlines in Figure 10(b) are the most uniformly distributed for the best lubrication. It can also be seen that due to the effect of the spin flow caused by high-speed rotating gears, the jet oil streams deviate from their straight-line trajectories. Figure 11 demonstrates the behavior of the oil pressure and oil volume fractions in the plane parallel to the gear symmetry plane and near the meshing point. During the meshing process, the trends of oil volume fraction and oil pressure generally increase. Both the oil volume fractions and oil pressure reach maximum values under the jet elevation angle of 97" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000093_j.engfailanal.2019.104180-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000093_j.engfailanal.2019.104180-Figure6-1.png", "caption": "Fig. 6. Placement schematic of the loading bridge\u2019s elements.", "texts": [ " 3 shows the kinematics of the support of the loading bridge and Fig. 4 presents the point of support of the loading bridge on the undercarriage of the excavator. It is in this area where the mining product is transferred between the excavator\u2019s and loading bridge\u2019s belt conveyors. The failure took place while the machine was changing location. At the moment of the incident, the crawler-mounted loading bridge was at higher level than the excavator (Fig. 5). The placement schematic of the loading bridge\u2019s elements is shown in Fig. 6. The loss of stability took place due to crossing the first overturning edge located between the axle of six a wheel bogie and the axle of a swing girder (Fig. 7). It means that the location of the center of gravity moved outside the area of the stability field. Overturning edges define the boundaries of the field of stability. Due to the loss of stability, the loading bridge leaned onto the second overturning edge located between the drive wheel of the fixed girder\u2019s track and the axle of swing girder" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000093_j.engfailanal.2019.104180-Figure13-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000093_j.engfailanal.2019.104180-Figure13-1.png", "caption": "Fig. 13. Location of exemplary regions for which the yield strength was exceeded.", "texts": [], "surrounding_texts": [ "The strain gauge testing was aimed at determining variable loads being the derivative of dynamic phenomena happening during the operation of the machine and its movement. Knowing the real values of stress in deformed cross-sections, it is possible to estimate the expected durability and the structural nodes\u2019 ability to work. Measurement points have been defined on both sides of the crawler-mounted loading bridge at its top and bottom girders. The measurement cross-section has been defined in a place with uniform load distribution (read from the numerical model) at a small distance from the deformed nodes (Fig. 15). A series of measurements aimed at registering changes in strain has been conducted both during excavation and movement of the machine. Thanks to that, it was possible to chart the range of load variations for all occurring operation cases. It is important to note that the registered loads do not depict the values of static loads from dead weight of excavator\u2019s elements. Eventually, 3 measurements during normal operation and 6 runs in specific configurations of the position of the superstructure relative to the undercarriage (in parallel and perpendicularly to the direction of movement) have been conducted. Table 3 presents a detailed comparison. Fig. 16 presents examples of diagrams of stress registered in chosen measurement cases. The analysis of measurement data shows small values of stress changes both for mass loads (changing the configuration of the machine, mining product) and dynamic loads (vibrations). There is no possibility dynamic phenomena could result in plastic deformation of the load-carrying structure. Average values oscillate around 20MPa. The highest registered maximum equals roughly 37MPa. Table 3 presents the widest ranges of global stress for each measurement case. The conducted tests show that variable operation loads are not significant from the point of view of ultimate tensile strength and fatigue strength of the load-carrying structure of the crawler-mounted loading bridge. The maximum value of the range of stress registered during the measurements is lower than the lowest specification limits for welded joints, which confirms the safety of the loading bridge\u2019s structure in the area of its permanent deformations caused by the loss of stability at the caterpillar undercarriage." ] }, { "image_filename": "designv11_5_0001665_iros45743.2020.9341101-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001665_iros45743.2020.9341101-Figure1-1.png", "caption": "Figure 1. Mechanism of the Ankle-foot prosthesis end-effector: (Top) CAD model with detailed description, and (Bottom) Hardware implementation. We printed strain relief assembly, toes (white part), etc. We improved the robustness of the heel (purple part) using carbon fiber. Pulling the inner cable provided plantarflexion, releasing the inner cable and using rubber bands between screws on toes and the main part was for dorsiflexion.", "texts": [ "0 0 \u00a9 20 20 IE EE | D O I: 10 .1 10 9/ IR O S4 57 43 .2 02 Authorized licensed use limited to: Robert Gordon University. Downloaded on May 26,2021 at 20:10:04 UTC from IEEE Xplore. Restrictions apply. presented our controller implementation, including human-inthe-loop optimization, and then a pilot study with two participants. In section III, we described results and discuss the outcomes. The two degrees of freedom ankle-foot prosthesis [15] was redesigned and manufactured as an end-effector for a tethered emulator system (Fig.1). The ankle-foot prosthesis is composed of a frame, two toes, and a compliant heel. The main body of the ankle-foot prosthesis (including the toes and frame) consists of 7075-T6 aluminum and was machined with a computer numerical control (CNC) method. Contained within the frame are needle bearings, which are press-fit, allowing two notched shafts to rotate about the same axis. Each shaft is then fixed to a single toe by the use of a set screw. The rotation of each shaft-toe assembly is similar to the dorsiflexion and plantarflexion of a human ankle joint" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002450_j.compstruct.2021.113608-Figure8-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002450_j.compstruct.2021.113608-Figure8-1.png", "caption": "Fig. 8. Three typical finite e", "texts": [ " At the same time, after the above simplified analysis, the contact surface at the contact point can be described with the directional curvature, and then only the deformation of the roller at the extreme position of the directional curvature of the surface need to be analyzed, which will greatly improve the efficiency of finite element analysis. The general problem of contact between the compaction roller and the curved surface can be transformed into three forms through analysis above as shown in Fig. 8. The directional curvature mentioned above refers to the normal curvature of a curved surface along a specific direction at a certain point. And it is calculated according to the maximum and minimum principal curvature at that point using the Euler's theorem. kdir \u00bc kmax cos2\u03c8 \u00fe kmin sin2\u03c8 \u00f024\u00de Where \u03c8 represents the angle between the specified direction and the direction of the maximum principal curvature. Considering that the ply angles are generally constant, of which 0, 90 and \u00b1 45 are the most common four angles" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001397_1350650120942329-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001397_1350650120942329-Figure2-1.png", "caption": "Figure 2. Illustration of jet nozzle layout position and orientation parameters for spiral bevel gears.", "texts": [ " The basic idea for calculating oil jet impingement is: first, an initial state or moment (t0\u00bc 0) is set when the outer edge of the gear top passes, and through geometric calculations the positions of the gear and the pinion in this state or moment are determined. Then, after a period of time (t1\u00bc t), the positions of the gear, pinion, and jet flow were determined. Finally, since the jet stream flowing time is equal to the gear rotation time from the initial state (t0\u00bc 0) to the final state (t1\u00bc t), the impingement points on the meshing surfaces of the gear and the pinion can be determined, respectively, by solving the combined equations. Obviously, the depth is directly related to the oil jet orientation parameters x0, y0, z0, , \u2019, as shown in Figure 2. The x0, y0, z0 denotes the nozzle exit position; the elevation angle represents the angle between the jet flow and the z direction, which is always restricted to /24 4 ; and the azimuth angle \u2019 represents the angle between the jet stream projection line on the xO0y plane and the x direction, which is restricted to 4 \u20194 2 , and the point O0 is defined as gear coordinates origin. To identify the constraints for the oil jet orientations, the equivalent helical cylindrical gear of the spiral bevel gear is adopted as shown in Figure 3" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001464_issc49989.2020.9180186-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001464_issc49989.2020.9180186-Figure1-1.png", "caption": "Fig. 1. Parallel arrangement of multi-sensor (camera, LiDAR and Radar) used for environment perception in autonomous vehicles for large heavy vehicles off-road in industrial environments.", "texts": [ " The first phase of our research evaluates the potential to synchronise sensor data where external reference clocks are not conveniently accessible. III. SENSORS This section discusses the benefits and limitations of the sensors used in our proposed sensor system. Here we consider a Falcon-IQ Internet Protocol (IP) industrial zoom camera of the model EZIP-T030(E), a Velodyne LiDAR Ultra Puck VLP-32C sensor, and a SmartMicro automotive radar UMRR-96 Type 153 sensor in our proposed systems. The arrangement of multiple sensors on large heavy vehicles off-road is shown in Fig. 1. Cameras are generally reliable and relatively inexpensive. Vehicle applications including advanced driver assistance systems rely on cameras to perform high resolution tasks such as environmental perception, object classifications and colour perception. The output of a conventional \u2018RGB\u2019 camera is typically an array of several hundred by several hundred (or more) \u2018pixels\u2019, each pixel having a red, green and blue intensity level [13]. Image noise may include periodic moire-type patterns cause by imperfect lenses, filters and image sensors [12] and most camera systems will have some degree of image distortion which may require calibration [29-30]" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001037_tec.2019.2951659-Figure16-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001037_tec.2019.2951659-Figure16-1.png", "caption": "Fig. 16: Motor TestRig Setup", "texts": [ " However, the shift direction in the voltages has a similar behavior for the tested faults regardless of the operating conditions. This makes the proposed method valid under different operating conditions. A voltage source inverter with Field Oriented Control (FOC) is used to operate and control the tested motor. Fig. 15 shows the basic block diagram for the FOC. For this type of controller, the commanded voltages (V cd and V cq ) are internally available and are used to detect the fault type and estimate the severity. Fig. 16 shows the motor test setup. A National Instrument Real Time LabVIEW and FPGA system is used to operate and control the tested motor. Two rotor faults were introduced and studied experimentally, RGS (positive and negative) and DIR. RGS is implemented by controlling the motor airgap for four severities in each direction (20, 40, 60, and 80%). DIR is implemented during the assembly of the motor in which the rotor was tilted in two severities: 40% and 60% (i.e. \u03b5tilt = 40% and 60%). Fig. 17 shows the experimental results for the change in the commanded voltages under the two tested faults" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001879_tits.2020.2983522-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001879_tits.2020.2983522-Figure2-1.png", "caption": "Fig. 2. Definition of the body coordinate RU and Euler angles.", "texts": [ " (45) Thus, the proof is complete. In this section, the state equation of a multi-rotor UAV system and its parameters are firstly presented. Additionally, in order to achieve the control target of the multi-rotor UAV, three types of controllers need to be designed, e.g., attitude angle controllers, height controller and position controllers. In order to set up the dynamic model of the UAV, we takes a four-rotor UAV as an example. The coordinate system RU of the UAV is firstly established as shown in Fig. 2. Furthermore, detailed descriptions about the coordinate of the UAV are presented as follows. \u2022 The origin of RU is at the center of the UAV; \u2022 The X-axis is along the direction of the arm \u2460; \u2022 The Y-axis is along the direction of the arm \u2463; \u2022 The Z-axis passes through the origin, perpendicular to the XOY plane, forming a right-handed coordinate system; \u2022 The roll angle \u03c6 denotes the rotation angle of the X-axis; \u2022 The pitch angle \u03b8 denotes the rotation angle of the Y-axis; \u2022 The yaw angle \u03c8 denotes the rotation angle of the Z-axis" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000676_s00170-018-03220-w-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000676_s00170-018-03220-w-Figure1-1.png", "caption": "Fig. 1 Four types of sub-processes of MLDD", "texts": [ "\u201d Therefore, MLDD can be further divided into four types of subprocesses based on its forming characteristics, named singletrack single-layer deposition (SSD) process (usually called laser cladding process), single-track multi-layer stacking * Haiying Wei why@hnu.edu.cn 1 State Key Laboratory of Advanced Design and Manufacturing for Vehicle Body, Hunan University, Changsha 410082, China 2 Key Laboratory for Intelligent Laser Manufacturing of Hunan Province, Hunan University, Changsha 410082, China (SMS) process, multi-track single-layer lapping (MSL) process, and multi-track multi-layer accumulation (MMA) process, as shown in Fig. 1. Compared to conventional manufacturing processes, MLDD holds many promising advantages, three stand out: & Metal parts with excellent mechanical properties can be realized, because it can promote grain refinement due to the characteristic of rapid cooling of molten pool [14]. & Complex metal parts can be manufactured with a high degree of freedom design, and extended parts\u2019 life can be achieved through repair or refurbishment [15, 16]. & Design time can be shortened and small batch customization products can be rapidly manufactured [17]" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002461_j.procir.2021.01.056-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002461_j.procir.2021.01.056-Figure1-1.png", "caption": "Fig. 1. Scheme of the linear winding process [11] (left) and difference between wild, layer, and orthocyclic winding [15] (right)", "texts": [ " According to DIN 8580 [14], winding is classified as a part of the main group \u2018Joining\u2019 and categorized in the subgroup \u20184.5.1 Joining by forming wire-shaped bodies\u2019. The norm defines winding as joining a core, e.g. a bobbin, with a wire by continuously bending around the core. The main difference between the various winding processes is in the way of applying the wire to the bobbin. In linear winding, the bobbin to which the wire is fixed is rotated by a spindle, while the wire guide is moved parallel to the rotation axis from one flange to another (Fig. 1, left) [11]. Depending on the requirements for coil quality and production costs, coils are manufactured in different winding schemes, such as a wild, layer, or orthocyclic winding (Fig. 1, right) [15]. The degree of utilization of the available winding space is described by the mechanical fill factor [11]. If one only considers the copper without the insulation layer, the so-called copper filling factor can be calculated alternatively [11]. Due to the undefined wire placement, a wild winding comes with a low mechanical fill factor and high variations in wire length [9]. As can be seen in Fig. 1, the wires are closest together in the orthocyclic winding, resulting in the highest achievable mechanical fill factor [15]. In contrast to the helical layer winding, the wires of the upper layer are laid into the valleys of the lower layer with orthocyclic winding [15]. For rectangular bobbins, the orthocyclic arrangement can only be realized on three sides; on the fourth side, the winding meets the incoming wire and must therefore perform a turn step [9]. With stator segment and single tooth bobbins, this turn step must be performed on the short side, i", " With linear winding machines, the wire is usually drawn from a barrel and passed through a wire tension control system, consisting of straightening rolls, a brake wheel, and a balancing system [10]. The tension control system, which is often simply called wire brake, applies a preset wire tension force and compensates for free wire lengths that occur when winding noncircular bobbins due to the cyclically changing wire speed [10]. While insufficient wire tension leads to loose wires, excessive tension causes plastic deformation, resulting in a tapered wire cross-section [11]. As labeled in Fig. 1, the angle between the axis of the wire guide and the tensioned wire is defined as the caster angle [11]. A positive caster angle ensures that the coils are braced against each other, preventing the formation of gaps between the turns. If a maximum caster angle is exceeded, the accurate positioning of the wire is no longer possible [11]. If a wire slips, defects like a double winding, loose wire, gap, or crossing can occur as summarized in [9]. Other common defects in the layer structure are a faulty flange winding or layer step" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002011_j.jmapro.2020.05.026-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002011_j.jmapro.2020.05.026-Figure5-1.png", "caption": "Fig. 5. Schematic diagram of the LMD system with deposition height control devices.", "texts": [ " When the nozzle is inclined, the drag forces on the powder flow from both carrier gas and shielding gas to the same axial direction can reduce the action of the gravity significantly and achieve a better straightness of the powder flow. With an inclined angle of 135\u00b0 in Fig. 4c, the powder flow of the coaxial powder feeding nozzle diverges due to gravity. In contrast, the powder flow of IBPF nozzle maintained a slight divergence and excellent straightness out of the powder pipe outlet, and exhibited almost no parabolic trajectory. The schematic diagram of the LMD system is illustrated in Fig. 5. The system consisted of an IPG YLS-2000-TR fiber laser as the heat source, a GTV PF 2/2 powder feeder, a robot of KUKA KR 60-3 F and a Beckhoff EK1100 Ethernet I/O. In order to measure the actual deposition height, a CCD camera was fixed on the IBPF nozzle, which is capable to move and rotate with the nozzle synchronously (Figs. 4b and 5). Following the frames acquisition by the image card, the height value calculation was processed in an industrial PC. A real-time ProfiNet communication was installed between the Industrial PC and the robot controller to build a closed-loop control system. The control algorithms were programmed in the KUKA KRL software. In general, a closed loop control system requires a good real-time responsiveness. In Fig. 5, except for a few seconds delay for changing of the powder feeding rate, the response times for the change of scanning speed, laser power, and defocus distance are all in the millisecond level. The both programs in the robot control cabinet and industrial computer work in independence threads with a real-time data transmission through ProfiNet. The substrate material used was 304 stainless steel. The powder material was the Fe-based alloy Fe313 (composition in wt,\u2212%: C\u22480.1; Si= 2.5\u20133.5; Cr= 13\u201317; B=0" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001525_icra40945.2020.9197565-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001525_icra40945.2020.9197565-Figure3-1.png", "caption": "Fig. 3. AMPRO II\u2013a custom-built powered transfemoral prosthesis", "texts": [ " The dorsiflexed swing equilibrium angle ensures foot clearance to avoid trips. The equilibrium angle for Set D resembled a foot-drop condition\u2013the state a human foot would conform to when physically unconstrained. It was anticipated that the footdrop condition would pose a challenge during swing phase. It is likely that certain compensatory actions will be needed to ensure sufficient foot clearance during swing. The proposed sets of impedance parameters were tested on a custom-built powered transfemoral prosthesis shown in Fig. 3. AMPRO II (Fig. 3) has an actuated ankle and knee joint, and a passive spring-loaded toe joint. While the proposed impedance controller was implemented at the ankle, a previously published controller\u2013a hybrid of impedance and trajectory tracking\u2013was used to manipulate the knee. The latter has been discussed in [28]. The prosthesis was operated under a time-based scheme that utilizes a parameter that linearly increases from 0 to 1 as the gait progresses from 0% to 100%. This parameter is used to identify the progress in the gait cycle", " Since the focus of this study was to evaluate the performance of continuously varying impedance parameters, it was preferred the performance be unaffected by the possible inaccuracies of state-based control. Further, it was unclear whether the stiffness at AMPRO II\u2019s toe joint would impact the ankle\u2019s performance. To study the effect of each impedance controller, in an isolated manner, the toe joint was restrained using a rigid element. Future studies will investigate the impact of toe stiffness on the generated gait. To validate the proposed idea, an indoor experiment was designed using the aforementioned powered prosthesis in Fig. 3. A healthy young subject (male, 170cm, 70kg) participated in the experiment using an L-shape simulator that helped emulate prosthetic walking. The subject was asked to walk on a treadmill at his preferred walking speed (0.7 m/s). The subject\u2019s safety was assured by handrails located on either side of the treadmill. The experiment protocol has been reviewed and approved by the Institutional Review Board (IRB) at Texas A&M University (IRB2015-0607F). The recruited subject showed considerable height difference between his limbs while wearing the prosthesis" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000457_1.4033888-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000457_1.4033888-Figure2-1.png", "caption": "Fig. 2 Boundary conditions for rotor thermal deformation", "texts": [ " Current research abandons the assumptions in earlier studies where the journal temperature is linear radially and uniform axially, and the thermal deflection is estimated by certain analytical expressions [10\u201316]. Boundary conditions should be carefully handled to provide realistic deformation and avoid singular matrixes in the FEM solution. Considering that the temperature rise is negligible outside the TR, the TR is assumed to be a cantilever with fixed degreesof-freedom (DOFs) on the right section of TR1 and left section of TR2, which is illustrated in Fig. 2. Note that on the left section, the nodal DOFs are fully fixed for the innermost layer and are fixed axially for other nodes. TR1 has similar boundary conditions as TR2. The rotor thermal bow induced by the asymmetric thermal load is shown in Fig. 3. Note that the rotor centerline deviates from the horizontal position due to the noneven radial expansion. For nodes located between both TRs, the thermal bow is zero (dashed rectangle); for nodes within TRs, shaded in Fig. 3, the nodal thermal bow is directly calculated by the FEM and expressed with \u00f0eTn;UTn\u00de for magnitude and phase; for nodes close to both rotor ends, the thermal bow is the linear extension of the TR end" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001981_s10846-020-01213-0-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001981_s10846-020-01213-0-Figure1-1.png", "caption": "Fig. 1 ODQuad overall assembly (a), roll joint axis (b), pitch joint axis (c), grasping tool (d)", "texts": [ " The use of omnidirectional platforms in lieu of standard quadrotors, is suggested by their capability of exerting both forces and moments at the contact points, without modifying the orientation of the load, and by reacting to external disturbances in a more effective way. The effectiveness of omnidirectional platforms in performing contact tasks has been experimentally verified in [4]. More in detail, the ODQuad aerial vehicle, firstly presented in [17], is adopted. It is composed by the following components, sketched in Fig. 1: \u2013 A platform, where computing hardware, electronics, batteries and sensors are hosted; \u2013 A rotor frame, hosting the propellers; \u2013 A mobile frame, linking the two previous components; \u2013 A grasping tool attached to the platform. Two rotational joints, with orthogonal axes, connect the platform to mobile frame (roll joint, see Fig. 1b) and the mobile frame to the rotor frame (pitch joint, see Fig. 1c). This aerial vehicle is able to move in all directions without changing the platform attitude, but only modifying the thrust direction by acting on the roll and pitch joints. In order to grasp and move the object, let us assume that a grasping tool (see Fig. 1d) rigidly connects the platform to the object and that a force-torque sensor is mounted on the tool, in order to measure the contact wrench between the object and the ODQuad. Let us define the following coordinate frames (see Fig. 2): \u2013 The inertial coordinate frame, {O, x, y, z}; \u2013 The coordinate frame k{Ok, xk, yk, zk}, attached to the kth robot platform (k = 1, 2, . . . , N); \u2013 The coordinate frame ek {Oek , xek , yek , zek }, attached to the tool and with origin at the contact point; \u2013 The coordinate frame o{Oo, xo, yo, zo} attached to the object center of mass" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000093_j.engfailanal.2019.104180-Figure11-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000093_j.engfailanal.2019.104180-Figure11-1.png", "caption": "Fig. 11. Discrete model of the crawler-mounted loading bridge with mass elements positioned in the same way as at the moment of failure.", "texts": [], "surrounding_texts": [ "Numerical analysis of the event has been conducted with the use of the finite element method. A series of numerical models of the loading bridge and its assembly with the excavator have been prepared. They have been built based on the technical documentation of the machine and measurements conducted on the real object. Elements of the caterpillar undercarriage have been modeled by applying adequate boundary conditions. The discrete model has been composed of both one- and two-dimensional finite elements. The areas located far from the damage have been modeled with the use of beam elements and the damaged regions with shell elements. The model of the loading bridge has been positioned in the same way relative to the excavator as at the moment of failure. It has also included the location of various subassemblies fixed to the load-carrying structure, such as operator\u2019s cabin, transformer stations, drive elements, mechanical elements of a conveyor located on the bridge, and moveable booms of the loading bridge. The elements of the loading bridge that have not been included in the geometrical model have been modeled with the use of mass elements. Thanks to that, the mass of the object and the location of its center of gravity conforming with the machine\u2019s documentation have been obtained. Figs. 11 and 12 present the discrete model with the location of mass elements. The presented discrete models have been used to perform a numerical analysis of the crawler-mounted loading bridge and the part of the excavator\u2019s undercarriage. Calculations have been conducted for two cases in which different loads were acting upon the loadcarrying structure of the crawler-mounted loading bridge. The machine\u2019s load was reflected by static loads (elements that are fixed to the structure of the excavator), the load exerted by mining product located on belt conveyors, and the load exerted as the result of encrustation. The crawler-mounted loading bridge was also under the force caused by the tilt of the machine. The last load was caused by displacement. Table 1 presents in which cases the particular load was taken into account. Calculations have been conducted for two cases: loss of stability and normal operation. All loads for the normal operation case were defined according to the standard [6]. Table 2 presents the results of analyses represented as the maximum values of stress. Figs. 13 and 14 presents regions for which their yield strength has been exceeded." ] }, { "image_filename": "designv11_5_0000471_j.ijsolstr.2016.07.004-Figure23-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000471_j.ijsolstr.2016.07.004-Figure23-1.png", "caption": "Fig. 23. The minor principal stress (MPa, M/ (2 \u03c0\u03c30 a 2 t) = 2 . 9 ).", "texts": [], "surrounding_texts": [ "C\nC\n\u03d5\nw\n\u03d5\n6\nt t t r a m e D r\nfollowing equations:\n1\nA\n+\n1\nB\n\u2212 ln\n( B\nA\n) \u2212 2\n3\n= 0 (63)\n2 1 \u2212 [ 1 + 2 C 1 ( a\nb\n)2 ]2\nR\u0304 4 + M\u0304 2 = 0 (64)\n2 = { R\u0304 + [ 1\nR\u0304\n+ 2 \u0304R\n( a\nb\n)2 ]\nC 1\n}2\n+\n( M\u0304\nR\u0304\n)2\n(65)\nwhere\nM\u0304 =\nM\n2 \u03c0a 2 \u03c30 t , R\u0304 =\nR a , A = C 2\nM\u0304\n2 \u2212 1 , B = R\u0304\n2 C 2\nM\u0304\n2 \u2212 1 (66)\nIt is clear that Eqs. (63) \u2013( 65 ) are nonlinear and can only be numerically solved. After the variables such as M\u0304 , R\u0304 , C 1 , C 2 are obtained from the equations, the nondimensionalized major and minor principal stresses and the torsional rotation can be derived according to Eqs. (67) \u2013( 69 ) as\n\u03c31 \u03c30 =\n\u23a7 \u23a8 \u23a9 C 2 / r \u221a C 2 \u2212M 2 / r 2\nr < R\n2 C 1 a 2 / b 2 + 1 +\n\u221a\nC 2 1 + M\n2\n2 r > = R\n(67)\nr\n\u03c32 \u03c30 =\n{\n0 r < R\n2 C 1 a 2 / b 2 + 1 \u2212\n\u221a\nC 2 1 + M\n2\nr 2 r > = R\n(68)\n\u00af =\n3 M\u0304\n8 ( 1 \u2212 a 2\nb 2\n)[ 1 R\u0304 2 \u2212 1\nB\n+ ln\n( B\nA\n) + 1\nR\u0304\n2 \u2212 8\n3\n( a\nb\n)2\n+\n5\n3\n] (69)\nhere\n\u00af =\n3 E\u03d5\n4 \u03c30\n( 1 \u2212 a 2\nb 2\n) (70)\n.2.1. Verification of accuracy\nIn this sub-section, the annular membrane is analyzed with he new wrinkling model and the results are compared with he theoretical solutions obtained from Eqs. (63) to ( 65 ), ( 67 ) o (69). The finite element model is shown in Fig. 20 . Its outer adius b and inner radius a (shown in Fig. 19 ) are 1250 mm nd 500 mm, respectively. The thickness t is 1 mm. The Young\u2019s\nodulus E = 10 0 0 MPa and the Poisson\u2019s ratio \u03c5 =1/3. The inner dge is restrained in the radial direction and the circumferential OFs of nodes on the edge are coupled to have the same torsional otation. The outer edge is firstly pretensioned to introduce a", "u c t\nr r a m p M i a s c b p r o a p a\nw fl v t\n6\nm ( 0\n8 As for the numerical examples in Section 6.2 , their geometrical and physical parameters are the same as given in Section 6.2.1 except the quantities assigned specifically.\nniform radial stress \u03c3 0 and then fixed along the radial and ircumferential directions. With a twisting moment M applied to he inner edge, wrinkles arise and their radii increase gradually.\nThe relation between the twisting moment and the torsional otation is given in Fig. 21 . In the figure, the horizontal axis repesents the dimensionless torsional rotation (please see Eq. (70) ) nd the vertical axis denotes the dimensionless twisting mo-\nent (please see Eq. (66) ). Contours of the major and minor\nrincipal stresses and the wrinkling strain corresponding to\n/ (2 \u03c0\u03c30 a 2 t) = 2 . 9 in Fig. 21 are illustrated in Figs. 22\u201324 , where t can be found that the principal stresses and the wrinkling strain re all uniformly distributed along the circumferential direction. It ignifies that the variation of stresses and strains along the radial oordinate may be typical of their distribution over the mem-\nrane. Fig. 25 illustrates the distributions of the major and minor rincipal stresses along the radial direction with R\u0304 = 1 . 2 and 1.6 espectively. The horizontal axis in the figure represents the ratio f the radial coordinate r to the inner radius a and the vertical xis denotes the major or minor principal stresse ( \u03c3 1 or \u03c3 2 ) to the restress \u03c3 0 . Figs. 21 and 25 indicate that the wrinkling model is ccurate and the results agree well with the theoretical solutions.\nFig. 21 also manifests that the major principal stress in the inkled region is much larger and its distribution curve becomes at in the taut area. Conversely, the minor principal stress almost anishes in the wrinkled region while it increases gradually with he radial coordinate in the taut zone.\n.2.2. Influence of the prestress\nThe annular membrane shown in Fig. 19 subjected to a twisting\noment M = 10,0 0 0 N \u2022mm is analyzed 8 with different prestresses \u03c3 = 0 . 1 \u03c4, 0 . 2 \u03c4, 0 . 3 \u03c4, 0 . 4 \u03c4 and 0.5 \u03c4 , where \u03c4 is the uniformly", "o u m\n\u223c s i r\ndistributed shear stress along the circumferential direction at r = a and obtained from the boundary condition with \u03c4 = M\n2 \u03c0a 2 t ).\nThe influences of the prestress on the major and minor principal stresses and the wrinkling strain are illustrated in Fig. 26 (a) and (b), respectively. The vertical axis in Fig. 26 (a) represents the ratio of the principal stress \u03c3 1 or \u03c3 2 to the shear stress \u03c4 and the vertical axis in Fig. 26 (b) denotes the ratio of the wrinkling strain \u025b w to the shear strain \u03b3 ( \u03b3 = \u03c4 G , G is the shear modulus). The solid\nr dashed lines in Fig. 26 (a) represent the major principal stresses nder different prestresses while the discrete symbols denote the\ninor stresses.\nFig. 26 (a) shows that when the prestress increases from 0.1 \u03c4 0.5 \u03c4 , the major and minor principal stresses are augmented lightly in the taut area. Fig. 26 (b) indicates that the prestress s able to decrease the wrinkling strain notably in the wrinkled egion." ] }, { "image_filename": "designv11_5_0001225_indicon47234.2019.9030298-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001225_indicon47234.2019.9030298-Figure3-1.png", "caption": "Fig. 3. The mobility model.", "texts": [ " Firstly, relations between these two probability density functions are given as follows: f V,\u03c6|V2( v,\u03d5|v2) = f V1,\u03b8 |V2 (v1, \u03d1 | v2) J = fV1,\u03b8 (v1, \u03d1) J ; (7) since V1 and \u03b8 are independent of V2, and where J = \u2223\u2223\u2223\u2223\u2223 \u2202v \u2202v1 \u2202v \u2202\u03d1 \u2202\u03d5 \u2202v1 \u2202\u03d5 \u2202\u03d1 \u2223\u2223\u2223\u2223\u2223 . Next, we calculate the Jacobian as \u2202v \u2202v1 = v1 \u2212 v1cos\u03d1\u221a v2 1 + v2 2 \u2212 2v1v2cos\u03d1 ; \u2202v \u2202\u03d1 = v1v2sin\u03d1\u221a v2 1 + v2 2 \u2212 2v1v2cos\u03d1 \u2202\u03d5 \u2202v1 = \u2212v2sin\u03d1\u221a v2 1 + v2 2 \u2212 2v1v2cos\u03d1 ; \u2202\u03d5 \u2202\u03d1 = v2 1 \u2212 v1v2cos\u03d1\u221a v2 1 + v2 2 \u2212 2v1v2cos\u03d1 J = v1\u221a v2 1 + v2 2 \u2212 2v1v2cos\u03d1 = v1 v = \u221a v2 + v2 2 + 2vv2cos\u03d5 v Thus we get: f V,\u03c6|V2( v,\u03d5|v2) = v 2\u03c0Vd \u221a v2 + v2 2 + 2vv2cos\u03d5 . (8) As defined before \u03b1, represents the original direction of Node 1 relative to Node 2 (as shown in Fig. 3(b)). Furthermore, as indicated before, \u03b1 is uniformly distributed and is independent of \u2212\u2192 V2, \u2212\u2192 V, and \u03c6. Thus, we have: f\u03b1 (\u03b1) = 1 2\u03c0 = f \u03b1|V2 (\u03b1| v2) , 0 \u2264 \u03b1 \u2264 2\u03c0 (9) f \u03b1,V,\u03c6|V2 (\u03b1, v, \u03d5| v2) = f \u03b1|V2 (\u03b1| v2) \u00b7 f V,\u03c6|V2 (v, \u03d5| v2) = v 4\u03c02Vd \u221a v2 + v2 2 + 2vv2cos\u03d5 (10) According to Fig. 3(b), RL L1 = R1cos (\u03c6 + \u03b1) + \u221a R2 \u2212 R2 1sin2 (\u03c6 + \u03b1) V (11) Assuming, without loss of generality, that \u03b1 = 0, RL L1 = R1cos (\u03c6) + \u221a R2 \u2212 R2 1sin2 (\u03c6) V (12) F RL L1|V2 (\u03c41| v2) = P ( RL L1 \u2264 \u03c41| V2) = P \u239b \u239d R1cos (\u03c6) + \u221a R2 \u2212 R2 1sin2 (\u03c6) V \u2264 \u03c41| V2 \u239e \u23a0 = P \u239b \u239dV \u2265 R1cos (\u03c6) + \u221a R2 \u2212 R2 1sin2 (\u03c6) \u03c41 \u2223\u2223\u2223\u2223\u2223\u2223 V2 \u239e \u23a0 (13) = \u03c0\u222b \u2212\u03c0 v2+Vmax\u222b R1cos(\u03c6)+ \u221a R2\u2212R2 1 sin2(\u03c6) \u03c41 f V,\u03c6|V2 (v, \u03d5| v2) dvd\u03d5 = \u03c0\u222b \u2212\u03c0 v2+Vmax\u222b R1cos(\u03c6)+ \u221a R2\u2212R2 1 sin2(\u03c6) \u03c41 v 2\u03c0Vd \u221a v2 + v2 2 + 2vv2cos\u03d5 dvd\u03d5 f RL L1|V2 (\u03c41| v2) = d F RL L1|V2 (\u03c41| v2) d\u03c41 = d d\u03c41 \u03c0\u222b \u2212\u03c0 v2+Vmax\u222b R1cos(\u03c6)+ \u221a R2\u2212R2 1 sin2(\u03c6) \u03c41 v 2\u03c0Vd \u221a v2+v2 2+2vv2cos\u03d5 dvd\u03d5 (14) \u03c0\u222b \u2212\u03c0 d d\u03c41 \u239b \u239c\u239c\u239d v2+Vmax\u222b R1cos(\u03c6)+ \u221a R2\u2212R2 1 sin2(\u03c6) \u03c41 v 2\u03c0Vd \u221a v2+v2 2+2vv2cos\u03d5 dv \u239e \u239f\u239f\u23a0 d\u03d5 Next, we use the Leibniz integral rule on (14) to obtain: f RL L1|V2 (\u03c41| v2) = \u03c0\u222b \u2212\u03c0 ( R1cos\u03d5 + \u221a R2 \u2212 R2 1sin2\u03d5 )2 2\u03c0Vd\u03c4 2 1 \u221a\u221a\u221a\u221a\u221a\u221a\u221a ( R1cos\u03d5 + \u221a R2 \u2212 R2 1sin2\u03d5 )2 + v2 2\u03c4 2 1 + 2v2\u03c41cos\u03d5 ( R1cos\u03d5 + \u221a R2 \u2212 R2 1sin2\u03d5 ) d\u03d5 (15) Formulation of f RL L2|V2 (\u03c41| v2) can be derived by following the same process" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000951_tro.2019.2936302-Figure13-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000951_tro.2019.2936302-Figure13-1.png", "caption": "Fig. 13. Final steady-state of one arm of the robot when the displacement is too large.", "texts": [ "01 N), and tr,d is the actual stabilization time of the robot parking. When tr,d = th,d and \u0393t = \u03b2, the robot astronaut can be considered an ideal humanoid robot in terms of the stabilization time; when tr,d > th,d, \u0393t increases exponentially, which indicates that the efficiency of parking decreases. In robot astronaut parking, the maximum displacement of the body is limited by the mechanical structure and arm posture. If the theoretical displacement of the robot body during parking is large, then the steady state of the robot might be in the position shown in Fig. 13. The mechanical structure of the various parts of the robot arm may cause interference. When d \u2192 0, the robot arm will reach a singular pose, which causes the robot body to collide directly with the handrail, resulting in a failure to park. A piecewise function based on the actual configuration of the robot is introduced to evaluate the displacement characteristics of the robot, as follows: \u0393s = { \u03b3 sr sh sr < smax +\u221e sr \u2265 smax (27) where \u03b3 is the weight of the displacement in the evaluation, sr is the robot stable state parking displacement, sh is the human stable state parking displacement, and smax is the maximum theoretical displacement determined by the configurations of the robot (in this article, smax = 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001122_tie.2020.2965484-Figure13-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001122_tie.2020.2965484-Figure13-1.png", "caption": "Fig. 13. Case temperature at rated-load. (a) Actual view. (b) Thermal imager.", "texts": [ " The one-way utilization of Poisson\u2019s equation without coupled paths in [26] is inaccurate to consider the tooth as an isothermal surface, especially when the tooth owns an unignorable gradient. Fig. 12 shows the percentage of differences between the results from ETN and FEA. A good agreement is achieved in the internal region with the maximum error is below 3%. The main difference lies in slot boundary which is caused by limited meshes in the tooth region. With more meshes in tooth, the transition of temperature gradient on the boundary will be smoother, which will be closer to FEA and decrease the percentage of differences. Fig. 13(a) shows the actual view of the prototype. Fig. 13(b) shows the measured temperature by using FLUKE at the rated-load with ambient 23\u00baC. Table III compares the ETN, FEA and measured results. It can be seen that the temperature in slot can be accurately predicted with the other parts like case and yoke. The measured results are in good agreement with the ETN one, verifying the accuracy in steady-state. Slight discrepancies in ETN are within the allowable ranges compared to FEA. To estimate the dynamic temperature rise process of SPM machine, the steady-state (5) in ETN model is related to thermal capacity matrix C and replaced by: 1( ) d dt T C P GT (28) Besides, the instant copper loss Pcopper is sensitive and dependent on the temperature, which affect greatly on copper loss and increase the overall temperature" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000976_s11044-019-09704-1-Figure17-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000976_s11044-019-09704-1-Figure17-1.png", "caption": "Fig. 17 Radially loaded ball", "texts": [ " (30) For static equilibrium, the applied radial load must equal the sum of the vertical components of the ball load [43]: Fr = Qmax \u03c6=\u00b1\u03c6l\u2211 \u03c6=0 [ 1 \u2212 (1 \u2212 cos\u03c6)/2\u03b5 ]1.5 cos\u03c6. (31) Equation (31) can also be written as [42] Fr = ZQmaxJr(\u03b5). (32) The radial integral Jr(\u03b5) of Eq. (31) has been evaluated numerically for various values of \u03b5 shown in Fig. 16. Corresponding to zero clearance, \u03b5 = 0.5. For one pair point contact of angular contact ball bearings having zero clearance and subjected to a simple radial load, the maximal load of the ball was determined by Stribeck [43] (see Fig. 17(a)) to be Q = Fr cos\u03b1 , (33a) Qmax = 4.37Fr Z cos\u03b1 . (33b) There\u2019re four-point contact interactions of ball-to-ring raceway surface contacts when the thin-walled four-point contact ball bearing has zero clearance and is subjected to a simple radial load shown in Fig. 7. The four-point contact interactions consist of two pairs of the crossing point contact interactions shown in Fig. 17(b). Similarly, for static equilibrium, the two components of ball load are equivalent to each other. The applied radial load must equal the sum of the vertical components of the ball load. Fig. 18 Load distribution of thin-walled four-point contact ball bearing The maximal load of the ball at the \u03c6 = 0 azimuth can be expressed as Q = Fr 2 cos\u03b1 , (34a) Qmax = 4.37Fr 2Z cos\u03b1 . (34b) From Eq. (30), \u03b5 = 0.5, the load of ball i at the \u03c6 azimuth is expressed as Q\u03c6i = Qmax cos1.5 \u03c6i. (35) The simulation tests were carried out by using ADAMS", " The maximal contact force and the angle of load distribution of the main contact pair are both larger than those of the secondary contact pair. The frequent impact of cage-to-outer ring guidance surface contacts is presented in Fig. 27. Compared to Fig. 22, the impact of cageto-outer ring guidance surface contacts under the conditions of radial load and axial preload combination is less frequent than that of pure radial load in the non-load zone. The angular velocity and motion trajectory of cage\u2019s center are presented in Figs. 28 and 29. Compared to Fig. 17, the angular velocity of the cage\u2019s center is much more stable and much less slight fluctuating than that of pure radial load. The motion trajectory of the cage\u2019s center is similar to a circular motion trajectory. The cage has whirling motion and is slightly fluctuating. The angular velocities of the balls are presented in Fig. 30. The rules of angular velocities of the balls in the cage small and big pockets are very similar to each other and varying in load and non-load zone, respectively. Their values reduce not too much in the non-load zone" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001885_tvt.2020.2986395-Figure18-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001885_tvt.2020.2986395-Figure18-1.png", "caption": "Fig. 18. Prototype of MSCP PM machine. (a) CP PM rotor. (b) MS with the windings. (c) Assembled MS and CP PM rotor.", "texts": [ " The maximum efficiency of the MSCP-AW1 PM machine is the smallest due to its long end winding, in addition, its flux-weakening capability is the worst due to its open circuit back-EMF is the largest. In order to verify the eliminate principle of even-order back-EMF harmonics and the torque characteristics in the MSCP PM machines. The MSCP and MSCP-AW2 PM prototype machine are manufactured, the AW2 is realized by reconnecting the windings. The CP PM rotor, stator with concentrated armature windings and auxiliary distributed windings are given in Figs. 18(a) and (b), respectively. The assembled machine is shown in Fig. 18(c). It should be noted that the stator outer diameter and teeth width of the prototype machine are 105mm and 6mm, respectively. The measured cogging torque is obtained from the reading of electronic weight and the arm length [17]. It is compared with the FE predicted ones, as shown in Fig. 19, the \u201cFE\u201d and \u201cMEA\u201d denote the FE predicted and measured results, respectively. The measured magnitude is higher than the FE predicted owning to the machining and measure errors. While the period of measured and the FE predicted coincides well" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001844_s12541-020-00333-9-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001844_s12541-020-00333-9-Figure3-1.png", "caption": "Fig. 3 Definition of variable parameters for the flexspline tooth profile", "texts": [ " The (6) \u23a7 \u23aa\u23aa\u23a8\u23aa\u23aa\u23a9 x1 = b cos( ) + al cos( )\u221a (a sin )2 + (b cos )2 y1 = a sin( ) + bl sin( )\u221a (a sin )2 + (b cos )2 (7)\ufffd\ufffd\u20d7\ud835\udf14 = \ufffd\u20d7L \u2212 \ufffd\ufffd\ufffd\u20d7rm (8) \u23a7\u23aa\u23aa\u23a8\u23aa\u23aa\u23a9 v = \u2212 \u222b 0 ( )d = \u2212 0 2 sin (2 ) = arctan \ufffd \ufffd \ufffd \u2248 1 rm d d = 1 rm \ufffd 2 0 sin(2 ) \ufffd straight line-segment near the index circle is the tangent of two circular arcs. A dynamic coordinate-system (x1, o1, y1) is used with regard to the axial section of the flexspline tooth, where y1 is the symmetrical axis of the tooth and o1 is the intersection point between the neutral curve and the y1 axis. The variable parameters are shown in Fig.\u00a03. According to the definition shown in Fig.\u00a03, the tooth profile can be defined by Table\u00a02. We use the tooth-profile arc length s as an independent variable to describe the double arc profile function as follows: 1. Tooth-flank arc segment (AB)\u2014see Eq.\u00a0(9) where 2. Straight line segment (BC)\u2014see Eq.\u00a0(10) where 3. Tooth-flank arc segment (CD)\u2014see Eq.\u00a0(11) (9) { X1 = R \u22c5 Xu + T Xu = d \u22c5 Ru s \u2208 [ 0, lAB ] = 2 + arctan ( xA \u2212 xo d yA \u2212 yo d ) , = 2 + arctan ( xA \u2212 xo d yA \u2212 yo d ) , T = ( xod , yod , 1 )T , Ru = [ cos ( s d ) , sin ( s d ) , 1 d ]T (10)X1 = ( s \u2212 lAB ) Ru + T s \u2208 [ lAB, lAB + lBC ] T = ( xod , yod , 1 )T , Ru = [ cos s,\u2212 sin s, 1 s \u2212 lAB ]T , T = ( xoa , yoa , 1 )T 1 3 where 2" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000187_s10846-019-01113-y-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000187_s10846-019-01113-y-Figure5-1.png", "caption": "Fig. 5 D-H coordinates", "texts": [ " From another perspective, Block A* can be seen as a special case of Subgoal Graphs, where subgoals were calculated rather than pre-defined. However, in this paper, the environment is so complicated that the real high-dimensional accessible region of searching space is not in regular shape. This causes difficulties in making the the low-resolution map which may not correctly represent the real accessible region. Hence, Lazy Theta* may be the most suitable option. Details will be discussed in Section 6. Forward kinematics is based on the D-H coordinate modelling shown in Fig. 5. The world coordinate coincides with the base coordinate (Coordinate 0) of CMM initially. Both two manipulators share the same kinematic model with respective parameters. Parameters of D-H coordinates are listed in Table 1. \u03b8i stands for each joint angle. A set of joint angles is called a joint configuration. Transfer matrix of Coordinate i relative to Coordinate i-1, denoted as Ai i\u22121, has the form Ai i\u22121 = \u239b \u239c\u239c\u239d cos i \u2212sin icos\u03b1i sin isin\u03b1i aicos i sin i cos icos\u03b1i \u2212cos isin\u03b1i aisin i 0 sin\u03b1i cos\u03b1i di 0 0 0 1 \u239e \u239f\u239f\u23a0 (1) End-effector\u2019s pose (Coordinate VII) relative to Coordinate 0, denoted as Trelat ive, follows Trelat ive = \u239b \u239c\u239c\u239d nx ox ax px ny oy ay py nz oz az pz 0 0 0 1 \u239e \u239f\u239f\u23a0 = 7\u220f i=1 Ai i\u22121 (2) The structure of 7DOFs makes inverse-kinematics challenging because one single pose of end-effector corresponds to infinite configurations" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0003154_s0368393100103104-Figure18-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0003154_s0368393100103104-Figure18-1.png", "caption": "Fig. 18. Determination of hx-distribution in a single cell tube.", "texts": [ " j Now using these relations, 0 ) is ctually the amount of contraction of the membrane ABCD along ector w owing to the wrinkling deformation. If the wrinkling train is defined as lim A B \u2032 \u2192 0 B B \u2032 A B \u2032 = \u03b2 1+ \u03b2 for the case of the uniaxial de- ormation, it is clear that the wrinkling strain reflects the proporion of contraction of an infinitesimal membrane compared to its riginal length ( A B \u2032 ). The more severely a membrane is wrinkled, he larger its contraction will be. This means that the wrinkling train can be adopted as a measurement of the wrinkling level. Fig. 4 illustrates the motion of an infinitesimal element AB f a wrinkled membrane. The element AB moves to the position \u2019B\u2019 after the occurrence of out-of-plane deformation. w is the ut-of-plane displacement of Point A and w + d w is that of Point . Given that there is no compressive strain when a membrane s wrinkled, i.e. AB is rigid in the motion, AB = A\u2019B\u2019 = d x. \u03bb is the ariation of AB\u2019 s projection on the x- axis after wrinkling and \u03bb = \u221a (d x ) 2 \u2212 (d w ) 2 \u2212 d x \u2248 \u22121 2 ( d w d x )2 d x (8) It is assumed in Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000967_j.ijepes.2019.105539-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000967_j.ijepes.2019.105539-Figure1-1.png", "caption": "Fig. 1. A sample 4-pole OR-PMSM.", "texts": [ " At the outset, an analytical model based on MEC method (which is both fast and flexible) has been introduced in order to analyze OR-PMSMs performance and Demag. fault detection. Finally, proposed model is verified by Finite Element Analysis (FEA) and their processing time are compared. Innovations of this research can be summarized as follows: A \u2013 Modeling the OR-PMSMs by a new flexible MEC method. B \u2013 Determining the accuracy level of the model as desired. C \u2013 Demagnetization fault detection for OR-PMSMs by current signature analysis. A sample 4-pole OR-PMSM is shown in Fig. 1, where all of the used dimension are shown in the figure. In order to obtaining of a model with selective accuracy, the machine structure is divided into 11 zones for modeling a p-pole machine as shown in Fig. 2. Rij and ij are reluctance and flux of jth element in ith region, respectively. The model accuracy can by tuned by the number of the considered elements defined by nr, and ns (see nomenclature). The leakage fluxes in the slots and inter-magnets regions are modeled by considered reluctances in the air parts" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001800_978-3-030-40410-9-Figure5.5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001800_978-3-030-40410-9-Figure5.5-1.png", "caption": "Fig. 5.5 SCARA manipulator", "texts": [ "27), where we have seen that the position vector is the first three elements of the fourth column of the transformation matrix while the z-vector is the first three elements of the third column of the homogenous transformation matrix. For a prismatic joint, there is only translational movement, thus, there is no contribution to the orientation of the robot. Consequently, the z-vector of the prismatic joint is taken to be zero [ 0 0 0 ]T , as we will see in Example 5.1. Example 5.1 SCARA Manipulator The SCARA manipulator has four degrees of freedomwith a combination of three revolute joints andoneprismatic joint. Figure 5.5 illustrates the frame assignment of the robot, while the spatial parameters are given in Table 5.1. By substitution of the spatial parameters in Eqs. (4.4) and (4.4a), we can find the homogenous transformation matrix for each link in the manipulator. Equations (5.43)\u2013(5.46) represent the four transformation matrices of the four links. We have to represent the orientation and position of each link to the reference frame; this what is happened in Eqs. (5.47) and (5.48), where we have calculated H 0 2 , H 0 3 , and H 0 4 , respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001038_0142331219879338-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001038_0142331219879338-Figure6-1.png", "caption": "Figure 6. The vertical rotational vector-based potential field.", "texts": [ " The virtual control input are still designed in two cases: for upward direction fxrv = ko v1v2 v3 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi v1 2 + v2 2 p z zo\u00f0 \u00de cos fn\u00f0 \u00de fyrv = ko v1v2 v3 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi v1 2 + v2 2 p z zo\u00f0 \u00de sin fn\u00f0 \u00de fzrv = ko v3 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi v1 2 + v2 2 p v1v2 x xo\u00f0 \u00de cos fn\u00f0 \u00de v3 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi v1 2 + v2 2 p v1v2 y yo\u00f0 \u00de sin fn\u00f0 \u00de \u00f033\u00de or in downward direction fxrv = ko v1v2 v3 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi v1 2 + v2 2 p z zo\u00f0 \u00de cos fn\u00f0 \u00de fyrv = ko v1v2 v3 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi v1 2 + v2 2 p z zo\u00f0 \u00de sin fn\u00f0 \u00de fzrv = ko v3 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi v1 2 + v2 2 p v1v2 x xo\u00f0 \u00de cos fn\u00f0 \u00de + v3 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi v1 2 + v2 2 p v1v2 y yo\u00f0 \u00de sin fn\u00f0 \u00de \u00f034\u00de The vertical rotational vector field is shown in Figure 6. qn is the angle between the line linking the UAV and the obstacle, and the horizontal plane. gn is the angle between the velocity of the UAV and the horizontal plane. zn is the angle between the tangent direction of vector field and the horizontal plane. Similarly, the angles can be calculated as follows gn =arctan _z, ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi _x2 + _y2 q zn =arctan\u00bd v1v2 v3 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi v2 1 + v2 2 p z zo\u00f0 \u00de, v3 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi v2 1 + v2 2 p v1v2 x xo\u00f0 \u00de cos fn\u00f0 \u00de v3 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi v2 1 + v2 2 p v1v2 y yo\u00f0 \u00de sin fn\u00f0 \u00de qn =arctan zo z, ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi xo x\u00f0 \u00de2 yo y\u00f0 \u00de2 q \u00f035\u00de Obstacle avoidance algorithm in 3D Space" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001366_j.measen.2020.100004-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001366_j.measen.2020.100004-Figure5-1.png", "caption": "Fig. 5. (a) Schematic representation of the automated testing rig and (b) the actual testing rig setup showing the modifications to a stock 3D printer for enabling automated sensor characterization.", "texts": [ " The serial data received from the Arduino and the load cell is displayed on an in-house LabVIEW GUI program which also programmatically writes the tactel information, capacitance and load values to a data file for further analysis. A G-code moves the x, y, and z-axes of the Prusa i3 printer to synchronously position the load cell-indenter to the tactel whose capacitance the Arduino is currently measuring. A schematic representation showing the automated testing rig setup and the actual testing rig is shown in Fig. 5a and Fig. 5b respectively. Multiple ECT sensors were fabricated in 2 PE configuration (up to 9 sensors per batch of composite) (Fig. 6a) to measure and establish the sensor characteristics and performance metrics. Three 4 PE ECT sensors were also fabricated for a comparative study. The 2 PE ECT sensors fabricated from the process described above, as seen from Fig. 6b, are highly elastic and flexible and the fabrication process shows the ease of encapsulation or embedding of the active sensing layers within a passive layer which can very well be the elastomeric layer of any soft robotic system" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002546_ccnc49032.2021.9369615-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002546_ccnc49032.2021.9369615-Figure2-1.png", "caption": "Fig. 2. Robot node prototype.", "texts": [ " These environments will provide a digital replica of the physical setup and are easily reconfigurable via the web portal provided. These playgrounds can be utilized for the experiment validation, prior to the deployment on the arena robots, or for the evaluation of more complex robotic usecases and configurations. The Gazebo web client interface will be used for visualising the GUI output of the simulator to the users within WebGL compliant browsers (see Fig. 4). 4) Robot nodes The robot nodes (see Fig. 2) have many sensors and radios as well as a lifter actuator for moving pallets either individually or in a collaborative manner. The available sensors consist of: \u2022 16 IR Laser time-of-flight distance \u2022 4 Cameras \u2022 9-DoF IMU \u2022 Temperature, pressure, humidity \u2022 Strain gauges \u2022 Robot health (various voltage, current and temperature readings) The robot node processor is based on the RockPi\u2122 4B with GPU support. This permits machine learning and vision application experiment containers to be run on the robot nodes, that can exploit OpenCL" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001347_j.conengprac.2020.104486-Figure12-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001347_j.conengprac.2020.104486-Figure12-1.png", "caption": "Fig. 12. 3-DoF helicopter experimental prototype.", "texts": [ " The noise in the measurements strongly affects the residual detectors (18a)\u2013(18b) as can be seen in Fig. 8. However, filtering of such residuals is suggested in order to mitigate the consequences of the noise. 6. The computational effort required to implement the FDI-SMD scheme is almost the same as needed by the FDI-SMO technique, which makes the proposed algorithm suitable to be implemented instead of the classic FDI-SMO approach. The 3-DoF helicopter prototype developed by Quanser is shown in Fig. 12. It is instrumented by three quadrature encoders with resolution 4096 pulse/rev for the measurements of elevation \ud835\udf00 and pitch \ud835\udf0c angles, and 8192 pulse/rev for the travel angle \ud835\udf03. Signal processing is computed using a DSpace 1103 controller board composed by a high-speed Digital Signal Processor (DSP) PPC750GX of Texas Instruments developer. The Simulink scheme is exported to the EPROM memory of the board in which the Euler\u2019s integration method, with fixed step \ud835\udf0f = 10\u22124 s, is used to solve the dynamical system of SMD (19)" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001443_rpj-11-2019-0287-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001443_rpj-11-2019-0287-Figure6-1.png", "caption": "Figure 6 Schematic of Vickers microhardness measurement on the samples", "texts": [ " The influences of the support structures removal on the hardness and residual stress of the samples were ignored in this work. The force used in the measurement was 980.7mN for a dwell time of 15 s. In total, 30 points were measured for each sample, and the distance between every two consecutive points was 0.5mm. Samples with and without support structures were measured at the same 30 points, and the value corresponding to each point was obtained. The measured points are marked in the schematic of Figure 6. The average value of Vickers microhardness was calculated by averaging the 30 points on the surface of each sample. Various methods have been devised for measuring residual stress, such as hole-drilling, crack compliance, X-ray diffusion, neutron diffraction, Vickers micro-indentation, etc. Among these measurement techniques, the Vickers micro-indentation method is the easiest and fastest way to measure residual stress because a series of points were previously used to measure the microhardness (Lu et al" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002437_s00366-020-01236-z-Figure14-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002437_s00366-020-01236-z-Figure14-1.png", "caption": "Fig. 14 Details of a substructure of the 600-bar single-layer dome truss", "texts": [ " It should be noted that the proposed method requires 15,560 structural analyses to find the optimum solution while WOA, EWOA, CBO, and ECBO require 18,840, 19,300, 17,700, and 19,700 structural analyses, respectively. Stress ratios and nodal displacements in all directions evaluated for the best design achieved by MWQICBO are shown Fig.\u00a012. The maximum stress ratio and the maximum nodal displacement are 99.95% and 3.1488 in, respectively. The sizing optimization of a 600-bar single-layer dome structure schematized in Fig.\u00a013 is the last test case. The entire structure is composed of 216 nodes and 600 elements. Figure\u00a014 shows a substructure in more detail for nodal numbering. The cross-sectional area of each of the member in this substructure is considered to be an independent variable. Therefore, this is a size optimization problem with 25 variables. Table\u00a010 presents the coordinates of the nodes in the Cartesian coordinate system. The elastic modulus is 200 GPa and the material density is 7850\u00a0kg/m3 for all elements. Non-structural masses of 100\u00a0kg are attached to each free node. The minimum and maximum admissible crosssectional areas are 1 cm2 and 100 cm2, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002316_rcar49640.2020.9303297-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002316_rcar49640.2020.9303297-Figure5-1.png", "caption": "Fig. 5. The strategy for grasping five cucumbers simultaneously in (a) and the gripper design in (b).", "texts": [ " Moreover, the new fabrication method facilitates mass production because of the simplified process, and therefore can lower the cost of the gripper. To package multiple cucumbers simultaneously, we need to minimize the gaps between neighboring cucumbers while packaging to save package space. At the grasping position, the cucumbers are approximately aligned by using a wave conveyor and have enough gaps for the gripper to enter. After grasping, the gripper should be able to shrink the gaps before packaging. To this end, we proposed a grasping strategy as shown in Fig. 5a. Each cucumber was grasped by two gripping units and the gripping positions were shifted between neighboring cucumbers. Therefore, while packaging, the gap between neighboring cucumbers can be shrank to one thickness of the hybrid finger, which is 3.0mm. Based on this strategy, we designed a shell gripper as shown in Fig. 5b. A parallel link mechanism was used to shrink the gaps and a pneumatic cylinder was used to drive the mechanism. To obtain the relationship of pressure and inflation, we performed finger inflation tests on hybrid fingers fabricated using both conventional and new methods. The experimental setup is shown in Fig. 6a. A laser distance sensor (HGC1050-P, Panasonic, Japan) was used to measure the displacement of the finger. A snapshot showing the inflated finger at a pressure of 50kPa is shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002233_10426914.2020.1832686-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002233_10426914.2020.1832686-Figure1-1.png", "caption": "Figure 1. (a) Schematic of LSM experiments; (b) representative sample upon LSM.", "texts": [ " In some studies, applicability of LSM is checked/shown toward a given implant (combination of surface characteristics) at very specific combination of the process parameters (Dental implants,[29] Orthopedic implants,[30] Hip implants,[31] etc.). However, these studies do not focus on how combination of the surface characteristics changes with variation in process parameters. The objective of this work is to analyze the combined effect of LSM process parameters on the overall surface characteristics (combination of surface roughness, microhardness, and wettability of Ti6Al4V) for various biomedical implants. Figure 1(a) shows the schematic of the LSM process. The material investigated in the study is commercially available Ti6Al4V alloy. ASPL continuous wattage fiber laser of 100 W (nominal output, 1060 nm wavelength) was used to create surface textures of size 10 mm \u00d7 10 mm, on a rectangular plate with overall dimensions of 215 \u00d7 160 \u00d7 7 mm3. The beam characteristics were Gaussian and a minimum beam diameter of 30 \u03bcm could be achieved.The LSM was performed with constant power of 100 W and beam diameter (BD), scanning speed (SS), and beam overlap (BO) were varied across the experiments. The experiments were performed with combinations of three BDs (0.5 mm, 1.1 mm and 1.4 mm), four SSs (60 mm min\u22121, 200 mm min\u22121, 400 mm min\u22121, and 600 mm min\u22121), and three BOs (30%, 50%, and 70% of BD), a total of 36 samples. The levels of process parameters were chosen to generate data over large range of process parameters. Figure 1(b) shows the representative Ti6Al4V sample upon LSM. The laser modified samples were cut (perpendicular to the raster-scan direction) to reveal the cross-section. The cross-sections were polished and etched using Kroll\u2019s reagent. The microstructure of modified samples was characterized by scanning electron microscopy (SEM). Vickers hardness and surface profile of the laser modified samples were determined using HMV microhardness tester and Zeta roughness measurement instrument (OLYMPUS U-D6RE-EDS-2), respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001150_indin41052.2019.8971967-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001150_indin41052.2019.8971967-Figure6-1.png", "caption": "Fig. 6. The final product, i.e., assembled box, of the use case task", "texts": [ " This use case is structured in 4 subsections that describe the main steps of the process to create the VR simulation. The workflow of this process is presented in Fig. 4 : The use case consists on a VR simulation of an assembly task with YuMi, which is the ABB commercial cobot. As a cobot, it will be used in a process that requires a robot to work together with a human in a secure way. The task performed is the assembly of a box made of wood. Further, the scenario includes a table which contains all the parts for assembling the box, i.e., box sides, bolts and screws. Fig. 5 and Fig. 6 illustrate the elements involved in the task, corresponding to YuMi robot and the assembled box respectively. In the proposed scenario, YuMi\u2019s work is to pick up the different parts of the box and assembling them in a proper position for the user\u2019s activity. The human operator must link the parts that the robot is holding with bolts and screws. This is an iterative process that ends when the box is completed. As the goal of the training is getting the user used to working with the cobot, the simulation must be as similar as possible to the reality" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001087_ab62c8-Figure13-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001087_ab62c8-Figure13-1.png", "caption": "Figure 13. Case1: (a) the experiment rig and (b) the rolling bearing with an outer race fault.", "texts": [ " In section\u00a0 3, we validated the efficiency of the proposed method for extracting periodic transient signal. In this section, we adopt the proposed method to analyze two vibration signals measured from a bearing with an implanted fault and a run-to-failure bearing to diagnose the bearing faults. To further verify the superiority, the proposed method is compared with two methods namely the L1-norm denoising method and the SK method. 4.1.1. Data acquisition. The bearing vibration signal is collected from an engine rotor-bearing-casing fault simulator [43], as described in figure\u00a013(a). The experimental system is composed of a rotor-bearing-casing system, a miniature accelerometer of type 4508 from Br\u00fcel & Kj\u00e6r Sound & Vibration Measurement A/S, and a data acquisition module of type NI9234. Figure\u00a0 13(b) presents the bearing with implanted outer race fault, which is installed at the end of the shaft. The vibration signal is acquired by an accelerometer installed on the top of the bearing block. Table\u00a03 lists the parameters of the test bearing to be diagnosed. The data is sampled at 10 kHz, and the length of the signal to be analyzed is 32768 points. The bearing fault characteristic frequencies are described in table\u00a04. 4.1.2. Results and analysis. Figure 14 shows the time domain and Fourier spectrum of the collected bearing vibration signal" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000173_01691864.2019.1680316-Figure8-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000173_01691864.2019.1680316-Figure8-1.png", "caption": "Figure 8. Stairs and basic posture of the robot. (a) Stairs. (b) Basic posture of the robot when climbing stairs.", "texts": [ " In contrast, the region of the joint part in the entire body of the robot developed in the paper is larger than that of [15] because the developed robot has both a pitch joint and yaw joint at the joint part. Thismeans that it is easy for the joint part of the robot to contact the surroundings. If the robot uses a backward traveling wave without considering the relative relationship between itself and the stairs, it is possible for the robot to become stuck or fall down. Thus, we designed a novel stair climbing motion for the articulated mobile robot so that it can avoid becoming stuck or falling. Figure 8 shows the stairs assumed in this section. The riser heights and tread depths are all the same. We define that depth d satisfies the condition r \u2264 d < L + r, (1) where L is the module length of the robot and r is the wheel radius, as shown in Figure 3 and Table 2. The robot cannot climb stairs whose parameters satisfy condition (1) using the method of [16]. The arm is treated as a virtual module similar to the case of basic steering control considering the folding arm. We assume that the robot approaches the stairs parallel to the x axis, as shown in Figure 8. The posture shown in Figure 8 is used as the basic posture of the robot when climbing stairs. One pair of wheels contacts one tread. The part of the body connecting two treads is called the connecting part. The posture of the connecting part is designed so that the distance between the contact points between the robot and tread is d along the x axis.We assume that the connecting part consists of two modules and the angle of the yaw joint on the connecting part is zero, as in [16]. Then, d, h, and L have to satisfy the following condition: d2 + h2 \u2264 4L2" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001669_gcwkshps50303.2020.9367568-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001669_gcwkshps50303.2020.9367568-Figure2-1.png", "caption": "Fig 2. MmWave Flat-Top antenna model", "texts": [ " Most of the links at air communications are line of sight (LOS) ones. Hence, the received power at UAV n from UAV m is calculated from the Friis transmission equation as [12] where is the UAV m transmitted power, is the separation distance between UAVs m and n, and are the wavelength and the path loss exponent, respectively. is the transmitter (TX) mmWave beamforming gain while is the same but for the receiver (RX). is the beam offset position of the TX beam route towards the RX position, while is the same but from RX beam to TX position, as shown in fig.2. reflects the -3dB beamwidth value. We assume a flat-top antenna model, in which can be formulated as [13]: ! ! \" #$ % % & \" '()*+,#-* where \" refers to the gain of sidelobe, . & \" & . Though, other mmWave BF techniques can be utilized by the proposed scheme. and are tuned during BT by directed antenna arrays of UAVs TX and RX. Assume that i access UAVs and j gateway ones are spread within the post-disaster area coverage. Every round & &/, every access UAV m choose gateway UAV n to fly to it and relay its gathered data, where T is the horizon that equals the battery lifetime of access UAV before recharging needs" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000536_j.mechmachtheory.2016.08.008-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000536_j.mechmachtheory.2016.08.008-Figure3-1.png", "caption": "Fig. 3. The spherical four-bar linkage and local Cartesian reference frames definition according to Denavit convention.", "texts": [ " For this reason, these equations need not be included here. 13. Numerical examples 13.1. Kinematic analysis of the spherical four-bar linkage The previous theory can be applied to extend the Denavit method of mechanism position to velocity and acceleration analysis. For ready comparison, the numerical example reported herein refers to the same spherical four-bar linkage analyzed by Denavit [3]. In particular, the angular dimensions of the links follow: \u03b1 = \u00b0601 , \u03b1 = \u00b0302 , \u03b1 = \u00b0553 , \u03b1 = \u00b0454 . The crank, link 2 of Fig. 3, has an absolute constant angular speed of \u03c9 = 1 rad/s2 . For position analysis, although not strictly necessary, the set (65) of nonlinear equations was solved by means of an iterative method for \u03b81, \u03b83 and \u03b84. Velocity and acceleration values are obtained upon solving the set of redundant linear equations (75) and (78). In fact, from Eq. (75) four linear equations in the unknown angular velocities \u03b81\u0307, \u03b83\u0307 and \u03b84\u0307 are obtained. Similarly, the unknown angular accelerations \u03b8\u03081, \u03b8\u03083 and \u03b8\u03084 follow from Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000380_s1758825116500393-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000380_s1758825116500393-Figure1-1.png", "caption": "Fig. 1. Schematic of core/shell hydrogels at three different states shows core and shell domains. Core and shell occupy distinct domains in the initial state. The reference state is differentiated from the initial dry state. The equilibrium state coincides with the actual state of gels.", "texts": [ " (4) The molar fraction of H+ and OH\u2212 ions are found from the concentrations of ions in the bath as [H+] = k c\u0304H+ c\u0304 , [OH\u2212] = k c\u0304OH\u2212 c\u0304 , [Cl\u2212] = k c\u0304Cl\u2212 c\u0304 , (5) where c\u0304, c\u0304H+ , c\u0304OH\u2212 , and c\u0304Cl\u2212 represent the concentration of water molecules, hydroxide ions, hydroxyl radicals and chloride ions in the bath, respectively, and k = 55.5\u0304 is the molarity of water. As a common assumption, the infinite amount of solvent molecules and solutes are available in the bath which maintains the concentration of mobile ions constant. The ionization of water molecules and hydrochloric acid and also, association of H+ mobile ions with water molecules in Eq. (3) are disregarded in this model. 2.2. Kinematics of a hydrogel The macro-deformation of a hydrogel has been described with the assumption of three individual states as shown in Fig. 1. The dry undeformed state of a hydrogel 1650039-5 In t. J. A pp l. M ec ha ni cs 2 01 6. 08 . D ow nl oa de d fr om w w w .w or ld sc ie nt if ic .c om by W E IZ M A N N I N ST IT U T E O F SC IE N C E o n 07 /0 4/ 16 . F or p er so na l u se o nl y. 2nd Reading May 10, 2016 13:6 WSPC-255-IJAM S1758-8251 1650039 N. Hamzavi et al. coincides with the initial state, and the as-prepared state presumed to be the reference state. The initial state of a hydrogel is differentiated from its reference state" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000427_0954410016643978-Figure17-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000427_0954410016643978-Figure17-1.png", "caption": "Figure 17. Stress extracted area.", "texts": [ " The natural frequencies and the mode shapes of the structure are measured by the hammering method, and there are 240 points for measuring the mode shapes. One acceleration sensor is placed in the structure. LMS SCADAS III type vibration measurement and analysis system is used for the data processing as shown in Figure 16. Based on the material properties and the structure size, the sub-section of the bolted joints is modeled by the 3D solid FE model. Because of the same bole preload, the mean contact stress of the flange interface shown in Figure 17 is obtained as the overall flange interface contact stress. For the material of the partitioned thin-layer elements in different areas, each needs three independent parameters E, Gy, Gz to simulate three stiffness respectively, and the material parameters in each area are different from those of other areas, thus there are 3N\u00fe 3 (N is the bolt number) material parameters to model the partitioned thin-layer elements of the bolted joints. Therefore, the orthogonal anisotropic material is served as the material of the partitioned thin-layer elements" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000718_978-1-4471-7278-9-Figure6.1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000718_978-1-4471-7278-9-Figure6.1-1.png", "caption": "Fig. 6.1 Various door opening challenges for a mobile robot from left to right: door knob, door handle, door press latch, card reader, and combination keypad", "texts": [ " Because of their mobility, smartphones have more potential to be adapted to new situations than desktop computers. A more comprehensive case study is how to enable an autonomous mobile robot to adapt itself to challenging situations, such as opening a door in a hospital, without having to reconfigure itself. Assume a mobile robot has a basic architecture \u2013 an arm and gripper that can reach the height of any door accessory, e.g. doorknob, door handle, door press latch, card reader, or combination keypad. See Fig. 6.1. For turning a doorknob, it is possible for the robot to use its gripper to rotate the knob with enough friction. A sliding sensor would be needed to detect sliding in order to apply sufficient pressure on the gripper. In addition to rotating the door handle, the gripper must exert some force in the downward direction. The gripper may push the door press latch forward in order to open the door. When additional force is needed, the robot may back up and use momentum to add thrust. The gripper may also slide a card through the slot in a vertical card reader" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002700_s12541-021-00510-4-Figure31-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002700_s12541-021-00510-4-Figure31-1.png", "caption": "Fig. 31 Schematic diagram of an involute internal gear pump", "texts": [ " Involute internal gear pairs have a characteristic of separability, implying that the center distance between the gear and the ring gear changes within a certain range and the transmission ratio remains constant [13]. In contrast, meshing points of a conjugated straight-line internal gear pair correspond to each other. If the center distance changes, the gear pair fails to mesh effectively. This characteristic leads to a different radial structure design. The involute internal gear pump shown in Fig.\u00a031 can be designed as a floating crescent plate by introducing a compensation structure and dynamically adjusting the center distance. However, the conjugated straightline internal gear pump shown in Fig.\u00a032 can only use fixed crescent plates. This fixed clearance design cannot dynamically adjust the center distance to compensate for the wear of friction pairs. Therefore, this design is more complicated, and the requirement for materials is high. 2. The involute internal meshing gear pair shown in Fig.\u00a031 meshes as a straight line, and the meshing characteristic changes linearly with the gear rotation angle. In addition, the flow characteristic changes with the gear rotation angle in a quadratic relationship. On the contrary, the meshing line of the conjugated straight-line internal gear pair shown in Fig.\u00a032 is a curve. The meshing curve changes nonlinearly with the gear rotation angle. Thus, the trapped oil volume gradually increases dur- 1 3 ing transmission, and the formation of negative pressure between gear pairs is more conducive to stable transmission (Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000718_978-1-4471-7278-9-Figure16.1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000718_978-1-4471-7278-9-Figure16.1-1.png", "caption": "Fig. 16.1 Wall-following robot", "texts": [ " We can simulate the wall-following function with a simple mobile robot. Assume we have two continuous rotation motors to drive two active wheels with one passive wheel behind. The speed difference between the two wheels enables the robot to move left or right. We can have two infrared distance sensors on the one side of the robot to monitor the distance to the wall and keep the robot moving parallel to the wall. In fact, we only need one distance sensor that is at 45\u0131 to the centerline of the robot, similar to using one hand to probe at the wall in dark (Fig. 16.1). The control logic is simple: first, the robot moves around to find the wall. When the sensor detects that the distance to the wall is too large, the robot keep moving toward the wall. The robot keeps going forward while maintaining a fixed distance to the wall with a certain tolerance. Instinctive Computing 321 The wall-following algorithm is just a feedback control system to ensure the distance to the wall is constant. The same principle can be used in different applications. For example, line following, shadow-following, or turning the distance sensor into a smell sensor to follow the trace of a smell" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001138_s42235-020-0011-x-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001138_s42235-020-0011-x-Figure3-1.png", "caption": "Fig. 3 Cycloidal trajectory diagram.", "texts": [ " Hence, the system\u2019s output is described as: mea m mea ( ) ( ) ( ) ,\u02c6 ( ) x x q k k x k p k y C (15) where mea ( )xq k and mea ( )xp k denote the measured CoM position and ZMP position, respectively, and m 1 0 0 . 0 0 1 C (16) Taking sensor measurement into consideration, the discrete time system with an observer can be modified as: 1 0 m\u02c6 \u02c6 \u02c6 \u02c6 ,k k k k kx x y x u A L C b (17) where L is an observer gain, which can be obtained by the discrete LQ regulator method. Omnidirectional walking can be easily realized based on this preview control scheme. Fig. 3 shows the method of generating a swing foot trajectory. pswingx and pswingy are the positions of the swing feet in x and z direction, respectively. Let ps = [xs, ys, zs] T and pe = [xe, ye, ze] T be the initial and landing positions of the swing foot, respectively. A cycloid curve is adopted for the swing foot trajectory. Given a desired walking speed vc = [vcx, vcy] T and turning speed , the position on the following time instant can be given as: c c s, 2 , f x x f p v v y (18) where \u03b1 = T/100 , ys is the distance between the two feet along the y-axis when the robot stands still, and T represents the locomotion period in seconds" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001951_j.isatra.2020.05.046-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001951_j.isatra.2020.05.046-Figure1-1.png", "caption": "Fig. 1. The schematic of truck-trailer for forward motion.", "texts": [ " The limitation of the separated kinematic and dynamic controls is that the velocity of the truck-trailer should be set low enough, i.e., the frequency of kinematic system should be lower than the dynamic system, in such a way that the dynamic controls, which actuate traction and steering motors, can track the velocity and steering angle designed from the kinematic system. This section derives the kinematic equations for forward and backward directions of the truck-trailer system. The velocity of the truck-trailer on vertical direction with respect to the surface is considered to be zero for all time. Fig. 1 shows the schematic of a truck-trailer for forward motion at a global coordinate system XOY. The lengths of the axle-toaxle of the head-truck and of the trailer are lh and lt, respectively, and the free-rotating-joint conducting the head and the trailer is located at the point Oh (xh, yh). The head-truck comprises of two-steering front wheels and two-fixed rear wheels while the trailer has two fixed wheels at the rear-side. The center point of the steering axle is denoted by O\u03b4 (x\u03b4, y\u03b4) and that of the trailer\u2019s rear-wheel axle is Ot (xt, yt)" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002322_tmech.2020.3047476-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002322_tmech.2020.3047476-Figure1-1.png", "caption": "Fig. 1. Human arm and anthropomorphic arm, both of which have the S-R-S structure. (a) Human arm; (b) Kinematic model of anthropomorphic arm; (c) Anthropomorphic arm and its self-motion.", "texts": [ " Thus, the human-like motion planning for an arm with multiple constraints plays an important role in the operation of humanoid robots. In the past decades, an anthropomorphic arm that has a structure similar to that of a human arm has been developed to enhance the dexterity and versatility of the robotic arm. An anthropomorphic arm is a special type of manipulator that has a spherical-revolute-spherical (S-R-S) structure. The self-motion of anthropomorphic arm is defined as the rotation motion of the elbow joint about the shoulder-wrist line while the pose of end-effector remain fixed (see Fig.1). However, the most promising arms seem to be those with human-like characteristic not only in their structural types but also in their motions. Some studies have been performed to generate human-like motion of the anthropomorphic arm by imitating human arm motions. A usual assumption of these approaches is that humans attempt to minimize an unknown cost function or criterion while doing everyday manipulation tasks, e.g., joint jerk [1], joint torque [2], manipulability ellipsoid [3], rapid upper limb assessment (RULA) [4], mechanical work [5], [6] or a convex combination of several cost functions with weighting factors [7]", " At a particular end-effector pose, the anthropomorphic arm yields a self-motion around the axis connecting the shoulder and wrist. The self-motion is subject to joint limits and collisions. The arm angle parameter is introduced to represent the self-motion manifold. 1) Joint limits The self-motion manifolds under joint limits are obtained by directly solving the inequality of the joint limits and the arm angle parameter. The kinematic equations of joint 2 and joint 6 are generally represented by the cosine function (see Fig.1) [39] cos sin cos ( 2,6)i i ia b c i = + + = (1) where ,a b and c are the corresponding factors of the kinematic equations respectively. Thus, the self-motion manifolds satisfying the two joint limits in the region [ , ] \u2212 are calculated by using the following: Authorized licensed use limited to: University of Prince Edward Island. Downloaded on June 06,2021 at 14:08:23 UTC from IEEE Xplore. Restrictions apply. 1083-4435 (c) 2020 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission", " Validity checking of Cartesian sampling points based on self-motion manifolds A novel validity checking method of Cartesian sampling points based on self-motion manifolds is presented in this section. The operation sequences are explained as follows: Step 1. Manipulation workspace analysis: Calculate the angle 4 of the elbow joint as follows: 0 2 2 4 arccos 2 sw se ew se ew x d d d d \u2212 \u2212 = (9) Where 0 swx is the vector from the shoulder to the wrist\uff0c ,se ewd d are the lengths of the upper arm and the lower arm, respectively (see Fig.1). If 4 exceeds its joint limit or 0 2 2 1 2 sw se ew se ew x d d d d \u2212 \u2212 , the sampling point exceeds the workspace of the anthropomorphic arm, thus, it should be invalid. Step 2. Validity checking of sampling points: Let SampleFree( n , free ) be a function that returns a set of n N points sampled independently and identically from the uniform distribution on free . This guarantees that the end-effector path of the anthropomorphic arm is collision-free and satisfies the task constraints, but it cannot guarantee that the corresponding arm configurations satisfy the joint limits and collision avoidance criteria" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002120_j.ijfatigue.2020.105894-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002120_j.ijfatigue.2020.105894-Figure5-1.png", "caption": "Fig. 5. Schematic diagrams showing the shapes and dimensions of 3D ellipsoidal/triangular pores (a) x = y and z varies with different length scales (0.1 \u2013 1.5 mm), (b) y = z and x varies with different length scales (0.1 \u2013 1.5 mm), (c) x = z and y varies with different length scales (0.1 \u2013 1.5 mm), and (d) triangular pore with two dimensions fixed at 0.2 (x in this case) and 0.1 mm (y in this case), respectively, and another dimension vaies with different length scales (0.15\u20132.9 mm), where the x or y axis may vary depending on which plane and orientation the triangular pore is resided.", "texts": [ " The triangular pores were denoted in the form of x-y-z as well, but the x, y and z do not follow the axes of the above semi-ellipsoidal pores, while it depends on the side (either short side or long side) and orientation (perpendicular or parallel) being resided in the model as illustrated in Fig. 4b. But all the triangular pores are denoted literally as x fixed at 0.2 mm and y fixed at 0.1 mm (with 10 \u00b5m root radius for the top triangular corner) and varying length scale z from 0.05 mm to 2.9 mm (through pore along short side) or 5.8 (through pore along long side). The 3D morphologies of the semi-ellipsoidal/spheroidal and triangular pores were shown in Fig. 5. The effect of the aspect ratios on stress concentration was studied from 1 to 50 for the semi-ellipsoid/spheroid pores, which was calculated by major axis over minor axis. For the ratio of z/x or z/y, z was fixed at 500 \u00b5m and x&y were varied from 10\u2013500 \u00b5m. For x/y or x/z, x was fixed at 500 \u00b5m and y&z were varied from 10 to 50 \u00b5m. The aspect ratios of the triangular pores were ranged from 0.25 to 14.5, which were calculated by length z divided by x or y, where z varied from 0.05 mm to 2.9 mm and x&y were fixed at 200 \u00b5m" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002210_j.robot.2020.103670-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002210_j.robot.2020.103670-Figure2-1.png", "caption": "Fig. 2. Drive unit for the wheelchair.", "texts": [ " Section 4 hows the results of our theoretical analysis, while Section 5 decribes our experiment and its results. Finally, Section 6 contains ur conclusions. 2. Mechanical architecture of cooperating wheelchair and robot In this research, we used a wheeled robot that was developed in our laboratory (Fig. 1). Table A.1 in Appendix A lists the specifications of the robot. The NOVA Integral-ME wheelchair (OX Engineering, Chiba, Japan) used in this study is a typical commercially available rear-wheel drive model equipped with an electric drive unit (TRD-1, Acritech Co., Ltd., Kanagawa, Japan, Fig. 2), that provides motive power to both rear wheels. Table A.2 in Appendix A provides the specifications of this wheelchair. As will be described below, a number of sensors were added to the wheelchair. In operation, the robot\u2019s hand grasps the back p t c w f d art (rotary shaft of the push handle) of the wheelchair [27] and he robot and the wheelchair are deployed in a forward-and-aft onfiguration in order to maneuver over a step (Fig. 1). This robot is equipped with three pairs of wheels, each of hich consists of a left wheel and a right wheel (Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000113_j.finel.2019.103319-Figure16-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000113_j.finel.2019.103319-Figure16-1.png", "caption": "Fig. 16. Pre-twisted beam test (L\u00bc 12.0; b\u00bc 1.1; E\u00bc 29E6; \u03bd\u00bc 0.22; F1\u00bcF2\u00bc0.5).", "texts": [ " In this configuration, bending strains were predominant than membrane effects. As presented in Fig. 15, results from QUAD\u00fe, Shell63 and lane condition (thickness\u00bc 0.05). S4 are satisfactory even with a very coarse QUAD mesh (\u201cQUAD 6 6\u201d). On the contrary, TRIA\u00fe andMITC3 elements appeared to be less efficient and a mesh finer than 12 12 2 was required to get good results. The pre-twisted beam was obtained by sweeping a line along the beam axis while twisting it at a constant rate from 0 to 90 (see Fig. 16). The beam was double-curved and the mesh was strongly affected by the warping effect, and thus, the bending-membrane coupling came into sight. Both in-plane (F1 load) and out-of plane (F2 load) load conditions were analysed. Moreover, in order to highlight the sensitivity of the thickness, both thin (thickness\u00bc 0.05) and thick (thickness\u00bc 0.32) shells were analysed. Displacements were calculated at load point in direction of the force. QUAD elements were increased from 6 to 48 along the longitudinal axis and from 1 to 8 along the transversal axis" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001360_s42417-020-00220-7-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001360_s42417-020-00220-7-Figure4-1.png", "caption": "Fig. 4 Slot across the bearing outer race", "texts": [ " The outer race of the bearing and the gear wheel is a single component, with the bearing rollers running directly on the inner circumference of the gear. Each planet gear is \u2018self-aligning\u2019 by the use of spherical inner and outer races and barrel-shaped bearing rollers (see Fig.\u00a03). The test procedure consists of three experimental conditions, including fault-free condition, minor seeded bearing damage and major seeded bearing defects. The bearing defects were inserted on one of the planet gears of the 2nd epicyclic stage. The minor defect was simulated by machining a rectangular section with 0.3\u00a0mm in depth and 10\u00a0mm in width as shown in Fig.\u00a04. The major defect was simulated as a combination of a defected inner race (natural spalling around half of the circumference) and an outer race (around 0.3\u00a0mm in depth and 30\u00a0mm in width) as shown in Fig.\u00a05. For each fault case, three loading conditions were added to the test rig, 110% of maximum take-off power, 100% and 80% of maximum continuous power. The load condition characteristics were detailed as follows: \u2022 110% Max take-off power: power 1760 Kw, rotor speed 265\u00a0rpm, right input torque 368 Nm and left input torque 368 Nm" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002337_iros45743.2020.9341751-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002337_iros45743.2020.9341751-Figure1-1.png", "caption": "Fig. 1: A velocity-based contact localization scheme. Contact is localized to a set of candidate points (circled) where the surface normal n is co-linear with the line joining the candidate point to the center of rotation (COR). Equivalently, the velocity at the candidate point v must be perpendicular to the surface normal. Inset: Example application of a quadruped robot on a flight of stairs.", "texts": [ ". INTRODUCTION Robots are proving to be increasingly useful in unstructured environments, such as cluttered homes and outdoor terrain. However, in these environments robots must deal extensively with the making and breaking of contact with uncertain or unknown object shapes and poses. This makes contact localization a vital skill. For example, Fig. 1 shows a legged robot walking up stairs, with one of its legs in contact with the edge of a step. In this scenario, noisy estimation of the height of a stair can lead to unexpected contact between the edge and the leg causing the robot to trip. Another example is an industrial robot gripper with planar fingers approaching an object to grasp or estimate its surface, Fig. 9. In this paper, we propose a generalized extension to the method used by Barasuol et. al. [1] for planar velocity-based contact localization, summarized in Fig. 1. The method is based on the observation that if a point is in contact with a rigid body, its velocity in the direction of the surface normal must be zero, i.e., its velocity is perpendicular to the surface normal. Calculating the instantaneous velocity at each point on the surface will yield a set of candidate points. The main requirement for this method to work is that the robot be rigid and have accurate position and velocity measurements on the collision link. Any motion due to collision is directly transmitted to the link\u2019s position/velocity sensors", " (6) Based on this, the scalar velocity in the normal direction, c\u0307n(t), of a point c at time t is, c\u0307n(t) = nLi,c \u00b7 c\u0307(t), (7) where nLi,c is the surface normal of link Li at c. There are two velocity constraints that must hold at the true contact point (c\u2217). First, during contact, c\u2217 must have zero velocity in the direction normal to the link surface, c\u0307\u2217n(t) = 0 \u2200t \u2208 [t0, tf ]. (8) This is necessary for persistent contact between the link and the object. For the planar case, this constraint is equivalent to having the line from the center of rotation to the point c\u2217 be perpendicular to the robot\u2019s surface, as shown in Fig. 1. Second, at the instant before contact, denoted as t\u22120 , the point of initial contact must have a positive velocity along the surface normal of the link, c\u0307\u2217n(t\u22120 ) > 0. (9) 7353 Authorized licensed use limited to: University of Prince Edward Island. Downloaded on May 30,2021 at 09:35:38 UTC from IEEE Xplore. Restrictions apply. Thus, to localize a contact, the proposed method simply identifies the set of possible contact points, denoted C, which contains candidate points that satisfy both constraints", " For the planar case, since the set B is one dimensional, the set C is zero dimensional. In many planar cases, C will contain a unique possible contact point at time t0. For a spatial robot, the set B is two dimensional, so the set C is one dimensional. Reducing contact location ambiguity in cases where C contains more than one point is discussed in Section VI-B. We now apply this method to an example robot with simple link geometry. Consider a Minitaur robot that has collided with the edge of a stair, Fig. 1. Using one of the methods described in [15], the robot has detected that the collision occurred at time t0 and lasts until time tf . Fig. 3 shows the link of the robot that has made contact. The portion of this link\u2019s surface that is expected to make contact, B, is highlighted in green. Any arbitrary point c on this highlighted region can be characterized by the variable l \u2208 R, where |l| is the distance away from the frame Li, and sgn(l) denotes which side of the link the point is on. If d is the width of the link, then c = [ |l| d 2sgn(l) ]T ", " To compare the velocity-based method and position-based method, one of the robot\u2019s legs was swept into a stationary object and the two methods were used to estimate the contact locations. These estimated contact locations were then compared to ground truth contact location measurements. To get these accurate ground truth measurements, the obstacle was rigidly attached to the body of the robot, and the dimensions of the rig connecting the obstacle to the robot were measured. This experiment provides a scenario similar to a legged robot on stairs (Fig. 1), but provides ground truth contact locations, which was used to evaluate contact location estimates, within 1 cm of the true contact location. The experiment setup is shown if Fig. 6. Six different contact positions were used. At each position, 100 estimations of the contact location were made using the three different methods. The actual contact locations are shown as the dots in Fig. 7 and the estimated contact locations from the position-based, and velocity-based methods are shown as the crosses" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002661_j.matdes.2021.109851-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002661_j.matdes.2021.109851-Figure1-1.png", "caption": "Fig. 1. Centrifugal impeller STL file domain used for the thermal simulation at the component scale. Open source STL file is obtained from [28]. The diameter of the impeller is d \u00bc 65 mm and the height of the impeller is h \u00bc 31 mm.", "texts": [ " The two scales are bridged using the temperature field h\u00f0x; t\u00de, that is an output of the thermal simulation model and an input to the precipitate model. In conjunction with the predicted thermal histories, we analyse the precipitate microstructure throughout the component under various printing conditions and conclude our analysis with the subsequent heat treatment simulations. For demonstration purposes, we choose the geometry represented by the open-source centrifugal impeller STL file [28] shown in Fig. 1. The diameter of the impeller is d \u00bc 65mm and the height of the impeller is h \u00bc 31mm. The domain of the STL file is first discretized into voxels that are interpolated using suitable shape functions to form the isoparametric finite elements. The voxelization algorithm and the visualization tools are written using the open-source VTK library functions [29] as part of a larger in-house C++ code. The voxelized domain of the size of 80 80 80 voxels that is used for the thermal simulations is shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000919_icuas.2019.8798308-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000919_icuas.2019.8798308-Figure7-1.png", "caption": "Fig. 7: Fail safe system of a hexacopter", "texts": [ " The attitude (\u03c6, \u03b8, \u03c8) and the position (x, y, z) of the quadcopter can be seen in Fig. 2 and Fig. 3, respectively. In the NMPC, the attitude is constrained from -0.2 to 0.2 rad. It can be seen from Fig. 4 that the diagnosis algorithm converges to the actual value. In this case, the adaptive gain Q = 0.01I12. The control inputs from NMPC can be seen in Fig. 5, while the position of the quadcopter in 3D can be seen in Fig. 6. The method is used to detect a complete actuator failure (\u03b8 = 14) and use this information to activate the parachute system (Fig. 7). The hexacopter used in this work is a S550 hexarotor with the PixHawk 2.1 Cube FC. The approximate weight of the hexacopter is 2kg. The diagnosis module is used to detect rapid changes in the accelerometer and subsequently shutdown the motor and activate the parachute system. The hexacopter operates in a normal condition in the first 6s before a complete actuator failure is detected and all motors are shutdown and the hexacopter experiencing free fall. The flight data from the accelerometer are recorded, which can be seen from Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002484_lra.2021.3061361-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002484_lra.2021.3061361-Figure5-1.png", "caption": "Fig. 5. The 3D design model of TAWE [9] on a right forearm with the major components labeled, where \u03b8FE is the approximate angle of flexion-extension, \u03b8RUD is the approximate angle of radial-ulnar deviation, and \u03b8SP is the approximate angle of supination-pronation.", "texts": [ " This observation is expected since the design of Jw in IO-RAC takes into consideration the larger disturbances at \u03b81 and \u03b83, which is unlike the case of PD where simply large control gains are used (ub for PD is designed as ub,PD = \u22124c1K2\u03be). Therefore, with the reasonable design of Jw, IO-RAC can potentially suppress disturbances at their origins, and prevent them from transmitting to other parts of the system. The stationary exoskeleton example allows us to observe the fundamental behaviors of IO-RAC. The next case studies IORAC used in a wearable exoskeleton for tremor suppression. Our team is developing TAWE [9] - a wearable wrist exoskeleton for active pathological tremor alleviation and movement assistance as shown in Fig. 5. TAWE features a 6-DOF rigid linkage mechanism that allows unconstrained wrist movement. Therefore, unlike the previous stationary exoskeleton case where the kinematics is defined by the exoskeleton mechanism, the kinematics of TAWE is defined by the biomechanism of the wrist. The detailed modeling of human-TAWE dynamics is explained and validated in [9]. For this study, we consider a simpler case where the forearm is fixed. This leads to the forearm being a 3-DOF system with q1 = [\u03b8RUD, \u03b8FE, \u03b8SP] T as the wrist 3D rotation angles (i" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001158_s00170-020-04957-z-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001158_s00170-020-04957-z-Figure3-1.png", "caption": "Fig. 3 The position of the cutting point changes after the hob moves in the Y-direction", "texts": [ " The calculation formula of KK\u2019 can be expressed as F\u03b1 \u00bc KK 0 \u00bc \u03b4xsin\u03b1 \u00f01\u00de where F\u03b1 represents the total error of tooth profile; \u03b1 is the angle between KK' and the horizontal direction. The error F\u03b1 will lead to the changes of the addendum, dedendum, tooth thickness, and the diameters of apex and root circles. This kind of error is relatively easy to eliminate, and it can be compensated by adjusting the feed speed in X-direction. The thermal drift of the workpiece spindle along the axial direction of the hob is usually defined as the error of Y-direction. As shown in Fig. 3, the calculation formula of KK' induced by \u03b4y can be expressed as F\u03b1 \u00bc KK 0 \u00bc \u03b4ycos\u03b1 \u00f02\u00de The thermal error in Y-direction will lead to the deviation of the gears\u2019 helices, and it is tough to eliminate. The numerical control compensation can be carried out only if the variation is figured out. The structure of hobbing machine is complex. During the operation, the temperature of each part of the machine is different, so a large quantity of temperature sensors were installed to obtain the data. The multiple collinearity existing among these temperature variables can be formulated as C0 \u00fe C1T1 \u00fe C2T2 \u00fe\u22ef\u00fe CnTn \u00bc 0 \u00f03\u00de where C0, C1, C2, ", " The GRNN algorithm has good performance on converging speed to the optimal result for both large and small sample data. The solution of GRNN is formulated as E Y !jX! h i \u00bc \u222b\u00fe\u221e \u2212\u221e Y ! f Y ! ; jX! dX ! \u222b\u00fe\u221e \u2212\u221e f Y ! ; jX! dX ! \u00f04\u00de where X ! denotes the input vector; Y ! denotes the predicted result of GRNN;E Y !jX! h i denotes the true value of the output Y ! ; and f Y ! ; jX! is the joint probability density function of X ! and Y ! [28]. GRNN consists of four layers, respectively, which are the input layer, the pattern layer, the summation layer, and the output layer, as shown in Fig. 3. In the input layer, neurons receive the input vector X ! , and then it is transferred to the pattern layer. In the pattern layer, a nonlinear transformation from the input neurons to the pattern neurons is processed [28]. The neuron numbers of the input layer and the pattern layer equal to the dimension of input vector and the number of learning samples separately. The Gaussian function of Pi used in transformation is expressed as Pi \u00bc exp \u2212 X !\u2212X!i T X !\u2212X!i 2\u03c32 2 64 3 75 i \u00bc 1; 2; :::; n\u00f0 \u00de \u00f05\u00de where \u03c3 denotes the smoothing parameter; X " ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002700_s12541-021-00510-4-Figure8-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002700_s12541-021-00510-4-Figure8-1.png", "caption": "Fig. 8 Complete gear profile", "texts": [ "\u00a0(15), if the point rotating at 1 on the straight-line gear profile becomes a meshing point, it must meet the following conditions. Hence, the tooth profile of the ring gear can be expressed as and the meshing line is The basic three lines are displayed in Fig.\u00a07. (10)\ufffd\ufffd\ufffd\ufffd\u20d7v12 \u00d7 \ufffd\u20d7n = 0, (11)\ufffd\u20d7R = x1(u) \ufffd\ufffd\u20d7i1 + y1(u) \ufffd\ufffd\u20d7j1, (12)\ufffd\ufffd\ufffd\ufffd\u20d7v12 = \ufffd\ufffd\ufffd\u20d7w1 \u00d7 \ufffd\u20d7R \u2212 \ufffd\ufffd\ufffd\u20d7w2 \u00d7 ( \ufffd\u20d7R + \ufffd\ufffd\ufffd\u20d7E1), (14)(r1 cos 1 \u2212 y1) dy1 du \u2212 (x1 \u2212 r1 sin 1) dx1 du = 0. (15)k cos 1 + sin 1 = k2x1 + kb + x1 r1 (16)1 + k2 \u2265 ( k2x1 + kb + x1 r1 )2 (17)G(x2, y2) = M21R(x1, y1), (18)M(xm, ym) = Mg1R(x1, y1). 1 3 The traditional tooth profile shown in Fig.\u00a0 8 has four parts\u2014the arc de is the addendum circle and is formed during the processing of the gear blank, the arc cd is the working profile, the arc bc is the transition curve and is formed naturally by tooth cutter processing and do not participate in the meshing motion, and the arc ab is the dedendum circle. The tooth profile of the ring gear is displayed as Fig.\u00a09. It is noticeable that the tooth profile does not intersect the dedendum circle and cut through the addendum circle. Hence, the singularity of a part of the tooth profile over the addendum circle arc is noticed" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002035_j.measurement.2020.108224-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002035_j.measurement.2020.108224-Figure1-1.png", "caption": "Fig. 1. Test system for vibration and strain testing of centrifugal fan.", "texts": [ " Firstly, a set of test system is established. Accurate data for fatigue analysis and fault diagnosis are available. Secondly, a comprehensive test and analysis platform for blade fatigue are established. Different methods are used to confirm and to make sure of the objectivity of the conclusion. Finally, data acquisition is finished and various blade damages are classified. The extraction, recognition and classification of fault feature are combined with optimization method for more accurate fault pattern recognition. As shown in Fig. 1, a set of experimental system is established to realize the function of fatigue testing and fault diagnosis. The system can test the fan under different working conditions. As core of system shown in Fig. 2, PXI (control and acquisition) and SCXI (signal conditioning) system can simultaneously measure strain and vibration signals. The main system is produced by National Instruments Co Ltd. (NI). The system contains 3 functions as follows. In the first part, the experimental platform is fan and drive functional module" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002120_j.ijfatigue.2020.105894-Figure13-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002120_j.ijfatigue.2020.105894-Figure13-1.png", "caption": "Fig. 13. Schematic illustration of stress concentration points for different shapes of pores from left to right: triangular, spherical/elliptical_z, elliptical_y, and elliptical_x.", "texts": [ " 11j (with refined mesh) shows higher Kt (Kt = 3.0) than the spheroidal pore (mostly around Kt = 2.0 for all the surface spheroidal pores, as shown in Fig. 12a), comparable with the ellipsoidal pore of elli_0.5\u20130.1\u20130.1 (Kt = 2.8 as shown in Fig. 12c with refined mesh) but lower than other triangular pores demonstrating that the length, radius as well as shape of the pores are the key factors influencing the stress concentration. Based on the above study, a rule of stress concentration locations is generalized in Fig. 13. The mouth locations at the sharp corners of the open pores residing along the plane normal to the loading direction favor fatigue cracking. For the triangular pores the points are mostly located at the root of the pores or at subsurface or in middle positions. For the semi ellipsoidal and semi-spheroidal series, the stress concentration points shift slightly to the sub-surface areas as illustrated by the dotted lines in Fig. 13. The effects of the proposed pore size, position, shape and orientation on stress concentration factor Kt are discussed in this section. Some of the results are shown in Table 3 for triangular surface types of pores. Those data are plotted in Fig. 14a for the correlation of Kt to the length of pores. For different pores, Kt essentially shows a wavy pattern along with the length of a pore. Overall, the triangular pores show higher Kt values than the semi-spheroidal/ellipsoidal pores for most of the configurations, which are mainly due to the sharp radius of the triangular pores set as 10 \u00b5m, while the radius of most of the semi-ellipsoidal pores is much larger than 10 \u00b5m" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002193_s12206-020-0903-z-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002193_s12206-020-0903-z-Figure4-1.png", "caption": "Fig. 4. (a) Faulty gears; (b) rotors as medium and heavy loads.", "texts": [ " The sensitivity of the device is 10.2 mV/ (m/s\u00b2). The vibration signals captured using the accelerometers were recorded in a personal computer by means of a 4-channel DAQ system. Fig. 3 shows the experimental setup of the machinery fault simulator for capturing the vibration signals of various faulty and non-faulty bearings and gear system. The vibration data is sampled at a rate 25.6 KHz and the shaft speed is kept at 600 rpm with 3 load conditions: no load, medium load and heavy load for bevel gear and rolling bearings. Fig. 4(a) shows the faulty gears and Fig. 4(b) shows the rotors applied as loads for medium load (450 gm) and heavy load (5 kg) conditions. The 1D vibration signal of healthy bearing and 3 faulty conditions of bearing, namely inner race fault, outer race and ball fault are illustrated in Fig. 5. Fig. 6 shows the 1D vibration signals of healthy gear, gear with missing teeth, gear with chipped teeth and the shafts in balanced and unbalanced conditions. The 1D vibration signal is converted to 2D gray scale image. The process is done in two steps", " 16 depicts that the best validation performance for rolling bearing is achieved at epoch 38 and Fig. 17 shows that at no load condition the best validation performance is achieved at epoch 25 for denoised gear system. At no load, accuracy of the classifier is 100 % for both the rolling bearing and the gear as shown by their confusion matrices in Figs. 18 and 19. For verification of effectiveness of the proposed method, experiments are also conducted by varying the load conditions. These loads are shown in Fig. 4(b). To lay emphasis on the improvement in classification accuracy by enhancing the vibration images, the classification results are computed for the noisy vibration images and are then compared with the classification results of the enhanced vibration images. From Table 1 and Figs. 20-22, it can be clearly observed that although both the raw vibration images and the enhanced vibration images have been classified using the same features, classification results have improved significantly after enhancing the raw vibration images" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000187_s10846-019-01113-y-Figure16-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000187_s10846-019-01113-y-Figure16-1.png", "caption": "Fig. 16 The trajectory of base (red line) and end-effector(blue line) by A* a CMM b EMM", "texts": [ " Then EMM should raise its end-effector and get prepared. After that, CMM attaches EMM and carries it to the cargo station where EMM can finish its ORU transport operation. The entire operation is briefly demonstrated in Fig. 14. Video \u2019Comprehensive Operation. avi\u2019 shows the detail. Figure 15 demonstrates the trajectory of base and end-effector for both CMM and EMMwith blue standing for the base and red for the end-effector. In comparison, the trajectory provided by the non-any-angle-path algorithm (basic A*), illustrated in Fig. 16, is neither smooth nor reasonable. Note that the collision between CMM and EMM as well as EMM against the environment should also be considered especially when CMM is approaching and carrying EMM. The collision between ORU and the station should also be avoided when being transported. In this case, planning by setting threshold to collision margin only takes 31.717% of time taken by planning without the improved collision prediction method. The time consumption ratio between planning using the proposed method and with the A* algorithm, using the same collision prediction method, is 112" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001836_j.jfranklin.2020.03.017-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001836_j.jfranklin.2020.03.017-Figure2-1.png", "caption": "Fig. 2. Two-link planar manipulator.", "texts": [ " Moreover, by (7) , we ave \u03c8 \u03b5 \u2016 q\u2016 = \u03c8 \u03b5 \u2016 z 1 + y r \u2016 \u2264 \u03c8 \u03b5 \u2016 z 1 \u2016 + \u03c8 \u03b5 \u2016 y r \u2016 , \u03c8 \u03b5 \u2016 v\u2016 = \u03c8 \u03b5 \u2016 z 2 + k 1 \u03c11 z 1 \u2016 \u2264 \u03c8 \u03b5 \u2016 z 2 \u2016 + \u03c11 \u03c8 \u03b5 | k 1 |\u2016 z 1 \u2016 , hich implies that sup t>t 0 \u03c8 \u03b5 \u2016 q\u2016 < + \u221e , sup t>t 0 \u03c8 \u03b5 \u2016 v\u2016 < + \u221e since all the terms on the ight-hand side of above inequalities are bounded on [ t 0 , + \u221e ) . . Application examples In this section, the theoretical results of the paper are validated by two application examles, i.e., two-link planar manipulator and nonlinear benchmark system. For each case, the imulation results are compared with those using the traditional PID control method. .1. Application example (1): Two-link planar manipulator Consider two-link planar manipulator system as shown in Fig. 2 [20] , where m i , l i denote he mass and length of link i , respectively; q i (i.e., the generalized coordinate) denotes the ngle from the vertical line to the i -th pendulum with unit rad; u i is the torque with respect o the i th joint; \u03be i \u2019s are the horizontal and vertical disturbance accelerations of the suspenion point, i = 1 , 2. Let q = (q 1 , q 2 ) T , v = \u02d9 q = (v 1 , v 2 ) T , u = (u 1 , u 2 ) T . By using Hamilton\u2019s Please cite this article as: J. Li, J. Li and Z. Wu et al., Practical tracking control with prescribed transient performance for Euler-Lagrange equation, Journal of the Franklin Institute, https:// doi" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001846_j.mechmachtheory.2020.103877-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001846_j.mechmachtheory.2020.103877-Figure4-1.png", "caption": "Fig. 4. Schematic picture of the CVT with specification of local references.", "texts": [ " The former is obtained by projecting the relative angular velocity vector onto a direction orthogonal to the contact surfaces at the contact spot. Spin velocity causes losses in torque transmission and can generate considerable side forces on the rotating bodies [16] . By Fig. 2 , the relevant relative angular velocity vectors can be evaluated as follows: \u03c9 S0 = \u03c9 S s \u2212 \u03c9 0 j = \u03c9 S ( cos \u03b3 j \u2212 sin \u03b3 k ) \u2212 \u03c9 0 j \u03c9 Sa = \u03c9 S s + \u03c9 a j = \u03c9 S ( cos \u03b3 j \u2212 sin \u03b3 k ) + \u03c9 a j \u03c9 Sb = \u03c9 S s + \u03c9 b j = \u03c9 S ( cos \u03b3 j \u2212 sin \u03b3 k ) + \u03c9 b j \u03c9 S2 = \u03c9 S s \u2212 \u03c9 2 j = \u03c9 S ( cos \u03b3 j \u2212 sin \u03b3 k ) \u2212 \u03c9 2 j (10) where the unit vectors are those defined in Fig. 4 . The spin velocity of two bodies in contact is the component of the relative angular velocity vector onto the normal direction evaluated at the contact point. Spin coefficients are then calculated as spin velocities normalized by a reference angular velocity modulus. Combining the definition with Eqs. (7) and (10) gives the four spin coefficients are derived as functions of the tilt angle and the creep coefficients: \u03c3S0 = \u03c9 S0 \u00b7 z 0 | \u03c9 0 | = \u2212\u02dc r0 tan (\u03b80 + \u03b3 )(1 \u2212 C R 0 ) + sin \u03b80 (11) \u03c3Sa = \u03c9 Sa \u00b7 z a | \u03c9 a | = \u2212 sin \u03b1 \u2212 \u02dc ra tan (\u03b1 \u2212 \u03b3 ) (1 \u2212 C Ra ) (12) \u03c3Sb = \u03c9 Sb \u00b7 z b | \u03c9 b | = \u02dc rb tan (\u03b1 + \u03b3 ) (1 \u2212 C Rb ) + sin \u03b1 (13) \u03c3S2 = \u03c9 S2 \u00b7 z 2 | \u03c9 2 | = \u02dc r2 tan (\u03b82 \u2212 \u03b3 ) (1 \u2212 C R 2 ) \u2212 sin \u03b82 (14) where the unit vectors of z-axes are defined as follows: z 0 = \u2212 sin \u03b80 j + cos \u03b80 k z 1a = \u2212 sin \u03b1j \u2212 cos \u03b1k z 1b = sin \u03b1j \u2212 cos \u03b1k z 2 = sin \u03b82 j + cos \u03b82 k Spin coefficients are purely kinematic quantities" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002011_j.jmapro.2020.05.026-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002011_j.jmapro.2020.05.026-Figure4-1.png", "caption": "Fig. 4. The inside-beam powder feeding (IBPF) technology.", "texts": [ " A closed-loop control strategy was devised; (2) The measured height values in different positions on a cladding track can be memorized after the deposition of a layer, and the scanning strategy and process parameters for the next layer can be planned in advance. The uneven top surface is then smoothed after multi-layers. In order to maintain the nozzle direction normal to the current layer shown in Fig. 1, our research group developed an inside-beam powder feeding (IBPF) nozzle [32]. The principle and product of this nozzle are illustrated in Fig. 4a and b respectively. Compared with the coaxial powder feeding technology [33,34], which is mostly studied and commonly applied in the market at present, the IBPF nozzle splits the solid laser beam into an annular and cone-shaped focusing laser beam. If the annular laser spot is scanning, the energy density of the annular laser beam displayed energy peaks at the both sides [35], which can be more uniform than the solid laser spot [12]. The powder beam is sprayed towards the central axial direction from a single powder pipe. The shielding/alignment gas sprays surround the powder beam coaxially, thus restricting the powder/gas flow from divergence and enabling an atmosphere protection for the molten pool. When the nozzle is inclined, the drag forces on the powder flow from both carrier gas and shielding gas to the same axial direction can reduce the action of the gravity significantly and achieve a better straightness of the powder flow. With an inclined angle of 135\u00b0 in Fig. 4c, the powder flow of the coaxial powder feeding nozzle diverges due to gravity. In contrast, the powder flow of IBPF nozzle maintained a slight divergence and excellent straightness out of the powder pipe outlet, and exhibited almost no parabolic trajectory. The schematic diagram of the LMD system is illustrated in Fig. 5. The system consisted of an IPG YLS-2000-TR fiber laser as the heat source, a GTV PF 2/2 powder feeder, a robot of KUKA KR 60-3 F and a Beckhoff EK1100 Ethernet I/O. In order to measure the actual deposition height, a CCD camera was fixed on the IBPF nozzle, which is capable to move and rotate with the nozzle synchronously (Figs", " The both programs in the robot control cabinet and industrial computer work in independence threads with a real-time data transmission through ProfiNet. The substrate material used was 304 stainless steel. The powder material was the Fe-based alloy Fe313 (composition in wt,\u2212%: C\u22480.1; Si= 2.5\u20133.5; Cr= 13\u201317; B=0.5\u20131.5; Fe= bal.) with the particle size of 75\u2212150 \u03bcm. Both shielding gas and powder carrier gas were nitrogen. The \u201cself-healing\u201d effect in the LMD process was validated in multiple studies [36,37]. In a certain range of the defocus distance d (Fig. 4a), the deposition height of the actual layer varies with the working distance w (Fig. 4a) inversely. That means in Fig. 3, the deposition height of a new layer is lower in convex points and higher in concave points with a varying working distance. Therefore, the uneven surface can be automatically smoothed after several layers under an open-loop control. The \u201cself-healing\u201d effect of an IBPF nozzle with a range of \u22123mm \u223c \u22125mm of the defocus distance d was verified in [37]. The model of the IBPF nozzle used in this study has a larger \u201cselfhealing\u201d effect range of \u22123mm \u223c \u22129mm. In a layer or a segment of an equal-height deposition, the \u201cselfhealing\u201d effect can be taken advantage of when the process parameters are invariant" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001122_tie.2020.2965484-Figure10-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001122_tie.2020.2965484-Figure10-1.png", "caption": "Fig. 10. The experiment platform. (a) The prototype assembly. (b) Locations of TS.", "texts": [ " According to the conservation of energy, the input power source of six nodes in boundary can be calculated and updated in the ETN model, which is the next step for updated temperature Tl1-Tl6 and forms an entire loop. Finally, the convergence of node temperatures in ETN is achieved and temperature distributions within slot are obtained. A maximum number of iterations is also set to avoid dead cycle. The complete flow chart of coupled ETN model calculating is shown in Fig. 9. IV. FEA AND EXPERIMENTAL VERIFICATION The coupled ETN model is implemented to verify the accuracy and validity by comparing with the FEA and experimental results, including the temperatures of main parts and distributions within slot. Fig. 10(a) shows the prototype experiment platform of the SPM machine. The thermal class of the magnet wire is F, whose tolerable temperature limit is 155\u00b0C. The magnet wire with thermal class F is used from the perspective of reliability. TSs (Thermal Sensors) consist of Ni-Cr/Ni-Si thermocouple wires (K-type) and Ni-Co alloy thermal resistances (KTY84/130), and own the excellent accuracy and required ranges. TSs are buried in winding with end-winding and attached on stator yoke. The exact locations of TSs are shown in Fig. 10(b). Thermal imager is utilized for the case temperature. Authorized licensed use limited to: University of Exeter. Downloaded on June 08,2020 at 00:27:11 UTC from IEEE Xplore. Restrictions apply. 0278-0046 (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. After the temperature reaches steady-state at the rated-current of 17A, the temperature distributions of winding obtained from ETN and FEA are compared in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001817_s40964-020-00123-9-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001817_s40964-020-00123-9-Figure4-1.png", "caption": "Fig. 4 LPBF-fabricated SS316L impeller along with the associated coordinate system", "texts": [ " Noted that the texture intensities in the pole figures were not as concentrated as those in the thin strut due to the presence of some stray grains. Therefore, the cube sample will be hereafter described as \u201csingle-crystalline-like\u201d. For the much larger wall, the {110}\u2329001\u232a Goss texture still held. It also exhibited a single-crystalline-like crystallograhic texture, albeit more stray grains were observed and the texture intensities slightly decreased. The simulated pump impeller shown in Fig.\u00a02 can be readily manufactured with good integrity, as presented in Fig.\u00a04. The blades of the impeller were categorized into vertical, 45\u00b0 and horizontal blades according to their orientations regarding the recoating direction. Electron backscattered diffraction (EBSD) analyses were carried out on the simulated pump impeller with emphasis on different blades and their connections with the base, as shown in Fig.\u00a05. Noted that for easy comparison, the two EBSD IPF colouring orientation maps of all the locations shown here were adjusted to be aligned with the universal X (SD1) and Z (BD) direction" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002450_j.compstruct.2021.113608-Figure21-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002450_j.compstruct.2021.113608-Figure21-1.png", "caption": "Fig. 21. The position space of the fiber placement machine.", "texts": [ " X/mm Y/mm Z/mm A/\u00b0 B/\u00b0 D/\u00b0 Travel \u22128970 to 8970 \u22123500 to 3500 \u2212235 to 1990 \u221220 to 20 \u221290 to 90 \u2212180 to 180 Nx \u00bc sin\u00f0A\u00de sin\u00f0C\u00de \u00fe cos\u00f0A\u00de cos\u00f0C\u00de sin\u00f0B\u00de \u00f034\u00de Ny \u00bc cos\u00f0A\u00de sin\u00f0B\u00de sin\u00f0C\u00de cos\u00f0C\u00de sin\u00f0A\u00de \u00f035\u00de Nz \u00bc cos\u00f0A\u00de cos\u00f0B\u00de \u00f036\u00de Where X, Y, Z are the coordinate origin of the automated fiber placement machine,CX ;CY ;CZ represents the position of the C axis rotation center in the machine coordinate system,BX ;BY represents the offset parameter of the B\u2010axis rotation center relative to the C axis rotation center, AX ;AY ;AZ indicates the offset parameter of the A\u2010axis rotation center relative to the B\u2010axis rotation center, DX ;DY indicates the offset parameter of the D\u2010axis rotation center relative to the A\u2010 axis rotation center, and RX ;RY ;RZ indicates the offset parameters of the center of the end of the roller relative to the center of rotation of the D\u2010axis. According to the algorithm above, the position accessible space of the AFP machine is shown in Fig. 21, and the mapping of the attitude accessible space on the Gaussian sphere is shown in Fig. 22. In the third section, two methods were used to establish the mathematical relationship between the Gaussian curvature of the surface, the width of the prepreg tow, and the deformation limit per unit length of the prepreg tow. Since the existing machines developed by our research team currently only supports the laying of 1/4 in. prepreg tow. Moreover, there was only domestic prepreg EH104 at that time, and its corresponding minimum turning radius was 1500 mm which was measured by experiment under the conditions of our common laying technology" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000539_s00170-016-9337-2-Figure22-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000539_s00170-016-9337-2-Figure22-1.png", "caption": "Fig. 22 Test surface", "texts": [ "1. Real blade machining experiments are shown in Sect. 7.2 to verify this algorithm. To verify that the hybrid method can further improve the efficiency, we made simulations using VERICUT\u00ae, which is presented in Sect. 7.3. The proposed methods are developed using Visual Studio 2010 as a model of CAM Software. To further verify that the tool path generation algorithm can be used in complex surfaces, wemodify the real platform surface to a test surface (with variation of curvatures), as shown in Fig. 22. This test case does not represent a realistic blade part, it is provided as an illustration only. The triangular meshes on the test surface are shown in Fig. 23. The spiral tool path is shown in Fig. 24; the step-over is controlled using the method in Sect. 5.2. The hybrid tool path (with zig ratio \u03bb=0.75) is show in Fig. 25. In this research, we make two blade machining experiment cases to verify the field-based tool path generation algorithm: five-axis blade platform surface machining (roughing and finishing) and four-axis blade surface roughing" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000378_2016-01-1221-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000378_2016-01-1221-Figure2-1.png", "caption": "Figure 2. Coil end schematics", "texts": [ "9%, respectively, and reduces the volume and weight by 24% and 26%, respectively. This HVW has been applied to Honda\u2019s new-structure (Honda Advanced Winding: HAW) motor that achieves top-in-class power and compactness. (1) CITATION: Ito, K., Shibata, T., and Kawasaki, T., \"Development of High Voltage Wire for New Structure Motor in Full Hybrid Vehicle,\" SAE Int. J. Alt. Power. 5(2):2016, doi:10.4271/2016-01-1221. 272 coil-forming without insulate paper, which also enables to reduce the size compared to the conventional distributed winding structure (Fig. 2). The motor application technology used HVW (made by Furukawa Electric Co., Ltd,) will be reported. Compared to the previousstructure motor, the developed HVW maintained the same efficiency, and helped to increase the maximum torque and maximum power by 2.6% and 8.9%, respectively, and reduce the volume and weight by 24% and 26%, respectively. The conventional i-MMD motor uses a round wire that is suitable for configuring a distributed winding structure, but limits the space factor within the slot" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001360_s42417-020-00220-7-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001360_s42417-020-00220-7-Figure3-1.png", "caption": "Fig. 3 Second-stage epicyclic gears", "texts": [ " This combined drive provides power to the tail rotor drive shaft and the bevel gear. The bevel gear reduces the rotational speed of the input drive to 2,405\u00a0rpm and changes the direction of the transmission to drive the epicyclic reduction gearbox module. The second section is the epicyclic reduction gearbox module which is located on top of the main module. This reduces the rotational speed to 265\u00a0rpm which drives the main rotor. This module consists of two epicyclic gears\u2019 stage, the first stage contains eight planet gears and the second stage with nine planet gears, see Fig.\u00a03. Specifications of the gears are documented in Table\u00a01. The epicyclic module planet gears are designed as a complete gear and bearing assembly. The outer race of the bearing and the gear wheel is a single component, with the bearing rollers running directly on the inner circumference of the gear. Each planet gear is \u2018self-aligning\u2019 by the use of spherical inner and outer races and barrel-shaped bearing rollers (see Fig.\u00a03). The test procedure consists of three experimental conditions, including fault-free condition, minor seeded bearing damage and major seeded bearing defects. The bearing defects were inserted on one of the planet gears of the 2nd epicyclic stage. The minor defect was simulated by machining a rectangular section with 0.3\u00a0mm in depth and 10\u00a0mm in width as shown in Fig.\u00a04. The major defect was simulated as a combination of a defected inner race (natural spalling around half of the circumference) and an outer race (around 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002259_j.measurement.2020.108723-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002259_j.measurement.2020.108723-Figure3-1.png", "caption": "Fig. 3. Winding distribution of the studied induction motor.", "texts": [ " Measurement 171 (2021) 108723 \u3008M(\u03b8,\u03c6) \u3009 = 1 2\u03c0\u3008g\u2212 1(\u03b8,\u03c6) \u3009 \u222b 2\u03c0 0 n(\u03b8,\u03c6)g\u2212 1(\u03b8,\u03c6)d\u03c6 (19) \u3008 g\u2212 1(\u03b8,\u03c6) \u3009 = 1 2\u03c0 \u222b 2\u03c0 0 g(\u03b8,\u03c6)d\u03c6 (20) To simplify those inductance calculation formula mentioned above, (18) can be substituted into (17). Then the simplification process can be expressed as followings: Lij(\u03b8,\u03c6) = \u03bc0rl \u222b 2\u03c0 0 ni(\u03b8,\u03c6) ( nj(\u03b8,\u03c6) \u2212 \u3008 Mj(\u03b8,\u03c6) \u3009 ) g\u2212 1(\u03b8,\u03c6)d\u03c6 = \u03bc0rl \u222b 2\u03c0 0 ninjg\u2212 1d\u03c6 \u2212 \u03bc0rl \u222b 2\u03c0 0 ni ( 1 2\u03c0\u3008g\u2212 1\u3009 \u222b 2\u03c0 0 njg\u2212 1d\u03c6 ) g\u2212 1d\u03c6 = 2\u03c0\u03bc0rl \u3008 ninjg\u2212 1\u3009 \u2212 2\u03c0\u03bc0rl \u3008nig\u2212 1\u3009\u3008njg\u2212 1\u3009 \u3008g\u2212 1\u3009 = 2\u03c0\u03bc0rl \u3008 NiNjg\u2212 1\u3009 \u2212 2\u03c0\u03bc0rl \u3008Nig\u2212 1\u3009\u3008Njg\u2212 1\u3009 \u3008g\u2212 1\u3009 (21) The simplified structure of the studied IM is shown in Fig. 3. Stator has 24 slots where four slots are assigned for each pole per phase, Rotor has 18 slots where three slots are assigned for each pole per phase. The reference of the stator circumambient (\u03c6) and the reference of the rotor circumambient (\u03b8r\u03b8r) correspond to the center of winding \u2018A\u2019 and the center of winding \u2018a\u2019 respectively. Considering the number of NsNs turns for each pole per stator\u2019s phase, the number of Ns/4 turns would exist in each slot. Also, considering the number of Nr turns for each pole per rotor\u2019s phase, the number of Nr/3 turns would exist in each rotor\u2019s slot [46]" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001877_j.promfg.2020.04.275-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001877_j.promfg.2020.04.275-Figure6-1.png", "caption": "Fig. 6. Scheme of the ring rolling automat.", "texts": [ " The sheet substrate is made of S235JR with a thickness of 5 mm. The mechanical properties of the pure welded material without heat treatment are provided by [9], stating an ultimate tensile strength of 500-640 MPa, a yield strength of more than 420 MPa and an elongation at break of more than 22%. The process parameters used for each layer are displayed in Table 2. The total WAAM process time was approximately 4 hours including cooling time. Ring rolling is realized on a mechanical ring rolling automat KFRWtneo (see Fig. 5 and Fig. 6) by SMS group GmbH. The machine name is derived from the former name of the inventor company Kreuser. KFRWt means Kreuser spring ring rolling mill semi-automatic. The design concept of the KFRWt has been known since the beginning of the 20th century [10]. One major advantage of the KFRWtneo is the low lead time of several seconds. Ring diameters of up to 500 mm can be manufactured on this machine. The KFRWtneo is equipped with four rolling stations installed on a rotation unit (see Fig. 6). Each station itself is equipped with a centering arm, a mandrel and a rolling table. The rotation unit surrounds a main roll. The axis of the main roll and axis of the rotation unit are eccentrically arranged. By this eccentrical displacement, each rolling station has a specific radial roll gap between the mandrel and the main roll. During one revolution of the rotation unit, the rolling gap decreases continuously and the ring wall thickness decreases, while the ring diameter increases. When the minimum gap is reached (i" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0003084_adem.202100611-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0003084_adem.202100611-Figure1-1.png", "caption": "Figure 1. Definition of the build direction: a) the 0 sample, b) the 45 sample, and c) the 90 sample.", "texts": [ " Young\u2019s modulus, the ultimate displacements, and the ultimate strength under different loading conditions were tested and obtained. Fracture morphologies were observed and analyzed. In our study, four types of specimens were designed and tested under compression load, bending load, shear load, and torque load. Each kind of specimen were classified into three types according to the angle between the cylindrical axial and the build direction: the 0 sample, the 45 sample, and the 90 sample, shown in Figure 1. All samples were built with the same AM parameters separately, and were annealed for 5 h at 750\u2013850 C, and then manufactured according to testing standards. A cross-hatching scanning strategy was used. The chemical compositions of the Ti6Al4V powder and the LPBF processing parameters are shown in Table 1 and 2, respectively. Table 2. LPBF processing parameters. Laser power [W] Laser beam diameter [mm] Scanning speed [mm/s] Layer thickness [mm] 350 0.08 1000 0.06 Table 1. Chemical compositions of Ti6Al4V powders (wt%)" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000731_j.humov.2019.02.003-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000731_j.humov.2019.02.003-Figure3-1.png", "caption": "Fig. 3. Simulation set up for the first of three validation and numerical convergence tests. The kayak is placed in a stagnant tank of water 3.0 m wide, 1.0 m deep and 40.0m long. The kayak moves with a prescribed constant forwards speed (u). The resulting drag force is calculated for comparison.", "texts": [ " However such test cases have not specifically addressed the fluid dynamics around a moving hull or a moving paddle. The first test case evaluates the ability of the SPH solver to accurately determine the drag on a streamlined body whilst the following test is for drag on a moving bluff body. The accuracy of the fluid simulation around the hull is evaluated here by comparison with experimental measurements of kayaks moving at constant speed (Banks, Phillips, Turnock, Hudson, & Taunton, 2014; Gomes, Ramos, Concei\u00e7\u00e3o, Vilas-Boas, & Vaz, 2012; Jackson, 1995; Tzabiras et al., 2010). Fig. 3 shows the model configuration used for this test case. The hull is the same as for the full athlete-kayak model (Fig. 1) and the tank of water is 3.0 m wide, 1.0m deep and 40.0 m long. The kayak is settled in the water in an initial simulation step and then the kayak is prescribed to move forwards at a constant speed with all other degrees of freedom fixed. Total force on the kayak is calculated at each timestep using Eq. (6). Fig. 4 shows the mean drag force for four values of forwards speed and the measured data from four studies by others (Banks et al" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000098_j.engfailanal.2019.104190-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000098_j.engfailanal.2019.104190-Figure5-1.png", "caption": "Fig. 5 Contact stress distribution from KISSsoft", "texts": [ " In case 1, the boundary element method approximates a pattern similar to FE. However, FE results from LDP show some stress concentration closer to one side. Moreover, for case 2, both the method approximated pattern shifting and higher maximum stress value at the end of the face width. In case 3, maximum stress value is also higher compared to the Abaqus result, but the pattern represents agreement as the pattern shifted back to the original case when no misalignment was introduced. Case Contact stress Root stress 1 2 3 Fig.5 shows the results for contact stress distribution throughout the teeth from KISSsoft. It is convincing that in case 1, the stress is distributed throughout the teeth but for case 2 the stress is shifted towards the corner at the end of face width for both the gears due to the inclusion of the misalignment. In case 3 the microgeometry modification compensated the misalignment and the contact pattern is better than case 2. In Table 5, normalized results are presented for comparison. For data confidentiality, all results are normalized to case 1 data resulted from Abaqus, as a reference" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002450_j.compstruct.2021.113608-Figure23-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002450_j.compstruct.2021.113608-Figure23-1.png", "caption": "Fig. 23. The surfaces of test models.", "texts": [ " In the third section, two methods were used to establish the mathematical relationship between the Gaussian curvature of the surface, the width of the prepreg tow, and the deformation limit per unit length of the prepreg tow. Since the existing machines developed by our research team currently only supports the laying of 1/4 in. prepreg tow. Moreover, there was only domestic prepreg EH104 at that time, and its corresponding minimum turning radius was 1500 mm which was measured by experiment under the conditions of our common laying technology. Two types of part surfaces were shown in Fig. 23. Firstly, the Gaussian curvature distributions of them were obtained by means of the CATIA secondary development tool as shown in Fig. 24. As shown in Fig. 24, the extreme values of the Gaussian curva- ture of first part are and, and the extreme values of the Gaussian curvature of second part are and. Next the suitable range of Gaussian curvature for EH104 was calculated according to Eq. (14), and then the placement suitability of two part were analyzed by comparing the extreme values of the Gaussian curvature with the suitable range of Gaussian curvature" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002700_s12541-021-00510-4-Figure12-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002700_s12541-021-00510-4-Figure12-1.png", "caption": "Fig. 12 Limit position without interference", "texts": [ " On point B, the trapped volume becomes larger as the gear tip clearance gets bigger and the meshing area becomes shorter, resulting in a smaller overlap ratio. The modified ring gear profile is displayed in Fig.\u00a010, and the numerical model of the ring gear is presented in Fig.\u00a011. 1 3 If point B is shifted to point C shown in Fig.\u00a09, the overlap ratio will increase and lead to interference, and it will occur easily when the tooth number difference between the gear and the ring gear is too small. The interference phenomenon of the tooth profile is exhibited in Fig.\u00a012. During interference, the gear tip and the ring gear tip collide. Before performing the interference analysis, the manufacturing process should be analyzed. The ring gear is formed by the corresponding gear as a cutter, and the cutting process is equivalent to a complete meshing motion. In Fig.\u00a013, the envelope line on the ring gear left by the gear (termed as the family of cutter curves) is the tooth profile of the ring gear. Therefore, based on the formation process of the ring gear, the interference could be checked by the tooth profile reversal method" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002175_icra40945.2020.9197401-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002175_icra40945.2020.9197401-Figure2-1.png", "caption": "Figure 2. View of the FRA with wrapped fiber element (black helix). lFRA: 90 mm, wFRA: 20 mm, rFRA: 7 mm.", "texts": [ " The manipulator can be used for different applications in health, industry and other specific scenarios in which soft actuators are required for improving the safety, but also an adequate stiffness and dexterity are essential features to be achieved. The designed manipulator is shown in Fig. 1. The bidirectional manipulator consists of an SCS inside an encapsulation layer and two identical FRAs both placed on each side of the layer. While the FRAs control the position of the manipulator, the SCS provides the variable stiffness. The FRAs and the SCS are driven by air pressure aiming to minimize the control complexity and response time while providing position independent stiffness. A. Fiber-Reinforced Actuator The FRA (Fig. 2) is designed in semi-circular geometry to minimize the resistance to bending of the pressurized actuator [22]. The FRA is the manipulator module for providing the bending to the entire structure. The body of the FRA is made of DragonSkin 10 (Shore Hardness: 10A, Smooth-On Inc., USA) silicone rubber, whereas DragonSkin 30 (Shore Hardness: 30A, Smooth-On Inc., USA) rubber was used for capping due to its higher hardness. Each FRA All the authors are with The BioRobotics Institute and Department of Excellence in Robotics & AI, Scuola Superiore Sant\u2019Anna Pisa, Viale Rinaldo Piaggio 34 Pontedera, PI 56025 Italy" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000967_j.ijepes.2019.105539-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000967_j.ijepes.2019.105539-Figure3-1.png", "caption": "Fig. 3. Illustration of reluctance computation and effective area calculation of elements for radially and tangentially flux tubes.", "texts": [ " The leakage fluxes in the slots and inter-magnets regions are modeled by considered reluctances in the air parts. Considering several flux tubes for the teeth and air parts, an improved MEC method is obtained for modeling purpose which allows to have better flux distribution in the air-gap compared to conventional MEC. It is clear that increasing the number of the mentioned parameters will lead to higher level of accuracy. In the proposed MEC model, the reluctance of the flux tubes are computed as Eq. (1) [25], where the radially and tangentially flux tubes are shown in Fig. 3. In Eq. (1), the reluctance of the flux tubes in the stator and rotor parts must be calculated based on the elements flux which are related to the flux density. The iron element reluctance (as written in Eq. (1)) is dependent on flowing flux through it, so the core nonlinearity should be modeled by \u00b5 Bi( )r j function. If a flux tube overlaps with two different materials (air and core material), the flux tube is divided into two parts, one for each material region. Then the lumped permeances for each part are calculated and then combined in series or parallel, using the same formulas employed for combining series or parallel conductances in electrical circuits" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001099_j.mechmachtheory.2019.103739-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001099_j.mechmachtheory.2019.103739-Figure4-1.png", "caption": "Fig. 4. Finite element model of hypoid gears.", "texts": [ " f = n p v p = 0 (14) \u03d5 1 = igp \u2217 \u03d5 c + z 0 z \u03b2 (15) where n p is unit normal vector and v p is direction of relative velocity in coordinate system S p . The parameter igp is ratio of roll. The parameters z and z 0 are the number of processing gear and blade groups, respectively. The main parameters of cutter plate are shown in Table 1 . The machine processing parameters are shown in Table 2 . By means of the mathematical model of wheel and pinion and these parameters, the finite element model is established in ABAQUS. The details of the finite element model are demonstrated in Fig. 4 . Firstly, the material property is isotropic with young\u2019s modulus 210,0 0 0 MPa and Poisson\u2019s ratio 0.3. The reference point P and P are established on pinion and wheel rotation axis and the surface 1 and surface 2 are coupled in reference points 1 2 P 1 and P 2 , respectively. Then, the contact surface be set to pinion concave and wheel convex and the boundary conditions are divide into pre-contact, applying load and mesh process. The torque load is applied on the wheel. Finally, the mesh of contact surface becomes more intensive by changing the size of local seed" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002465_j.mechmat.2021.103786-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002465_j.mechmat.2021.103786-Figure1-1.png", "caption": "Fig. 1. Schematics of the crack-growth and local strain field measurements: (1) Pre-stretching the specimens to the desired elongation \u03bby pre in pure shear geometry; (2) Imposing speckle patterns on the surface of the stretched specimens, subsequently triggering an initial crack with the observation of the crack propagation by a high-speed camera; (3) Measuring the displacement field based on the images captured by DIC method.", "texts": [ " These specimens were hot-pressed and cross-linked at 155 \u25e6C for 18 s min using sulfur (1.4 g per 100 g of rubber) as a crosslinking agent. For the crack propagation experiments, we employed the three types of specimen, i.e., (I) SBR5, (II) SBR21, and (III) a specimen prepared by imposing previously cyclic loading\u2013unloading cycles on SBR21 (SBR21-SY). The details of the preparation of SBR21-SY are described later. A custom-built experimental setup was used for the crack propagation and local strain field measurements (Fig. 1) (Mai et al., 2020; Morishita et al., 2016). The tensile measurements were conducted by a universal/tensile tester (AGS-X, Shimadzu Corp.) with a loadcell of 5 kN. Sheet specimens with a pure shear geometry (x \u00d7 y \u00d7 z = 175 \u00d7 20 \u00d7 1.0 mm3) were initially stretched to a desired elongation in the y-direction (\u03bby pre) with a constant crosshead speed of 0.3 mm s\u2212 1. In the pre-stretched state of \u03bby pre, speckle patterns were imposed on the sample surface using a custom-made speckle stamp. The crack growth in the x-direction was triggered by cutting at the edge of the sample midway between the y boundaries. The initial cut was made by sharp scissors, and the cut-length was about 5 mm or 50 mm for the fast- or slow-moving crack, respectively. A high-speed digital camera (FASTCAM Mini AX100 with a maximum frame rate of 540,000 Hz) with the aid of two high-power LED light sources (UFLS-75, U-Technology) was used to observe the crack propagation in the central region of the specimens (Fig. 1). The central region with a length of approximately 100 mm is sufficiently long to observe the crack growth in the steady-state, which was shown elsewhere (Mai et al., 2020). The measurement was conducted at 25 \u25e6C. The crack-growth velocity (V) at each \u03bby pre was evaluated from the images captured during the crack propagation. The input tearing energy (energy release rate; \u0393) at each \u03bby pre was calculated using the relationship of \u0393 = WLy,0 (Rivlin and Thomas 1953; Thomas 1955; Tsunoda et al., 2000), where W and Ly,0 are the stored elastic energy density for the stretching to \u03bby pre (W = \u222b [F/S0]yd\u03bby) and the initial length in the y-direction, respectively", " Indeed, the D values for SBR21-SY-x are smaller but comparable to those for SBR21, whereas those for SBR21-SY-y are as small as those for SBR5. SBR5 possesses inherently no appreciable stress-softening feature (Fig. S1a in SI), and exhibit no significant anisotropy in Mullins effect (Fig. S1b in SI). A similar anisotropic damage is observed in silica- and carbon black-filled SBR (Machado et al., 2012; Mai et al., 2017b) previously subjected to uniaxial loading. The steady-state crack-growth velocity (V) is measured at various degrees of imposed pre-stretching (\u03bby pre) (Fig. 1). Fig. 4a and b illustrate V as a function of \u03bby pre and the input tearing energy (energy release rate; \u0393) at each \u03bby pre, respectively. The input tearing energy at each \u03bby pre is obtained from the stored elastic energy density (W) for the stretching to \u03bby pre. W is evaluated using the unloading curve, instead of the pristine loading curve (inset of Fig. 4c). The same procedure was done in several studies (Diani et al., 2015; El Yaagoubi et al., 2018; Mzabi et al., 2011; T. Zhang et al., 2015) for elastomers undergoing considerable stress softening during pristine loading, like SBR21" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000859_1077546319856147-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000859_1077546319856147-Figure2-1.png", "caption": "Figure 2. Physical expression of nodal coordinates.", "texts": [ "rJz @\u2019 @t \u00f0d \u00de 9>>>>>>>>>>>>>= >>>>>>>>>>>>>; \u00f05\u00de Shape functions are derived from the exact solution of homogenous form of equations for either of the orthogonal planes. The displacements X, Y, \u2019, and are expressed in terms of shape functions. Applying Lagrange\u2019s equation Mtx\u00bd 0 0 Mty \u20acqe \u00fe Mrx\u00bd 0 0 Mry \u20acqe \u00fe 0 Gy Gy 0 \" # _qe \u00fe kx\u00bd 0 0 ky qe \u00bc Qe It may be noted that [Mtx] \u00bc [Mty], [Mrx] \u00bc [Mry], and [kx] \u00bc [ky]. Re-writing the above element equation in abbreviated form Me t \u00fe Me r \u20acqe \u00fe Ge\u00bd _qe \u00fe ke\u00bd qe \u00bc Qe \u00f06\u00de In equation (6), the nodal variables expressing displacements and rotations are as shown in Figure 2. Sub-matrices [Mtx], [Mrx], [kx], [Mty], [Mry], and [ky] are symmetric and of the order 4. The nonconservative forces {Fe i } for the x-z plane are given as Fe i \u00bc Z 1 0 f z, t\u00f0 \u00deNi z\u00f0 \u00dedz, i \u00bc 1, 2, . . . , 4 \u00f07\u00de The nonconservative force in the y-z plane can also be derived similarly. The x axis is vertical in the present formulation and gravity force acts vertically downward. To include inherent damping of shaft material, a proportional damping model is used, that is C\u00bd \u00bc a M\u00bd \u00fe b k\u00bd \u00f08\u00de The numerical values of a and b have been determined using the values of modal damping ratio as 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001553_tie.2020.3031516-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001553_tie.2020.3031516-Figure2-1.png", "caption": "Fig. 2. Rig for the magnet temperature measurement, (a) crosssectional view of a small-scaled motor, (b) front side and backside of a sensing board, and (c) a block diagram of measurement system.", "texts": [ " The dynamic permeability is defined as d\u03bbr ds/dir ds, whose value depends on the core saturation level. When the daxis and q-axis currents are considered as constants ir ds0 and ir qs0, the saturation level can be simplified to a function of residual flux density, Br. The Br is only affected by the magnet temperature, Tmag. Therefore, the d-axis flux can be rewritten as the function of Tmag in (6). Fig. 1 (a) shows the B-H curve of S09 35pn230 used in the rotor and stator core in the smallscaled motor in Fig. 2, which will be described in section IV. Fig. 1 (b) shows static permeability and dynamic permeability. In Fig. 1, green dashed arrows denote the direction of increasing the magnet temperature. When the magnet temperature increases, the magnet flux linkage decreases, and the static permeability increases, as shown in Fig. 1. It makes the d-axis static inductance, Lds, increases as Tmag increases because the static d-axis inductance, Lds, is directly proportional to static permeability. 0 0 0 0 0 0 , , , , ", " Under this analysis, the DC component in (4) can be exploited to estimate the magnet temperature. 0 0. r rds ds ds mag mag mag L i T T (7) 0. ds sc mag mag L T T m (8) IV. VERIFICATION OF INDUCTANCE VARIATION DUE TO MAGNET TEMPERATURE AND SPATIAL HARMONICS In this section, high-frequency inductance is measured to show inductance variation according to the rotor angle and the magnet temperature. This relationship is verified in the special rig, which is introduced in subsection A. First, a motor is designed to measure the magnet temperatures. Fig. 2 (a) shows the cross-sectional view of the designed motor. The motor has 32 holes to accommodate magnet temperature sensors. The size of the holes is 1 mm \u00d7 2 mm. In these holes, the temperature sensors are inserted and attached to the magnets in the rotor. The temperature sensor was selected as platinum temperature sensors, which has an error bound about 1.1oC in the range from -50 oC to 150 oC. Fig. 2 (b) shows the temperature sensing board, which is located on the special plate directly connected to the rotor. Therefore, the sensing board rotates with the rotor. The specification of the sensing board is listed in Table II. The digital signal processor in the sensing board has 16 analogdigital converters. Hence, the temperature at 16 points of the magnets is simultaneously measured. All the components in the board are powered by two batteries, whose type is selected according to the temperature range, as described in Table II. The overall block diagram of the measurement system is shown in Fig. 2 (c). The sensing board rotating with the rotor transfers the measured temperatures to the host on the stationary side via Bluetooth communication. The host at the stationary side receives and transfers the data to a control board for a motor drive by SCI communication. The measurement system had been described in [6]. The designed rig has been developed and employed to verify spatial harmonics. The rotor angle of the small-scaled motor in the test rig is kept as its reference by the load machine in the test set-up", " In low-speed regions, the conventional method utilized a low pass filter to extract the average of the high-frequency inductance, Ldh_av. The average value of Ldh, Ldh_av, means that only the DC component in (16) was considered in the conventional method to estimate the magnet temperature in the low-speed regions. This conventional method is not appropriate at a standstill condition because of the spatial harmonics. Fig. 5 shows the estimation errors caused by spatial harmonics when the conventional method is employed to the rig shown in Fig. 2. The error fluctuates from -25.3 oC to 16.8 oC depending on the rotor angle, whose maximum peak-to-peak value is about 42.1 oC. The result shows the necessity of considering the spatial harmonics in the magnet temperature estimation at a standstill. The coefficients for spatial harmonics can be reconstructed by using their magnitudes and phase delays. In this paper, the order of the harmonics up to 18th is only considered as (17). It means that n is 18 in (17). This approximation is reasonable because the magnitudes of harmonics, higher than 18th, are small enough to be neglected", " The closed-loop function can be derived under two assumptions that the reconstructed coefficients are the same as the original coefficients, and the difference between the estimated temperature and the real magnet temperature is small enough to be neglected. The closed-loop transfer function can be derived as (18), which looks like a 1st order low pass filter, which can suppress noise and fluctuation of the estimated temperature. _ _ __ _ _ _ 2 . 2 Ldh SS mag est Ldh SS imag est mag Ldh SS mag est Ldh SS i A T B kT T s A T B k (18) VI. EXPERIMENTAL RESULTS In this section, various experiments are executed. In subsection A, the magnet temperature has been estimated at a standstill in the implemented special rig shown in Fig. 2. In the rig with the designed small-scaled motor, the magnet temperature can be directly and accurately measured in realtime. Therefore, the estimated magnet temperature can be compared with the measured magnet temperature. The verification sequence is as follows: the rotor position is kept at a certain angle by the load machine for 12 seconds. The rotor position in an electrical angle is swept from 0 o to 358 o with a 2 o interval. The second experiment in subsection B has been designed and executed to verify the feasibility of the proposed method in an actual traction motor of a mass-produced electric vehicle in [29]" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001028_j.wear.2019.203135-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001028_j.wear.2019.203135-Figure5-1.png", "caption": "Figure 5: a) The distribution of inner and outer loads b) Applying lateral and axial displacements on the bearing", "texts": [ " Then, the value of the load distribution factor \u03b5 is calculated by using the approximate value of and the radial clearance ur as: max 1 1 2 2 r r u u \u03b5 \u03b4 = \u2212 + . (6) Then, the value of the radial load integral Jr(\u03b5) is found from Table 4.2 [26]. Therefore, improved approximate of Fmax,new is then calculated as: max, ( ) r new r F F ZJ \u03b5 = , (7) Equations 1-7 are then solved with this new estimate of Fmax,new as Fmax,initial until the difference between these is less than 1%. Afterwards, the inner static load distribution is calculated based on the roller location and its number (Figure 5a) as, ( )( )max 1 ( ) 1 1 cos 2 n iF F\u03c8 \u03c8 \u03b5 = \u2212 \u2212 , (8) According to the feedback shown in Figure 3, the external vibration might be either displacement input or force input. The angle of the lateral force applied to each roller is calculated as: = 90 \u2212 (9) where \u03c80=0, \u03c8i=360i/Z, i=1,2, 3\u2026., n and Z is the roller number. In the axial direction, all rollers are affected by the same amplitude (Figure 5b). 2.1.3 Contact mechanics In the case of displacement input, the maximum tangential load (T* ) can be calculated as [28], 3 2 * * 16 1 1 3 (2 )i i aG T F F v \u03b4\u00b5 \u00b5 \u03c6 = \u2212 \u2212 \u2212 , (10) where \u03b4* is the maximum tangential displacement during oscillation (it can be used as width of the wear), and Fi is the normal load in the outer race and inner race. T* is the maximum tangential load, \u00b5 is the friction coefficient, \u028b is the Poisson ratio, G is the shear modulus and \u03d5 is the correction factor that can be calculated as [28], ( ) ( ) 1 1 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000976_s11044-019-09704-1-Figure41-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000976_s11044-019-09704-1-Figure41-1.png", "caption": "Fig. 41 Load distribution of thin-walled four-point contact ball bearing under radial, axial load and moment of 50000 N mm", "texts": [], "surrounding_texts": [ "The dynamic performance of thin-walled four-point contact ball bearing is affected by the static and dynamic loads of the wrist, hand, and workpiece or tools of the robots, especially the impact force brought by inertial force and moment under high velocity or acceleration. Considering multi-clearances and multibody contacts, the dynamic contact and impact of ball-to-ring raceway surface contacts, ball-to-cage pocket surface contacts, and cage-toouter ring guidance surface contacts of the bearing are investigated by using the multibody contact dynamic analysis models in this work. The proposed models of the bearing are solved by HHT algorithm with ADAMS software package. The results of dynamic contact forces, impact forces, displacements, and velocities are discussed under the conditions of different complex loads. Due to the effect of multi-clearances and multibody contacts, the impact forces of ball-to-cage small pocket surface contacts are larger in the non-load zone under the conditions of pure radial load or combined rotating radial load, axial load, and positive moment. The dynamic characteristics will be underestimated when ignoring the effect of multibody contact dynamics of the bearing using the empirical design method. The rules of static load distribution and the simulation model of the bearing are verified by the theoretical value. Considering pure radial load, there\u2019re always four-point contacts of ball-to-ring raceway surface contacts in the load zone. The dynamic contact forces are similar to the rule of the static load distribution. The impact forces of ball-to-cage small pockets surface contacts are large, the angular velocities of ball rotation and cage are vary- ing greatly in the non-load zone. The dynamic characteristics of the bearing are stable under the condition of combined radial and axial loads. There\u2019re four-point contacts in the load zone and two-point contacts in the non-load zone. The dynamic contact force of the main contact pair is greatly different from the secondary contact pair. The maximal contact force and the angle of load distribution of the main contact pair are both larger than those of the secondary contact pair. The dynamic contact force of the main contact pair of ball-to-ring raceway surface contacts has one peak and one valley, and is nonzero in the non-load zone. The dynamic contact force of the secondary contact pair is similar to the rules of static load distribution and is zero in the non-load zone. The angular velocities of the ball rotation and cage are stable and periodically varying as a result of the small impact forces of ball-to-cage small pockets surface contacts. The moments are important to the rules of dynamic contact forces of ball-to-ring raceway surface contacts and angular velocities of the ball rotation. The dynamic contact forces of the main contact pair in the load zone are increasing as positive moments increase. The dynamic contact forces of the secondary contact pair in the load zone are increasing as the negative moments increase. The effects are an increase in the radial load and reduction in the preload. The impact forces of ball-to-cage small pockets surface contacts in the non-load zone are large as a result of the effect of rotating radial load. The motion trajectories of outer ring center are as from circular whirling motion. Under the conditions of the proposed loads, the motion trajectories of cage center are always similar to a circular motion trajectory. As a result of the slight impact forces of ball-to-cage big pockets surface contacts, the balls in the cage big pockets are always purely rolling in the load zone and slightly varying in the non-load zone. The motion stability is high and dynamics characteristics are complex for the bearing. In the proposed work, a new approach is presented and the calculated results are illustrated. The dynamics performance and motion accuracy of thin-walled four-point contact ball bearing are complicated. They\u2019re influenced by the complex load conditions, multi-clearances and multibody contacts of ball-to-ring raceway surface contacts and ball-to-cage pockets surface contacts. The influences of the geometrical parameter, multi-clearance and flexibility on multibody contact dynamic analysis, and motion accuracy of thin-walled four-point contact ball bearing will be investigated under varying complex working conditions in the future. Acknowledgements The authors would like to express sincere thanks to the referees for their valuable suggestions. This project is supported by National Natural Science Foundation of China (grant nos. 11462008 and 11002062) and Natural Science Foundation of Yunnan Province of China (grant no. KKSA201101018). This support is gracefully acknowledged. Publisher\u2019s Note Springer Nature remains neutral with regard to jurisdictional claims in published maps and institutional affiliations." ] }, { "image_filename": "designv11_5_0002267_tpel.2020.3038741-Figure13-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002267_tpel.2020.3038741-Figure13-1.png", "caption": "Fig. 13. Current phasors at the initial rotor position (\u03b8 = \u03b8dq) and 1000 r/min. (a) Minimum power factor operating point iph1 = 21.1\u2220 19\u25e6. (b) Maximum power factor operating point iph1 = 18.9\u2220 65\u25e6.", "texts": [ " 12 shows the excitation angles corresponding to the maximum power factor at different speeds for average torque of 1, 2, and 3 N \u00b7 m. It can be seen that the maximum power factor always happens at an excitation angle between 62\u25e6 and 67\u25e6. As mentioned before, the maximum power factor is not at \u03b8dq = 60\u25e6 due to the Im 2R losses and the different current magnitude Im for the operating range that achieves the same average torque. In order to show how the location of the current phasor affects the magnetization of stator poles, Fig. 13 shows the magnetic flux path at the aligned position for the maximum and minimum power factor operating points achievingTavg = 3 N \u00b7m at 1000 r/min. It can be observed that when the current excitation angle is 65\u25e6, the current phasor is approximately aligned with the stator poles of phase c. In other words, the 65\u25e6 current phasor is magnetizing the unaligned stator poles of phase c, which are responsible for torque production. When the excitation angle is 19\u25e6, the current phasor is closer to phase a" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001062_s00202-019-00874-x-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001062_s00202-019-00874-x-Figure4-1.png", "caption": "Fig. 4 Rotor cage details for 11\u00a0kW, 400\u00a0V, eight-pole, 50\u00a0Hz motor with R = 30 rotor bars: btr= 13.08, bor= 2.0, hor= 1.0, d1 = 7.3, d2 = 4.9, hr= 11.1, a = 8.8, b = 18.2, Dshaft= 48, l = 200.6 (all dimensions in millimetres)", "texts": [], "surrounding_texts": [ "After the initial design process for an 11\u00a0kW IM with S = 48 slots and an arbitrarily chosen R = 30 rotor bars, the model parameters were obtained, as shown in Table\u00a02. While the initial design was for 30 rotor bars, by employing the PWF model, the steady-state electromagnetic torque is acquired for any desired number of rotor bars and hence, the torque ripple factor can be calculated using (4)\u2013(7). IM operation with three different arbitrarily chosen number of rotor bars 26, 33 and 40, respectively, is examined initially with the PWF model and then with FE models to cross-verify the results. A wider study of possible rotor bar numbers was subsequently undertaken in the PWF model only due to the prohibitive execution time constraints associated with the FE model analysis. An example of the PWF and FE model outputs for electromagnetic torque for the three investigated cases is shown in Figs.\u00a05, 6 and 7. The presented transients simulate the process of IM direct start-up for no-load conditions, and once steady-state is reached at t = 0.6\u00a0s, the rated load is applied. A comparison between the corresponding PWF and FE model results reveals a good level of agreement, both quantitatively and qualitatively. This is particularly apparent for the R = 33 case, shown in Fig.\u00a06. The notable high spikes in the PWF torque during startup, for the case of non-skewed rotor bars, arise due to the stepwise shape of the mutual inductance derivative curve between stator phases and rotor loops, given later in the text (Fig.\u00a016). When the rotor bars are skewed, the inductance derivative function is smooth resulting in no spikes in the PWF model predicted torque (Fig.\u00a016) [18]. The natural consequence of rotor bar skewing is a somewhat reduced value of electromagnetic torque and hence, a longer time period needed for machine acceleration. A comparison of the PWF and FE model steady-state electromagnetic torque predictions for the three examined rotor bar number cases is shown in Figs.\u00a08, 9 and 10. Again, the obtained results are seen to be in good agreement, further confirming the PWF model predictions validity. It is interesting to note that the introduction of rotor bar skew has a significant mitigating effect on the electromagnetic torque ripple for R = 26 and R = 40. However, for the case of R = 33, rotor bar skew does not significantly modify the existing ripple level and is thus, an almost unnecessary measure in this respect. Furthermore, it can be shown that this observation is valid for almost all odd rotor bar numbers. To enable an evaluation of the rotor bar number choice for the examined IM from the perspective of torque ripple minimization, a wider study was performed with the PWF model for rotor bar numbers ranging from R = 20 up to R = 73 (approximately in the range 0.5S \u2264 R \u2264 1.5S). The obtained results for the torque ripple factor are shown in Table\u00a03 for each considered number of rotor bars. Furthermore, both skewed and non-skewed rotor designs were assessed for each evaluated rotor bar number. 1 3 Table\u00a03 also provides torque ripple factors obtained from the FE model for R = 26, R = 33 and R = 40. The corresponding FE model and PWF model data are seen to be in good agreement. Furthermore, identical trends are obvious from both models. For example, both report the smallest torque ripple factor value for the case R = 33, and the largest value 1 3 1 3 1 3 for the case R = 40. For ease of interpretation, the torque ripple factors from Table\u00a03 are also graphically presented in Figs.\u00a011 and 12." ] }, { "image_filename": "designv11_5_0000390_j.protcy.2016.03.068-Figure8-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000390_j.protcy.2016.03.068-Figure8-1.png", "caption": "Fig. 8. (a) Free body diagram and (b) cross-section of curved beam in normal plane.", "texts": [ " yx, coordinates of curve 11BA are measured from the drawing at twenty equal division points and exported in a data file by using an auto-lisp code. Deflection profiles of the master leaf under each step of loading are drawn in MATLAB\u00ae computational platform from the data files and presented in Fig. 7. The figure also shows results from direct measurement, which are in good agreement. At pre-load condition, SG circuit is balanced and at each of the six load steps, individual SG readings are observed by using a portable strain indicator (item D) and given in table 2. Free body diagram (FBD) of the master leaf is shown in Fig. 8 (a) by modeling it as an initially curved beam under three point bending. The FBD is drawn in equilibrium position attained after application of load and point 1A is considered as origin of the present coordinate system xyzO . At a particular load step, distance of load line from xyzO is WL , where horizontal restraining force P develops with load W due to the geometric asymmetry of the LS. Reaction forces at left and right ends of the curved beam are LR and RR respectively coming from the roller supports", " Horizontal components of reaction forces at left and right ends are LyLxL PP tan/ and RyRxR PP tan/ respectively, which yields restraining force P at load application point. At location x ( Lx0 ) shear force and bending moment are given by LxLforWPVandLxforPV WyLxWyLx ,0, (1) LxLyHPLxWyPxP LxyPxP M WWWxLyL WxLyL x ),()( 0, (2) According to Winkler-Bach formula, bending stress developed in the beam at location x ( Lx0 ) is given by )( yrAeyM nxbx . (3) For rectangular cross-section, as shown in Fig. 8 (b), radius of curvature of neutral axis is given by )]//[ln( 12 rrhrn . The beam is not only subjected to bending but there is in plane tensile load as well. At location x ( Lx0 ) in plane load is given by xxxxLtx VPP sincos , where x is slope of the beam profile at that location. This in plane load produces axial stress which is given by bhPtxtx . (4) Presence of circular hole in master LS gives rise to stress concentration. Overall stress distributions around the circular hole in an infinite plate subjected to nominal stress , as given by [2], are )]2cos()431()1[(2/ 242 r , )]2cos()31()1[(2/ 42 and 2/r 431[( )]2sin()2 2 , where a is the radius of hole (refer Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000200_s00202-019-00882-x-Figure17-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000200_s00202-019-00882-x-Figure17-1.png", "caption": "Fig. 17 Eddy current volumetric loss distribution, nin = 1100\u00a0rpm/nout = 200\u00a0rpm", "texts": [], "surrounding_texts": [ "Distributions of the eddy current losses are presented in Figs.\u00a015, 16, 17, 18 and 19 for the NdFeB magnets [27, 28]. These losses are limited to electrically conductive parts only. Results correspond very well with these for electrical conductivities of permanent magnets, shown in Table\u00a02, and total losses and efficiency in Table\u00a03. Eddy currents are estimated at five rotational speeds, for inner rotor 150, 550, 1100, 1650 and 2200\u00a0rpm and for outer rotor rotational speeds 27, 100, 200, 300 and 400\u00a0rpm, respectively. Output rotor torque reduction due to eddy current demagnetization effect for different rotational speeds is shown in Fig.\u00a020. Eddy currents, neglecting the temperature changes of resistivity, are linearly related to rotational speed [29, 30]. Eddy current losses rise with the square of rotational speed. Also, they are influenced by the huge number of magnetic pole pairs. At rotational speeds greater than 500\u00a0rpm for the magnetic gear construction under consideration, the heat effect of eddy current losses could not be neglected. At speed 10,000\u00a0rpm, only the eddy current losses are overcoming 4% of the transmitted power. Eddy current volumetric loss distributions show that most of the losses are located in the low-speed rotor magnets. They are induced by high-speed rotor magnet movement and are amplified by modulating segments at high harmonics of the magnetic flux. Frequency separation of eddy current losses is important for loss analysis. 1 3 8 Magnetic field harmonic distortion of\u00a0CMG In many existing researches, eddy current losses are often ignored in steady state because of direct analogy with single-rotor electrical machines with permanent magnets. In MGs, these losses still appear because of relative movement of two rotors and modulating segments. Harmonic flux distortion increases eddy current effects. The main rotational frequencies of the magnetic field of the magnetic gear construction for the outer and inner rotors are 55\u00a0Hz and 10\u00a0Hz. The radial components of the flux density in the air gap between the inner rotor and the steel segments of the magnetic gear are shown in Fig.\u00a021. The radial components of the flux density in the air gap between the steel segments and the outer rotor of the magnetic gear are shown in Fig.\u00a022. The fast Fourier transform (FFT) analysis of the radial components of the flux density in the air gap between the inner rotor and the steel segments is shown in Fig.\u00a023. The FFT analysis of the radial components of the flux density in the air gap between the inner rotor and the steel segments is shown in Fig.\u00a024. 1 3 FFT analysis shows significant harmonic distortion. All frequency harmonics different from main 55/10\u00a0Hz are creating torques with no proper frequency of rotation, limiting the output CMG torque. According to these estimations, more than 20% of the total magnetic flux is engaged in these undesired frequencies [31]. This distortion effect does not influence the power efficiency, but it limits the CMG output torque and does not use the available flux from the permanent magnets. The harmonic distortion is not related to direct power losses because we are not using electrical supply for these fluxes, but it limits the output torque value of the CMG design. Some words on hysteresis loss are necessary, considering frequency-dependent hysteresis models, such as Steinmetz equation variations; they are suitable for loss superposition over magnetic flux amplitudes and frequency harmonics, like shown in Figs.\u00a023 and 24. However, a closer look reveals that hysteresis loss is not substantial for MG operation, as a power loss, covered by electrical excitation as it is in rotational machines. In MG, magnetic hysteresis loop causes time-dependent flux non-linearity, decreasing this way slightly the dynamic magnetic torque interaction between rotors. The summarized results for torque reduction and losses according to rotational velocity are presented in Table\u00a04. Losses are estimated for 150\u00a0rpm, 2200\u00a0rpm and 10,000\u00a0rpm of high-speed rotor. Results are showing significant dynamic torque reduction in high rotational speeds and rise of eddy current losses. According to estimated losses, efficiency mapping at MG overload, at torques above 320\u00a0Nm, is shown in Fig.\u00a022. At low-speed overload, efficiency is influenced by rotor slipping, while in high speeds it is influenced by eddy currents (Fig.\u00a025)." ] }, { "image_filename": "designv11_5_0000888_j.promfg.2019.06.223-Figure18-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000888_j.promfg.2019.06.223-Figure18-1.png", "caption": "Fig. 18. Digital fabrication of the multi-material fixture", "texts": [ " The color STL model of the fixture model is sliced into 250 layers, and the hatch width for internal filling of the CFs is 1mm. According to their distribution, the eight materials are assigned to the two robotic arms. To model the region-based constraint, R7 and R3 are set as the work regions for these two actuators respectively. Moreover, a 20mm safety distance is assigned to each robotic arm according to the dimensions of their end-effectors. After generating concurrent toolpaths, the multi-material fixture can be digitally fabricated and visualized, as shown in Fig. 18. The total build time is about 12.05 hours with the concurrent toolpaths, and 13.26 hours with sequential toolpaths. For some applications, a single-material fixture may be good enough and cost-effective. In this case, a monochrome STL model of the fixture can be loaded, as shown in Fig. 19. The region-based constraint, the slicing parameters and the hatch width are the same as before. The total build time of the single-material fixture is about 11.82 hours with concurrent toolpaths, and about 13" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000557_aim.2016.7576989-Figure9-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000557_aim.2016.7576989-Figure9-1.png", "caption": "Figure 9. Schematic of motion pattern", "texts": [ " When the artificial muscle is pressurized, it expands only in the radial direction because the carbon fibers do not readily expand axially. Thus, the artificial muscle contracts in the axial direction. The motion of the peristaltic pump is specified by three parameters: wavelength, propagation speed, and wave number. The wavelength is the number of adjacent expanding units. Wave propagation refers to the number of wavelengths moved during one step and wavenumber is the number of wavelengths that move simultaneously. Fig. 9 shows the motion pattern specified by two wavelengths, one wave propagation, and one wavenumber, which is expressed as 2-1-1. We refer to the regular intervals between each successive states as the motion intervals, and one cycle is the motion from the starting state to the former state. In this section, we develop the geometric theory of maximum conveyance amount for the peristaltic conveyor. The inside space volume (V0 [mm3]) is obtained from the diameter (d [mm]) and the length (lD [mm]) of the conveying direction by " ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001037_tec.2019.2951659-Figure8-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001037_tec.2019.2951659-Figure8-1.png", "caption": "Fig. 8: Geometry and parameters for the tested machine", "texts": [ " In SIR, the airgap variation, for each computational plane is fixed in space, however, in DIR fault, the airgap variation changes with the rotor position. Fig. 7 shows an example for the variations in the airgap for each computational plane. For this example a total of 10 computational planes are assumed for an airgap of 0.5mm, outer radius of 30mm, inner radius of 10mm, and 50% fault severity. III. 3-D FEA SIMULATION TESTS Simulations using 3D FEA were carried out on a threephase, 12V , 8 poles single-stator single-rotor fractional horsepower AFPMSM. Fig. 8 shows the geometry and the parameters of the tested machine. Simulations were performed under healthy conditions, first to validate the quasi-3D computation modeling. Fig. 9 and Table I show a comparison of the motor voltages (average Vd and average Vq) between the simulation results and quasi3D computation under different operating conditions. For these tests, the motor was operating at a constant speed of 2800RPM . It can be noted that the quasi-3D computation can estimate the motor voltages with high accuracy for all the studied conditions" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002086_tia.2020.3015693-Figure26-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002086_tia.2020.3015693-Figure26-1.png", "caption": "Fig. 26. Magnet flux distribution of the unskewed PM-pole after removing the applied demagnetizing d-axis current.", "texts": [ " The differential probe (700924) with a bandwidth of 100 MHz is connected to the Yokogawa (SL100) to record the back EMF. Demagnetization of the unskewed PM-pole is easier compared to the skewed PM-pole. The applied demagnetizing current pulse is shown in Fig. 24. The flux density of the unskewed PM-pole before and after applying the demagnetizing current pulse is shown in Fig. 25. The demagnetizing current pulse brings the magnetic flux density close to zero which ensures complete demagnetization of the unskewed PM-pole of the VF IPMSM. The FEA results in Fig. 26 show that both the PMs are uniformly demagnetized. Fig. 23. VF IPMSM with six probe along the length of PM. Authorized licensed use limited to: Cornell University Library. Downloaded on September 01,2020 at 03:48:51 UTC from IEEE Xplore. Restrictions apply. 0093-9994 (c) 2020 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. Each step of the 3-step skewed PM-pole in the VF IPMSM has to be demagnetized uniformly" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002002_012076-Figure9-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002002_012076-Figure9-1.png", "caption": "Figure 9. Strain on tire with air package", "texts": [], "surrounding_texts": [ "The calculated forces like bump force were applied by the road to the tire of the wheel with a force of 2500N and the analysis is carried out in the no-slip condition as shown in figure 7." ] }, { "image_filename": "designv11_5_0002269_lra.2020.3039732-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002269_lra.2020.3039732-Figure2-1.png", "caption": "Fig. 2. The spatial pose relationship between the endoscope and the instrument during surgery.", "texts": [ " The RCM mechanism can effectively prevent the nasal endoscope from pulling the tissue at the nostril, thereby helping to improve the motion safety of the endoscope during surgery. In the early stage of this letter, an active endoscope robot for nasal surgery was developed [37]. The robot has 7 DOFs, and the RCM mechanism of 4 DOFs is adopted to realize the pulling protection of the nostril during surgery. In nasal surgery, to obtain a clear surgical FOV, the spatial pose relationship between the endoscope and the instrument is generally shown in Fig. 2 (the robot in the figure was developed during the early stage of this letter [37]). Through discussion with the otolaryngologist, generally, the endoscope tip is located on the spherical surface centered on the instrument tip, the axis extension line of the endoscope passes through the sphere center, and the sphere radius is equal to the focal length of the endoscope camera R. At this time, the surgical FOV output by the endoscope is optimal. Therefore, the main task of the robot is to dynamically adjust the endoscope Authorized licensed use limited to: Auckland University of Technology", " Therefore, two kinds of tracking motion control are analyzed separately in this section and then combined with the transition process from outside the nasal cavity to inside the nasal cavity, to establish the overall framework of tracking motion control for the entire surgical process. Finally, to improve the safety of the robot motion, VFs constraints and contact force control are performed on the tracking motion process. Fig. 3 shows the schematic diagram of surgical FOV control based on tracking motion without RCM constraints. First, through the navigation system, the poses of optical markers 1 and 2 can be obtained as NDI ndi1 T and NDI ndi2 T (see Fig. 2). Then the poses of the endoscope and instrument tips in Fig. 3 can be calculated as NDI End T and NDI Ins T (see Fig. 3). Since OFocus is the camera focal point of the endoscope, the coordinate system at the focal point can be expressed in the coordinate system of the endoscope tip. NDI FocusT = NDI End T \u00b7 \u23a1 \u23a2\u23a2\u23a2\u23a3 1 0 0 0 0 1 0 0 0 0 1 R 0 0 0 1 \u23a4 \u23a5\u23a5\u23a5\u23a6 (1) We keep the orientation of End FocusT at the focal point unchanged and move its origin position OFocus to the instrument tip OIns (NDI Ins T (1 : 3, 4) is the origin point of NDI Ins T ) NDI NewFocusT = [ NDI FocusR NDI Ins T (1 : 3, 4) 0 1 ] (2) Then, we keep the orientation of the coordinate system unchanged and move the origin of the obtained coordinate system to point O1 End; then we obtain: NDI O1 End T = NDI NewFocusT \u00b7 \u23a1 \u23a2\u23a2\u23a2\u23a3 1 0 0 0 0 1 0 0 0 0 1 \u2212R 0 0 0 1 \u23a4 \u23a5\u23a5\u23a5\u23a6 (3) At this time, the Z-axis vector of NDI O1 End T and the Z-axis vector of NDI Ins T can be extracted as follows: NDIzO1 End = NDI O1 End T (1 : 3, 4) (4) NDIzIns = NDI Ins T (1 : 3, 4) (5) Based on the two extracted Z-axis vectors, the current spatial angle \u03b81 between the endoscope and the instrument and the rotation axis NDIzaxis(kx, ky, kz) from the endoscope to the instrument can be calculated as follows: \u03b81 = acos \u239b \u239d NDIzIns \u00b7 NDIzO1 End \u2016NDIzIns\u2016 \u00b7 \u2225\u2225\u2225NDIzO1 End \u2225\u2225\u2225 \u239e \u23a0 (6) NDIzaxis = NDIzIns \u00d7 NDIzO1 End \u2016NDIzIns\u2016 \u00b7 \u2225\u2225\u2225NDIzO1 End \u2225\u2225\u2225 (7) According to the space required between the surgeon\u2019s hand and the end-effector of the robot, we can set the desired tracking angle \u03b82, and then the desired endoscope pose can be obtained as: NDI O2 End T = RK(\u03b8) \u00b7 NDI O1 End T (8) RK(\u03b8) = \u23a1 \u23a2\u23a2\u23a2\u23a3 kxkxv\u03b8 + c\u03b8 kxkyv\u03b8 \u2212 kzs\u03b8 kxkzv\u03b8 + kys\u03b8 0 kxkyv\u03b8 + kzs\u03b8 kykyv\u03b8 + c\u03b8 kykzv\u03b8 \u2212 kxs\u03b8 0 kxkzv\u03b8 \u2212 kys\u03b8 kykzv\u03b8 + kxs\u03b8 kzkzv\u03b8 + c\u03b8 0 0 0 0 1 \u23a4 \u23a5\u23a5\u23a5\u23a6 (9) where \u03b8 = \u03b82 \u2212 \u03b81,c\u03b8 = cos\u03b8,s\u03b8 = sin\u03b8,v\u03b8 = 1\u2212 cos\u03b8, and RK(\u03b8) is the equivalent axis angle rotation matrix [38]", " 7), it indicates that there is a collision. Because the surface of the human nasal cavity is a thin mucosa, the robot stops moving immediately, thereby protecting the operation safety. The computational efficiency of the VFs model is mainly determined by the search radius R and the pixel spacing. Because the search range for collision detection is local and relatively small, the computational efficiency of the VFs model is relatively high. 2) Collision Monitoring Based on Contact Force Threshold: As shown in Fig. 2, a six-dimensional force sensor is installed at the robot end-effector to detect the contact force between the endoscope and the external environment. After removing the zero drift of the sensor and the gravity influence of the endoscope installed at the end of the sensor, the force sensor is used to obtain the contact force between the endoscope and the nasal cavity during surgery. Then, we control the robot motion by setting the maximum safe contact force threshold. The reader is referred to [39], [40] for contact force threshold specification" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001286_lifetech48969.2020.1570620257-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001286_lifetech48969.2020.1570620257-Figure2-1.png", "caption": "Fig. 2. A case of food feeding machine.", "texts": [ " \u2022 Arduino IDE is a program used for coding, running, and uploading the developed codes to a microcoontroller board. \u2022 Minimum OS requirement for Android: Android 4.2. \u2022 Minimum OS requirement for iOS: iOS 9.0. The system can be divided into software and hardware. Software is a programming part to control various devices and provide a user interface. Hardware is an assembly of equipment into the project. The work flow of the proposed product (pet-food feeding) is shown in Fig. 3Work flow diagramfigure.3. Firstly, we began with a drawing in a 3D design software as shown in Fig. 2A case of food feeding machinefigure.2 (a). After getting the 3D model, we, secondly, built the dispenser with a paper box as shown in Fig. 2A case of food feeding machinefigure.2 (b). Lastly, we built a final product with acrylic. Acrylic was used to make the dispenser strong, durable, and able to withstand to environment better than normal plastic and also able to withstand heat from sunlight too. The language used in developing this project is C and C++ as it is easy to understand especially writing code for microcontrollers like Arduino. The benefit of object-oriented coding is to make coding easier, concise, easy to read, easy to understand and reusable. The user interface for this proposed product has a set of interfaces as follows: \u2022 Graphical amount of daily feeding food as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000888_j.promfg.2019.06.223-Figure13-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000888_j.promfg.2019.06.223-Figure13-1.png", "caption": "Fig. 13. A composite X-Y stage AM system with two actuators for fabrication of a multi-material brooch", "texts": [ "net, C++ and OpenGL library. Using this system, an AM fabrication scenario can be built conveniently, with the required object and actuator data extracted automatically and managed hierarchically. The resulting toolpaths and subsequent digital fabrication process can be visualized and validated. The following case studies demonstrate the proposed toolpath planning approach. For ease of comparison, the deposition speeds of all materials are set to be 30 mm/s, and the zigzagstyle contour filling style is adopted. Fig. 13 shows a composite X-Y stage AM system consisting of two actuators for fabrication of a multi-material brooch. A total of five materials are assigned to the end-effectors according to their distribution in the brooch. Since the brooch 590 Yi Cai et al. / Procedia Manufacturing 34 (2019) 584\u2013593 Author name / Procedia Manufacturing 00 (2019) 000\u2013000 7 skeleton is made of orange material, it is assigned the highest deposition priority. The color STL model of the brooch is sliced into 100 layers with a hatch width of 1mm" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000718_978-1-4471-7278-9-Figure7.9-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000718_978-1-4471-7278-9-Figure7.9-1.png", "caption": "Fig. 7.9 The definition of a Chain Code (top left). A saw-toothed flight path caused by mixing reports from two different flights (top right), the curvature histogram (bottom left) and the bending energy histogram (bottom right)", "texts": [ " For example, the momentum of a running animal prevents it from making a sharp turn at a high speed. A flying airplane normally does not make a sharp zigzag path at a high speed. Our instinct enables us to spot such anomalous movements. Here we will apply the instinctual curvature model to filter out noise along a physical object\u2019s path, for example, the trajectory of an aircraft. In aviation datasets, it is possible that two different flight reports are mixed together due to the proximity of aircraft in time and space. Figure 7.9 shows a saw-toothed flight path caused by mixing reports from two different flights that are overlapped in time. Intuitively, we know that it is impossible for an aircraft to fly like that.14 How do we then filter out 14Eddy FW and Oue S (1995) Dynamic three-dimensional display of U.S. Air Traffic. Journal of Computational and Graphical Statitics, Vol.4, No.4, pp. 261\u2013280. 1995 Instinctive Computing 135 the points that are not on the true path? We can do this by calculating the curvature and the bend energy. There are many methods for calculating curvature. For discrete segments of a curve, we can apply the simple method called Chain Code. It encodes the unit line segments into numbers. For example, we can assign 1 of 8 directions to a unit line segment: north (0), northeast (1), east (2), southeast (3), south (4), southwest (5), west (6), and northwest (7). The Chain Code for the flight path in Fig. 7.9 would be: : : : 2; 1; 1; 3; 3; 2 : : : The curvature of two adjacent segments is the difference of their chain code at the location a, b, d, f, e. : : : 1; 0; 2; 0; 1 : : : The bending energy (BE) of a curve is the sum of squares of the curvatures over the curve length, similar to the energy necessary to bend a physical rod. The bending energy: BE D 1 L LX kD1 c2.k/ (7.2) where L is the curve length and c is the curvature of the segments. Assume the total curve length is 6. We have the bending energy values at the location a, b, d, f, and e: : : : 0:167; 0; 0:667; 0; 0:167 : : : Now, we can locate the sharp bending locations" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000327_j.mechatronics.2015.06.014-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000327_j.mechatronics.2015.06.014-Figure1-1.png", "caption": "Fig. 1. Active over-constraint parallel manipulator 6PUS\u2013UPU.", "texts": [ " If the force screw is a pure force vector, element M can be described as M \u00bc r F , where r represents the application point\u2019s position vector of F. And if the force screw is a pure torque vector, F \u00bc 0 0 0\u00bd T and M represent its direction and magnitude. A unit force screw can be defined as $\u0302 when F is a unit force vector. So a force screw can be written as f $\u0302 where f represents the screw\u2019s magnitude. The resultant force screw is the vector sum of screws if there are a set of force screws applied on the same object. $F \u00bc $1 \u00fe $2 \u00fe \u00fe $n \u00bc f 1 $\u03021 \u00fe f 2 $\u03022 \u00fe \u00fe f n $\u0302n \u00f02\u00de An active over-constraint parallel manipulator, as shown in Fig. 1, contains a moving platform with six actuated limbs and one constraint limb. The actuated limbs interlink the moving platform to the base frame with spherical and prismatic joints, and the constraint limb with both universal joints. Each actuated limb is composed of a slider and a link rod of fixed length, which are connected to each other by universal joint. The sliders are driven by AC motors via linear ball screws which are all vertical to the ground. The spatial movement of the moving platform can be described with six parameters of q \u00bc x; y; z;a; b; c\u00bd T " ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002450_j.compstruct.2021.113608-Figure17-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002450_j.compstruct.2021.113608-Figure17-1.png", "caption": "Fig. 17. Contact model of compaction roller and concave surface.", "texts": [ " The silicone rubber layer of the compaction roller was treated as a linear elastomer over a range of laying forces. It is therefore assumed that the effective contact width of the compaction roller increases with its maximum deformation. When the compaction roller is in contact with the concave surface, whether the compaction roller can fully fit the surface is the focus of our concern, which is the compaction mentioned above. The contact model of compaction roller and concave surface is shown in Fig. 17. It can be seen from the figure that the most difficult position to compact for concave surface is in the middle of the compaction roller, while the compaction situation at this position is the most concerned position for fiber placement. The width of the roller determines the maximum distance \u0394l0 between the roller and the contact surface, and the effective compaction between the roller and the contact surface depends on the relationship between the maximum deformation of the roller and the maximum distance above" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000200_s00202-019-00882-x-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000200_s00202-019-00882-x-Figure1-1.png", "caption": "Fig. 1 Power loss flow diagram for CMG Fig. 2 Magnetic flux reduction in CMG", "texts": [ " The paper is organized as follows: First, a coaxial MG design under testing is presented; consideration of the losses is made; a description of the developed computational model is given; consideration of mechanical losses, magnetic reluctance, eddy current losses, dynamic torque reduction, demagnetization is made; and before the conclusion, the test verification results are finally shown. Typically, the coaxial MG (CMG) consists of three rotors. The inner and outer rotors come with permanent magnets mounted on them. The third rotor is built of steel segments, which are positioned between the two other permanent magnet rotors, and used for magnetic flux modulation. Power losses in the CMG (Fig.\u00a01) can be classified into three main groups: mechanical losses in bearing supports and air drag; eddy current power losses in all electrically conductive materials; and magnetic flux reluctance, eddy current demagnetization and harmonic distortion effects. In fact, the last three effects do not directly divert transmitted power, but limit the output torque of the CMG downwards. All three groups of losses are frequency-dependent and significantly increase at high rotational speeds. The main losses involve eddy currents in conductive materials" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002426_j.engfailanal.2020.105211-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002426_j.engfailanal.2020.105211-Figure5-1.png", "caption": "Fig. 5. TRB stress diagram.", "texts": [ " Engineering Failure Analysis 122 (2021) 105211 { Fa1 = Fa + S2 \u2212 S1 Fa2 = S2 (3) When Fa + S2 < S1, then { Fa1 = S1 Fa2 = S1 \u2212 Fa (4) The radial force creates an additional bending moment M on bearing 1 M = Fra (5) When the load Fa direction is changed, the bearing number 1 and 2 should be replaced. The above formula can be used to calculate the internal load of the bearing. Fr, Fa are the radial and axial forces received by the hub unit. a is the distance between the load center and the center of bearing 1. When the TRB bears radial load Fr and axial load Fa, there will be a radial displacement \u03b4r and a axial displacement \u03b4a correspondingly. The bearing\u2019s stress state and deformation are shown in Fig. 5. Q is the contact load on the roller. \u03b1v is the contact angle between the roller and the outer raceway. As shown in Fig. 6, when the bearing has a radial displacement \u03b4r, for the roller at position Angle \u03d5i, the radial displacement \u03b4ri generated by the bearing outer ring relative to the inner ring is: \u03b4ri = \u03b4rcos\u03d5i (6) When the bearing bears axial load, the axial deformation component of all rollers is the same, that is, the axial displacement of rollers at different angle positions \u03b4ai is equal to the axial displacement of the bearing \u03b4a L" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002450_j.compstruct.2021.113608-Figure28-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002450_j.compstruct.2021.113608-Figure28-1.png", "caption": "Fig. 28. The mapping of the normal vector of the part surface.", "texts": [ " Due to our limited conditions at present, we cannot realize the choice of different AFP machine, so we can only analyze the accessibility of gantry AFP machine shown in Fig. 20. First, the three\u2010dimensional space of the part surface was obtained as shown in Fig. 26, which was calculated according to the model of the mould surface. The comparison diagram of the position space of the AFP machine and the three\u2010dimensional space of the part was shown in Fig. 27. Next, the mapping of the normal vector of the part surface on the Gaussian sphere was obtained as shown in Fig. 28. Comparative analysis of attitude reachable space mapping and normal vector distribution range mapping was shown in Fig. 29. According to the analysis results in Figs. 27 and 29, the position and attitude of the gantry AFP machine can be reached for this part. Based on the above analysis, appropriate roller specifications are CR64_20, CR64_40 and CR64_60. Considering the laying efficiency, select CR64_60 from the three specifications of the compaction roller. According to the compaction roller specification and prepreg tow width (1/4 in" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001019_j.ymssp.2019.106426-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001019_j.ymssp.2019.106426-Figure4-1.png", "caption": "Fig. 4. The schematic of aero-engine ground test system; A, B, C indicate the locations of the sensors.", "texts": [ " Basic idea of the proposed collaborative sparse classification. sively in Section 4. In Section 5, numerous experiments are performed to verify the superiority and robustness of the proposed method with the start-of-arts. Conclusion and future work of the research are presented in Section 6. Due to the requirements of the aero-engine safety and reliability, a ground test system of the whole aero-engine is constructed to reveal the vibration mechanism and perform the diagnosis of the gear-hub, which is shown in Fig. 4. In the ground test, an aero-engine with a cracked gear-hub and the same engine with a normal gear-hub were tested for comparison. The gear-hub have total 97 splines. The location and the length of the preset crack on the fault gear-hub are shown in Fig. 5. The three cracks were generated through line cutting process, one is a line crack with a length of 16 mm, and another two are \u2018\u2018L\u201d-shaped cracks with the radius of 9 mm and 16 mm respectively. To obtain the vibration signal of the aero-engine, three accelerometers labeled as \u2018\u2018A\u201d, \u2018\u2018B\u201d and \u2018\u2018C\u201d were mounted on the engine casing, shown in Fig. 4, and their detailed locations are also summarized in Table 1. In the test, the sampling rate was set as 20 kHz. The trial run of the aero-engine experienced eight stable working conditions, which are symbolic of \u2018\u2018slow\u201d, \u2018\u20180.2\u201d, \u2018\u20180.4\u201d, \u2018\u20180.6\u201d, \u2018\u20180.85\u201d, \u2018\u20181\u201d and \u2018\u2018fly\u201d respectively. The \u2018\u2018slow\u201d working condition corresponds to the minimum speed that can maintain the stationary operation of the engine. The \u2018\u2018fly\u201d working condition corresponds to the maximum speed and maximum power of the engine. \u2018\u20181\u201d represents the standard working condition, where the engine works under the rated power and speed" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001946_tte.2020.2997730-Figure18-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001946_tte.2020.2997730-Figure18-1.png", "caption": "Fig. 18. HE machines with consequent pole rotor[44]. (a) Inset-PM HE machine. (b) IPM HE machine. (c) 2IPM HE machine. (d) V-type IPM HE machine. (e) Spoke-IPM HE machine.", "texts": [ " According to the derivation method, the conventional HE topologies for any PM machines can be obtained, which is a combination of the PM topology and HE topology of the same original machine. Authorized licensed use limited to: University of Exeter. Downloaded on June 09,2020 at 19:59:50 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. Besides, the HE machines based on different original machines (HEMDO) are investigated, as seen in Fig.17~Fig.18. Except the stator-PMs in Fig.17(b), the other PMs in Fig.17~Fig.18 produces bipolar flux in the armature winding, thus achieving a relatively high torque density. For each HEFS topology in Table II, the flux-linkages of the armature winding produced by the PMs and the field windings are bipolar, while for the hybrid excitation \u201cE-core\u201d flux-switching machine (Fig.17(e)), the flux-linkages of the armature windings produced by the field windings are unipolar, hence, it can be regarded as a HE machine based on different original machines (HEMDO), where the HEMDO is combined of variable flux reluctance machine (VFRM) and FSPM machine" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002450_j.compstruct.2021.113608-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002450_j.compstruct.2021.113608-Figure3-1.png", "caption": "Fig. 3. Relationship between deformation of prepreg tow and steering radius.", "texts": [ " Finally, the consistency of the two methods was compared, and how to determine the width of prepreg tow by placement suitability analysis was introduced. The minimum turning radius is one of the important characteristics of prepreg tow, which is usually obtained by experimental method and formula derivation [26\u201329]. The focus of this paper is not to study the deformation law of prepreg tow, so the minimum turning radius of the tow is assumed to be the intrinsic property of the tow as input. As shown in Fig. 3, R represents the turning radius of the center line of the tow, w represents the width of the tow and\u025bdenotes the capacity of deformation on the unit length of the tow. It is assumed that the middle layer of prepreg tow is the neutral layer, the outside is seen as under tension and the inside is under compression. The derivation process of the deformation limit per unit length of the prepreg tow is as follows: dlcenter \u00bc R d\u03b8 \u00f01\u00de dlin \u00bc R w=2\u00f0 \u00de d\u03b8 \u00f02\u00de \u025b \u00bc \u0394dlj j dlcenter \u00bc dlcenter dlinj j dlcenter \u00f03\u00de Combining Eqs" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002716_j.ast.2021.106899-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002716_j.ast.2021.106899-Figure6-1.png", "caption": "Fig. 6. Experimental Setup.", "texts": [ " As the stabilization of the quadrotor is a necessary condition to perform the experimental testing, the X and Y position and angles control uses cascading PID control to adjust the torques. For these components, the sensory information was provided by the accelerometer, the ultrasound sonar, the gyroscope, and the magnetometer. Since the tests were performed indoors, the X and Y position measurements are provided by an indoor multiple-camera motion capture system. The experimental setup is presented in Fig. 6. The load is added before the flight. In this section, a simple verification of the system stability is provided. The design mass is 0.445 kg, and the real mass is 0.47 kg. The resulting eigenvalues for the height control are {\u22121.4648 \u00b1 3.0317i, \u22120.7023 \u00b1 0.4351i, \u22123.5159, \u22120.1077}. As intended, the real component of the eigenvalues are negative, providing an overall closed loop system that is stable. The design inertia is 5.1 \u00d7 10\u22123 kg m2, and the real inertia is 4.8 \u00d7 10\u22123 kg m2. The resulting eigenvalues for the yaw control are {\u22120" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000953_coase.2019.8843316-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000953_coase.2019.8843316-Figure3-1.png", "caption": "Figure 3. Build for surface roughness study [11]", "texts": [ " The knowledge data feature types was queried from an AM knowledge base using on an AM ontology. The AM ontology is the main body of + /< /+= >?/ AM knowledge base developed with formal Web Ontology Language (OWL)/Resource Description Framework (RDF) representations. The ontology was developed to capture a priori knowledge of experimental reports or benchmark studies on AM [20, 21]. This study uses AM data obtained from a Laser Power Bed Fusion (LPBF) build with 9 identical parts, each of which has 72 surfaces oriented differently in either XY direction or Z direction, as shown in Figure 3 [11]. The 9 parts were built on 3 by 3 grids at the same z level 50mm from each other in both x and y directions. Each part has 8 ribs at increments of 45 degrees in XY plane. Each rib includes 9 surfaces oriented from 45 degrees to 165 degrees against Z axis at 15 degrees of increments. Measurements were conducted to evaluate the surface roughness for all the 648 full factorial designed surfaces. The surface height is measured based on ISO 4287 standards. TABLE I summarizes the datasets used for this study [11]" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002543_j.fusengdes.2021.112309-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002543_j.fusengdes.2021.112309-Figure6-1.png", "caption": "Fig. 6. Non-planar spherical surface segment configuration.", "texts": [ " Fusion Engineering and Design 167 (2021) 112309 and rinsing of the filler (Fig. 1, step 5) has been completed in late 2020 as shown in Fig. 5. The integral process chain is planned to be demonstrated in 2021 by completing the fabrication sequences and procurement as well as joining of a corresponding cover plate according to steps 6) \u2013 9) illustrated in Fig. 1. In late 2020 another non-planar configuration demonstrator has been developed and designed in collaboration with Hermle representing a spherical surface segment with a meandering channel flow path, see Fig. 6. The CS deposition will be done in early 2021, as well as the procurement and joining of the cover plate to demonstrate the integral manufacturing process chain. Another novelty in relation to the CS based process chain demonstrated firstly in 2020 is the realization of semi-detached rib structures [3,4] inside of cooling channel surfaces. The semi-detached ribs are positioned onto riveting pins machined on the bottom side (heat flux facing surface during operation) of the cooling channel and fixed by plastic deformation of the rivet pin" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001289_tmech.2020.2992711-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001289_tmech.2020.2992711-Figure2-1.png", "caption": "Fig. 2. Illustrative schematic for design.", "texts": [ " It is noteworthy that deployment of FLAM actuators to all four tensioned legs allows platform stabilization control against wind and wave of virtually any directions. Two FLAM actuators are deployed at each tendon in order to avoid excessively large actuator dimension. Under the current design, the actuator diameter is 0.569 m. If only one actuator were deployed at each tendon, the actuator diameter would be even larger, which might not be economically viable. As mentioned in [34, 36], an artificial muscle shown in Fig. 2 (b) can be made by twisting a high-strength polymer fiber, such as nylon 6 or nylon 66 into a coil. When heated, the twisted fiber untwists and generates a torsional actuation, which drives the coil to contract. Heating can be achieved by submersion in hot water, hot air or application of electric current through certain fashion of conductive coating deployed on the fiber surface. As the FOWT platform control features oscillatory cycles, the FLAM actuators for each pair of tendons are expected to contract (heating) and relax (cooling) in turn. For FOWT control, by far it is more realistic to use joule heating via electric current. Copper wires wrapped around the nylon fibers, as shown in Fig. 2(a), are cost effective to realize joule heating. In contrast, the cooling operation is a primary challenge for achieving desirable transient performance of FLAM based actuation. For the FOWT operation, sea water provides a natural cooling medium. In this application, artificial muscles are assumed to be cooled by sea water drawn by pump attached to the TLP spoke. For the heating process, to reduce heat loss, Authorized licensed use limited to: Auckland University of Technology. Downloaded on May 29,2020 at 00:10:36 UTC from IEEE Xplore. Restrictions apply. 1083-4435 (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. artificial muscles are heated by electric current in air. For applications that need significant level of actuation, e.g. required by AMLFC of FOWT, multi-strand bundles are fabricated, as illustrated in Fig. 2(c). The FLAM actuator consists of multiple bundles, as shown in Fig. 2(d), connected to the mooring line and tendon via flanges. On the flanges, an array of holes are designed as channel for cooling water to pass between the FLAM bundles. In the following design and simulation of the FLAM actuator, sea water is assumed as 30\u00baC, which is the upper bound for North America area on monthly average base [37]. For the FLAM actuator, cooling water temperature is not necessarily constant. Lower temperature will accelerate the cooling process, thus improve the actuator\u2019s dynamics" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002614_j.compositesa.2021.106449-Figure18-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002614_j.compositesa.2021.106449-Figure18-1.png", "caption": "Fig. 18. The illustration of the plane cut of FE mesh (a) 900, (b) -450.", "texts": [ "1) or some specific requirements, such as a cutout. For FE analysis, the TWM needs to be trimmed as required. A plane and hole cut is used in this section to demonstrate the trimming procedure of TWM. The planar cutting is achieved through a submodule in the trimesh library named \u201cintersections\u201d. This module can detect the intersections between the cutting plane and the FE mesh, and automatically generate new mesh based on these intersections. To clearly illustrate this procedure, a \u2212 900 and \u2212 450 ply cut are presented in Fig. 18. Four plane cut are performed in sequence to obtain perfect straight edges, which allows the specific boundary conditions to be applied in FE analysis. The hole cut, however, requires a mesh boolean process. This can be done in the FE package or 3D computer graphics software, such as Blender or OpenSCAD, which is used for backends mesh boolean computation. The mesh file of a cylinder is used as a hole cutter to trim the geometries. The open-hole geometry of 900 and -450 are used for demonstration purpose as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001717_jmes_jour_1959_001_016_02-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001717_jmes_jour_1959_001_016_02-Figure3-1.png", "caption": "Fig. 3 Turoidal shells", "texts": [ " Wissler (IS), in a thesis which does not seem to have been published elsewhere, examined several toroidal problems, and Stange (IZ), in a paper on corrugated membranest, gives a comparison between Wissler\u2019s theory and his own for some loading cases. A more recent paper by Clark (2) also gives a comparison between the stresses in a part torus subjected to a ring of edge moments as calculated by Wissler and by Clark himself. Clark solves his equations by using a method of asymptotic integration which is valid when the parameter p is large (for example, p > 7). Such large values of this parameter often imply that the radius of the ring crosssection is not small compared to the radius of revolution (that is, Fig. 3; b is of the same order as a) but for the reverse case of interest here, where b Q a the parameter p may well be small and Clark\u2019s method is not then valid. Clark\u2019s theory was applied mainly to the compression of corrugated pipe and omega-type expansion bellows. Dahl (3), de Lieris (6), Turner and Ford (14), and Salzmann (11), have all examined, mainly by strain energy methods, certain aspects of various expansion bellows shapes, including toroidal elements, without studying the torus per se. Internal pressure loading of a complete torus has been examined by Dean (4)", " The normal flat plate, stretched disc, sphere and cone solutions can easily be obtained from equations (14) and (15). The cylinder equations can be obtained by using equation (4b) and thence reworking the solution. Skce toroidal elements may well adjoin such other structural elements it is convenient to establish these already known solutions in terms of the variables used here, and such solutions are outlined in Appendix I. TOROIDAL SHELLS The toroidal shell can be defined by r = a+b cos 0 as in Fig. 3 4 or by r = a-b cos 0 as in Fig. 3b. JOURNAL MECHANICAL E N G I N E E R I N G SCIENCB Vol I No 2 1959 at UNIV OF VIRGINIA on June 5, 2016jms.sagepub.comDownloaded from 116 C. E. TURNER Either form may be the more convenient for a given problem and both may be covered by writing r = a+ bk cos 0, where k = f l for the first or second cask respectively. For problems such as the expansion bellows or a pressure vessel as mainly considered here, the body force due to the weight of the toroidal element itself is not often important. Thus neglecting the body force Y, writing a/b = r ) ; supposing the pressure p is constant, and the body force X is proportional to radius (or of course zero); letting Fr = pp(q+k cos 6)-1/2; 2hX P +- (3+v)b2 sin B(T+k cos e)1/2 " ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001469_tmag.2020.3019821-Figure13-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001469_tmag.2020.3019821-Figure13-1.png", "caption": "Fig. 13. Prototype of DF-N/Fe/N/Fe. (a) Stator. (b) Rotor. (c) Lamination. (d) Assembled prototype with slip ring.", "texts": [ " The efficiency of four machines considering both copper loss and core loss are investigated, as shown in TABLE 2. It can be seen that DF-NSNS and DF- N/Fe/N/Fe have higher core loss than their one armature counterparts due to higher saturation. But the core loss of four machines are quite small compared to the copper loss. Overall, DF- N/Fe/N/Fe enjoys highest efficiency owing to its large torque density. To validate foregoing analysis of torque performance, the DF-N/Fe/N/Fe prototype machine has been fabricated and its major parameters are listed in TABLE 3. As shown in Fig. 13, the prototype machine has two sets of armature windings and a slip ring attached to its shaft. Authorized licensed use limited to: Carleton University. Downloaded on September 21,2020 at 00:08:11 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 comparison of FE predicted and measured back EMF waveforms at rated speed of both stator and rotor armature windings is presented in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0003154_s0368393100103104-Figure19-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0003154_s0368393100103104-Figure19-1.png", "caption": "Fig. 19. Determination of hx-distribution in a multi-cell tube.", "texts": [ " Assuming the section to be cut at point / the function hxo for the resulting open section is statically determined since the origin of the integral (275) is at the point /. The distribution hx of the actual tube is given by, hx = hx0 + hxi . . . . . . (286) where hxi is the unknown value of hx at point /. Using equation (286) in (282), hxi = L \u00a7 hxo pBds (287) which with equation (286) determines completely the function hx. A similar analysis holds for h-i. For a multi-cell tube of the type illustrated in Fig. 19 each cell is supposed cut at points I, II, etc., in the outer walls of each cell I, II, etc. The hxo system for the resulting open tube may easily be obtained by integration from each cut with due attention to the sign of s. The positive direction of s may be assigned arbitrarily in each wall but once assigned must be adhered to strictly. In Fig. 19 the positive directions of s have been denoted by arrows. The closed iV-cell tube will have iV unknowns hxi, Axii hxn at the cuts, /, //, N. Now there are JV- 1 conditions for the equality of twist of each cell and the one condition given by equation (282), in all N conditions for the N unknowns. The hx distribution 768 GENERAL THEORY OF CYLINDRICAL AND CONICAL TUBES in the actual system can now be described as follows: In the outer walls of the Mth cell, hx = hxo+hxu in the left hand wall of the Mth cell, hx=hxo + hxm-i) - hKMand in the right hand wall of the Mth cell, hx=hxo + hxM-hx vdx, b = F \u2217d /nb takes effect to provide braking force for the tractor; when the intermediate control input Fd > 0 and vx < vdx, p = F \u2217d (V T\u0398)\u22121 comes into effect to provide a propelling force for the tractor; at the rest, a small constant p will be applied to the tractor. This switching law takes into account the fact that the throttle response is much slower than the brake response, thereby avoiding undesired switching chattering. The preview lateral deviation le and heading error \u03b8e are needed for lateral control [10]. Figure 4 shows their positive configurations, and the corresponding dynamics are given as l\u0307e = vx\u03b8e \u2212 vy \u2212 uD, \u03b8\u0307e = vx\u03ba\u2212 u, (32) where \u03ba is the reference path curvature, and D=2.5m is a constant preview distance. Also, from a practical perspective, the tractor\u2019s mechanical limits make the small-angle assumption reasonable. Therefore, from (3)-(6), (9) and (23), the dynamic equations can be rewritten as v\u0307y = c1 + d1\u03b4 + Zy/m, u\u0307 = c2 + d2\u03b4 + Zu/Jz, (33) where c1= ( \u03b7rC\u0304rl2 \u2212 \u03b7f C\u0304f l1 mvx \u2212 vx ) u\u2212\u03b7f C\u0304f + \u03b7rC\u0304r mvx vy+ F yh m , c2=\u2212\u03b7f C\u0304f l 2 1+\u03b7rC\u0304rl 2 2 Jzvx u\u2212\u03b7f C\u0304f l1\u2212\u03b7rC\u0304rl2 Jzvx vy\u2212 l2 + l3 Jz F yh , d1= \u03b7f C\u0304f m , d2 = \u03b7f C\u0304f l1 Jz , and d1, d2 are positive system parameters" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001072_j.compeleceng.2019.106507-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001072_j.compeleceng.2019.106507-Figure1-1.png", "caption": "Fig. 1. Quadrotor structure, forces, angles and frames .", "texts": [ " Section 4 studies the boundedness of the proposed TLUC. Section 5 includes ERSM control design and Lyapunov stability analysis for the whole system. Section 6 demonstrates the simulations with and without applying the TLUC. Section 7 demonstrates experimental results and analysis. The conclusion is in Section 8 . 2. Quadrotor model The quadrotor dynamics built and used by many researchers [18,19] are based on the work of Lagrange and Newton\u2013 Euler . Quadrotor configurations, frames, and forces are shown in Fig. 1 . The nonlinear quadrotor system is described as [20] : X\u0308 ( t ) = F T ( \u02d9 X ( t ) ) + G T ( X ( t ) ) U ( t ) + D ( t ) (1) Where X(t) = [ \u03d5(t), \u03b8 (t), \u03c8(t), x(t), y(t), z(t)] T is the state vector, and \u02d9 X ( t) , X\u0308 ( t) are velocity and acceleration vectors, respectively. F T ( \u0307 X (t) ) and G T (X(t)) are the total nonlinear dynamics of the quadrotor system which include known and unknown dynamics, D(t) denotes the external disturbance vector. The dynamic system in (1) can be rewritten by describing the nominal part and the uncertain/changing dynamic part as: X\u0308 ( t ) = F ( \u02d9 X ( t ) ) + F ( \u02d9 X ( t ) ) + ( G ( X ( t ) ) + G ( X ( t ) ) ) U ( t ) + D ( t ) (2) Where F( \u0307 X (t)) , G ( X( t) ) are the nominal dynamics and they are given as: F ( \u02d9 X (t) ) = \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 i y \u2212i z i x \u02d9 \u03b8 \u02d9 \u03c8 \u2212 j r i x \u02d9 \u03b8 \u03c9 r i z \u2212i x i y \u02d9 \u03c6 \u02d9 \u03c8 + j r i y \u02d9 \u03c6\u03c9 r i x \u2212i y i z \u02d9 \u03b8 \u02d9 \u03c6 0 0 g \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 G(X(t)) = diag [ 1 i x , 1 i y , 1 i z , 1 m , 1 m , \u22121 m cos \u03c6 cos \u03b8 ] (3) Functions F ( \u0307 X (t) ) and G(X(t)) are the uncertain terms of the dynamics" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000685_1077546318818694-Figure17-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000685_1077546318818694-Figure17-1.png", "caption": "Figure 17. The failure mechanism of the bearing at the drive end of generator: (a) initial pitting; (b) deep concaves; (c) bend of the cage; and (d) damage.", "texts": [ " The seriously damaged bearing is shown in Figure 16(a), and its disassembled rolling elements, inner race, and cage are shown in Figure 16(b), Figure 16(c), and Figure 16(d). Actually, all the fault characteristics of the tested bearing were extracted by PEWT with margin factor during the vibration test 24 days earlier. If appropriate precautions had been implemented at that time, the catastrophic results would have been avoided. Through observing the failure pictures of the tested bearing, the failure mechanism can be deduced as shown in Figure 17: first, due to insufficient lubrication, the inner race in heavy load area is prone to pitting, making the rolling elements rough, as shown in Figure 17(a). Then, continuous operation of the wind turbine aggravates the above defects, leading to possible deep concaves on inner race, as shown in Figure 17(b). Next, the deep concaves limit the revolution of the rolling elements, which generates a large bending moment to the cage, as shown in Figure 17(c). Finally, the bending moment destroys the cage and some of the rolling elements fall into the bottom of the bearing, as shown in Figure 17(d). At that time, the bearing is dead, and the shaft is not turning. On the basis of the failure mechanism, selecting an excellent lubricating grease and duly checking the state of lubrication of the bearing would be a promising measure for preventing the bearing failure in wind turbine generators. Meanwhile, vibration analysis is necessary for knowing the health state of the bearing. In the EM with maximum margin factor shown in Figure 15, all the bearing fault features are successfully found. To compare with the sensitivity of margin factor to bearing fault, all decomposed EMs are sorted according to kurtosis, and the EM with maximum kurtosis is shown in Figure 18" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000327_j.mechatronics.2015.06.014-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000327_j.mechatronics.2015.06.014-Figure2-1.png", "caption": "Fig. 2. Force analysis of actuated limb.", "texts": [ " Time differentiation yields kinematical mapping relationship between dependent and independent parameters, with result of _q \u00bc G _qs; G 2 R6 5 \u00f03\u00de \u20acq \u00bc G \u20acqs \u00fe _G _qs; _G 2 R6 5 \u00f04\u00de Dynamics of the redundantly actuated parallel manipulator includes forward dynamics and inverse dynamics. Forward dynamics refers to the solution of corresponding movements when driving forces/torques are given, while inverse dynamics refers to the desired driving forces/torques when motion state of moving platform is given. Ignoring the influence of friction, dynamic model of each limb can be achieved by virtual work principle. As shown in Fig. 2, the actuated limb i bears gravity mHi g applied on the slider, mLi g applied on the rod and inertia force and moment mHi aHi applied on the slider, mLi aLi applied on the rod. In addition, it bears contact force f i applied on the rod from the moving platform and driving force sHi applied on the slider from the linear ball screw as well. For the PUS actuated limb i, the contact force f i applied from the moving platform produces not only the limb\u2019s movement but also its deformation. Assuming that the little deformation have no effect on the movement of parallel manipulator, the actuated limb can be seen as rigid body and its dynamic model can be obtained by rigid body dynamics" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002450_j.compstruct.2021.113608-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002450_j.compstruct.2021.113608-Figure5-1.png", "caption": "Fig. 5. Relationship among the deformation capacity of tow, the width of the tow and the Gaussian curvature.", "texts": [ " In addition, automated fiber placement technology is generally used for non\u2010developable surfaces. In view of this, this paper intends to approximate a sufficiently small neighborhood of any point on the curved surface as a part of a spherical surface. And then, the suitability problem between the prepreg tow and the surface to be laid is greatly simplified. The number of calculations for suitability analysis at any point on the surface is greatly reduced, and the Gaussian curvature of the surface can be better introduced into the suitability analysis. As shown in Fig. 5, take any longitude (Pf ) of the spherical surface as a fiber placement path, and take a plane that passes through the center of the sphere and is perpendicular to the path Pf to cut the spherical surface. The direction of the placement path is perpendicular to the paper direction. The width of the prepreg tow is w, the radius of the spherical surface is r, and the angle of the prepreg tow width on the spherical surface is \u03c6. It is assumed that the length along the placement direction is dl, and the corresponding angle formed is d\u03d5" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001942_acsabm.0c00446-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001942_acsabm.0c00446-Figure1-1.png", "caption": "Figure 1. Schematic of biohybrid DSSC and suggested redox pathway.", "texts": [ " A gel analogue of a DSSC with an I-/I3 - mediator exhibited improved photocurrent stability after 125 days of operation versus an equivalent liquid cell.29 These advantages allow for the minimization of evaporation and enhanced portability while still harnessing MET. Device Design. The goal for this investigation is to design a low-cost, renewable photovoltaic device that integrates a DSSC electrode with a gel-based medium and a PSI-coated cathode to elevate the performance of the cell, as measured via the short circuit current, open circuit photovoltage, and power output. Figure 1 shows a schematic of the gel-based biohybrid solar cell. The device utilizes an FTO-coated glass slide as a transparent current collector upon which the TiO2 film is spincoated, annealed, and loaded with anthocyanin dye. The gel media are 0.5% wt agarose hydrogel with a supporting electrolyte and an AscH/DCPIP mediator. Agarose is a naturally derived polysaccharide that is a principal component of agar, a gel commonly used for gel electrophoresis. Agarose has been used in DSSCs and has shown performance enhancement over equivalent liquid cells" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002245_1350650120969003-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002245_1350650120969003-Figure4-1.png", "caption": "Figure 4. Test rig of novel GFJB. (a) Structure of air compressor; (b) fabricated rotor and bearing.", "texts": [ " The arch height of the upper bump foil is lower than that of the inserted shim foil when Dhb6\u00bc0, which means an initial installation clearance and preload effect for the novel GFJB. So there is a lower clearance at the location of double arches, except for the enhanced structural stiffness. Figure 3 displays the processed GFJB and the mould used to form the bump foil and shim foil with different Dhb. The arch number on the bump foil is 15, and the three groups of arches are evenly spaced on the shim foil. The thickness of shim foil and bump foil are ts\u00bc 0.08mm and tb\u00bc 0.1mm, respectively. And the other arch parameters of these two foils are ensured consistent. Figure 4(a) shows an air compressor supported by the oil-free foil bearings, which is used in a fuel cell vehicle. The T-rotor is supported by two GFJBs and two GFTBs, and the main parameters are shown in Table 1. The air compressor is electrically driven, and there is a Sieb-Meyer controller for the speed control. The compressor\u2019s impeller is removed to make it easier to observe the rotor through a laser displacement sensor, which is placed at the shaft end. To verify the effect of bump-type shim foil on the supporting performance, the novel and traditional GFJBs are manufactured and used, as shown in Figure 4(b). It should be noted that during the test, all parts of the air compressor, including GFTBs, remains consistent except for the replacement of GFJBs. To eliminate the impact of inserting shim foil on the nominal clearance, the cartridge diameter of the novel GFJB is increased by 0.2mm, which exceeds the inserted thickness as it\u2019s impossible to have the complete contact for the two foils. The rotor tracks at different rotating speeds is recorded. For the isothermal and isoviscous gas, the Reynolds equation for GFJB is @ @z ph3 12gg @p @z " ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002656_tmech.2021.3082935-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002656_tmech.2021.3082935-Figure7-1.png", "caption": "Fig. 7. The experiment platform of the SbW system. (a) Schematic diagram of the experiment platform, (b) Physical diagram of the experiment platform.", "texts": [ " 5, one can find that both the designed adaptive state observer and model-based HGO can achieve satisfactory observation performance without model uncertainty. As shown in Fig. 6, when the model uncertainty is considered in simulation II, i.e., when the model parameters are changed, the observation performance of the adaptive observer designed in this paper is better than that of the modelbased HGO. Besides, it can be seen from the above simulation that the designed observer and controller are robust to timevarying disturbance. The experiment platform of SbW systems is shown in Fig. 7. In this platform, the single board computer (dSPACEds1202) is used as the control unit of SbW systems, and the servo motor driver (XiNJE DS2-20P7) is used for driving the steering motor (XiNJE MS80ST-M02430B-20P7) equipped with a reducer. The linear sensor (KTR11-10) fixed on the steering arm measures the steering angle of the front-wheels. A computer is applied to display the experimental results of the experiment on-line and store the experimental data. The sampling period is chosen as 0.001s" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000378_2016-01-1221-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000378_2016-01-1221-Figure1-1.png", "caption": "Figure 1. Comparison of motor structures", "texts": [ " On the other hand, system size and weight reduction is needed to realize a system with high general flexibility for expansion to multiple models. Honda\u2019s Intelligent Multi-Mode Drive (i-MMD) system uses two motors for driving and electric power generation. The driving motor provides power performance and the generating motor uses the engine power to generate electricity. Further space saving and high insulation performance are required of each insulating part to meet demands for smaller size and higher voltages. Figure 1 shows the magnet wire and insulation structure of the previous Honda motor and the HAW motor. i-MMD is a high-voltage system with voltages up to 700 V. The conventional motor uses enameled wire and film thickness is 0.1 mm or less. The insulate paper is mainly used to secure inter-phase insulation. The inter-phase electric potential difference is larger than the in-phase electric potential difference, so the film of magnet wire is unable to secure insulation, and insulate paper is added. Increasing film thickness of the magnet wire is a method of achieving insulation performance without increasing the number of parts, and enabled the HAW motor structure" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002264_s40430-020-02659-x-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002264_s40430-020-02659-x-Figure1-1.png", "caption": "Fig. 1 Diagram of fault-free ball bearing", "texts": [ "\u00a03, the mechanism of time-varying stiffness of bearing is studied. In Sect.\u00a04, the effects of axial preload, radial load and rotational speed on time-varying stiffness of bearing are investigated. In Sects.\u00a05 and 6, the effect of structural parameters (nominal contact angle, raceway groove curvature radius, number of balls, ball diameter) and material parameters (ball density, Poisson\u2019s ratio, modulus of elasticity) on time-varying stiffness is compared and analyzed. In Sect.\u00a07, the conclusions are elaborated and the application values are discussed. Figure\u00a01 presents a diagram of an angular contact ball bearing, where ri and re are the inner and outer raceway radius, respectively. Exerting external force on the bearing, Fx,Fy are Journal of the Brazilian Society of Mechanical Sciences and Engineering (2020) 42:578 1 3 Page 3 of 24 578 the forces that exerted on the radial (x, y) directions, Fz is the force that exerted on the axial (z) direction, and Mx,My are static moments around the x-axis and y-axis. Let x, y refer to the relative deformation in the radial direction, z refers to the relative deformation in the axial direction, x, y are the relative rotation angle around x-axis and y-axis, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002397_1.1770817-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002397_1.1770817-Figure1-1.png", "caption": "FIG. 1. Schematic diagram of respirometer.", "texts": [ "124 On: Sun, 23 Nov 2014 22:22:46 684 CONNELLY, BRONK, AND BRINK Though the instrument is especially suited to meas uring the first few seconds of a transient change in rate of oxygen uptake, it can be employed to follow the entire time-course of a change in respiration. However, in the latter case the interpretation of the relation be tween the oxygen cathode current and time is compli cated by the inward diffusion of oxygen from the sur rounding gas phase. This problem has been analyzed approximately in order to permit quantitative interpre tation over most of the time-course of the change in respiration that takes place during prolonged neural activity or during recovery therefrom. Figure 1 is a diagram of one model of this respirometer as used in the study of nerve metabolism. The oxygen. cathode is a platinum wire sealed in a cylinder of glass and mounted in the bottom of the moist chamber. The nerve lies across the end of the cylinder of glass and the platinum wire. It is held firmly against the electrode by a Lucite foot, as illustrated in detail in Fig. 2. The nerve is in contact with four silver wires (black circles of Fig. 1) which serve as stimulating and recording elec trodes. An additional wire holds the nerve in contact with wet blotting paper adjacent to the oxygen cathode. The nerve is held in place so that Ringer's or other solutions may be made to bathe it as desired, without moving it. The plane of the oxygen cathode surface and of the blotting paper is inclined so that excess solution will drain away from the portion of the nerve in contact with the glass. Moistened gas of known oxygen content is made to flow past the nerve", " Accordingly, whenever the respirometer is used in the fashion here described, the limitations imposed by the size of the diffusion system and the rates of chemical events therein must be taken into consideration.9 Calibration of an oxygen cathode consists of the ex perimental determination of the dependence of the electrode current on oxygen concentration in the medium at its tip. Calibration of the electrode in this respirome ter is carried out with the nerve mounted across the electrode as shown in Fig. 1. Gases of different oxygen content are made to flow past the nerve, and the steady state electrode current corresponding to each is recorded. (The time-course of the change in electrode current 9 One method for obtaining a numerical solution of the diffusion equation, for an unknown (Y2(/), might be mentioned. An electronic analog of the respirometer system can be constructed so that one can control an input voltage analogous to change in rate of oxygen uptake, 0l2(t), and can observe an output voltage analogous to the concentration of oxygen, uo(t), at the center of the nerve" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000403_0954405416640171-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000403_0954405416640171-Figure6-1.png", "caption": "Figure 6. (a) Initial gear model and (b) tool trace distribution.", "texts": [ " It can be divided into three processes: rough machining of tooth slot, vice-finishing and then finishing of tooth flank. Figure 5 represents the general process of the rough machining of the slot. Additionally, the simulation process of the gear is similar to the pinion, but its finishing is simpler. Accuracy analysis of tooth flank after simulation This article provides an initial gear model of spiral bevel gear whose basic geometric parameters are represented in Table 1. It is outputted from the simulation process based on the universal machine milling center, as indicated in Figure 6. Distribution of tool trace on the tooth flank from the toe to the heel is varied from the dense to sparse, in which the greater the density is , the larger the deviation is. An observation of the initial model of the simulation process indicates that the tooth form error accuracy is very poor. The tooth flank is full of so many clear and visible tool traces, even with some modifications of the different tools, the feed rates, the cutting speeds and other conditions. Especially at the fillet area, rugged tool traces with different patterns can keep mutual overlap to make it worse" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000718_978-1-4471-7278-9-Figure18.6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000718_978-1-4471-7278-9-Figure18.6-1.png", "caption": "Fig. 18.6 The self-sufficient flow meter, powered by the flow itself", "texts": [ " The idea behind self-reliance is to use the resources available within the system. Consider the design of a gas flow meter. A tiny electrical power generator is embedded inside a gas pipeline to convert gas flow to electrical current to power the circuit, creating a self-reliant gas meter that requires no external power. When gas flow stops, the meter sets to zero by default. For the safety and security of critical systems, it is common to develop the code in-house, eliminating the dependency of Dynamically Linked Libraries or COTS (Fig. 18.6). 368 18 Survivability Episodic memory provides access to an event that is experienced personally. This kind of memory is not about regularity, but rather reconstructing particularities about when and where, like in a movie. Episodic memory implies a mental reconstruction of some earlier event, including hunger, hurt, escaping from danger, and emotional reaction. Based on previous experience, we can anticipate specific events in the future. Mental time travel into the future might include the planning of a specific event, such as food gathering" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000123_s1560354719050071-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000123_s1560354719050071-Figure7-1.png", "caption": "Fig. 7. Trajectories for K3 = 0.52 on the Poisson sphere, its involute ( \u03d5 = \u03c0 4 + arctan \u03b31 \u03b32 ) and the", "texts": [ " If 0 < \u03ba2 < \u03ba1 < 1, then on the Poisson sphere there are ten isolated fixed points of different types (see Fig. 6): \u2014 four saddles (\u03b81, \u03b82, \u03b43, \u03b44); \u2014 two stable foci (\u03b83, \u03b84); \u2014 two unstable nodes (\u03b41, \u03b42); \u2014 two slow stable foci (\u03935, \u03936). We note that, in this case, the unstable and stable manifolds of equilibrium points \u03b43, \u03b44 and the central manifold at \u03b41, \u03b42 form a closed contour. Let \u03ba2 < 1 < \u03ba1. Then the equilibrium points \u03b41, \u03b42, \u03b43 and \u03b44 disappear, and an unstable limit cycle arises from the closed contour (see Fig. 7). If 1 < \u03ba2 < \u03ba1, then of all the equilibrium points the system (4.1) has only \u03935 and \u03936 and the unstable limit cycle (see Fig. 8). A detailed classification of phase portraits on the Poisson sphere and of trajectories of the contact point depending on the system parameters remains an open problem. REGULAR AND CHAOTIC DYNAMICS Vol. 24 No. 5 2019 trajectory of the contact point with the fixed initial conditions \u03b1(0) = (1, 0, 0), \u03b2(0) = ( 0, 3 \u221a 11 50 ,\u2212 49 50 ) , \u03b3(0) = ( 0, 49 50 , 3 \u221a 11 50 ) , X(0) = 0, Y (0) = 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000716_s00170-019-03372-3-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000716_s00170-019-03372-3-Figure2-1.png", "caption": "Fig. 2 Schematic of coordinate system rotation in threedimensional view (a) and front view (b)", "texts": [ " The third term at the right side is the source term due to the frictional dissipation in the mushy zone. The drag coefficient is determined by the Carman\u2013Kozeny equation, which is given by [24] K \u00bc f l 3d2 180 1\u2212 f l\u00f0 \u00de2 \u00f05\u00de where d is the distance. Mass equation \u2202\u03c1 \u2202t \u00fe \u2207\u22c5 \u03c1V ! \u00bc Sm \u00f06\u00de where Sm is the source term. As shown in Fig. 1, during hybrid fillet welding, the hybrid welding torch is inclined, which can significantly affect the distribution modes of heat flux and the related forces. In this study, the inclination of hybrid heat source is dealt with by coordinate transformation. Figure 2 gives the schematic of coordination system rotation. When the coordinate system (x, y, z) rotates with the heat source together, the distribution function of the heat source in the new coordinate system (xr, yr, zr) is the same as that in the original one. Therefore, after obtaining the relationship between these two systems, a specific expression of inclined arc heat source can be determined easily through coordinate replacement. This relationship can be described by the following matrix equation xr yr zr 2 4 3 5 \u00bc 1 0 0 0 cos\u03b8 sin\u03b8 0 sin\u03b8 cos\u03b8 2 4 3 5 x y z 2 4 3 5 \u00f07\u00de where \u03b8 is the inclination angle" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001452_j.mechmachtheory.2020.104061-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001452_j.mechmachtheory.2020.104061-Figure3-1.png", "caption": "Fig. 3. Ring gear model.", "texts": [ " Nomenclature A b Swept area of the rotor b sp , b rp Gear backlash in sun-planet meshes, and ring-planet meshes C p Power coefficient c\u0304 s , \u0304c r Sun gear contact ratio, ring gear contact ratio E\u0304 , G\u0304 Young modulus, shear modulus f 0 Produced deterministic force by linearization process F d , F e Deterministic Force, Force generated by stochastic linearization f e Equivalent linear meshing function f l Node force array caused by moving mesh F r Moving mesh forces f sp , f rp Nonlinear meshing function I i Moment of inertia of component i ( i = s, c, p) k sp , k rp Sun-planet pair mesh stiffness, ring-planet pair mesh stiffness k\u0304 sp , \u0304k rp Average of the sun-planet, and ring-planet mesh stiffness k e Equivalent mesh stiffness k i Bearing stiffness in transnational directions ( i = s, c, p) k l Stiffness matrix of l th element k b, v , k b, \u03c4 , k b, \u03b8 Support points\u2019 stiffness (see Fig. 3 a) k su Torsional stiffness of sun gear m i Mass of component i ( i = s, c, p) m l Mass matrix of l th element N Total number of planets p i , q i Fourier series\u2019 harmonic coefficients ( i = sp, rp) R b Wind turbine blades radius R i Base circle radius of gear, radius of the carrier ( i = s, c, p) S 0 White noise spectral density T c Period of the carrier T in Torque caused by the wind T m Period of the mesh u \u03c5, u \u03c4 Ring displacements in radial and tangential direction V, V 0 , V t Wind velocity, mean wind velocity, turbulence ( V = V 0 + V t ) W ( t ) White noise x, y Transnational displacement for sun gear and carrier \u03b1s , \u03b1r Pressure angle of sun gear, and ring gear , \u03b3 Independent and identically distributed random variables \u03b4sp , \u03b4rp Displacement along action line sun-planet, ring-planet mesh \u02c6 \u03b4 Dirac delta function t Time step \u03c3 Variance of random variable \u03c9 Meshing frequency \u03c9 c Carrier angular velocity \u03c9 r 0 Rotation speed of the rotor. i Meshing phase ( i = sp, rp, sr) \u03c8 n Position of the n th planet (see Fig. 1 ) \u03c1 Density \u03c1a Air density \u03b8 i , u i Rotations of the component i ( i = s, c, p), u i = \u03b8i \u00d7 R i \u03b8 r Angular coordination of moving load (see Fig. 3 b) k a i Amplitude of the stiffness variation ( i = sp, rp) \u03c5\u03c4 Coordinates of ring\u2019s element in radial and tangential direction \u03be\u03b7 Transnational displacement for planet gears c Carrier p Planet gear r Ring gear s Sun gear rp Ring-planet pair sp Sun-planet pair dynamic behavior of the planetary gear of wind turbines by considering contact loss, mesh stiffness variation, and bearing clearance. They discovered that the PGT showed the chaotic responses because of internal excitation due to the clearance", " The equations of motion of the carrier then are: m c x\u0308 c + k c x c + N \u2211 n =1 k p,n [ x c \u2212 \u03ben cos \u03c8 n + ( \u03b7n \u2212 u c ) sin \u03c8 n ] = 0 (9) m c y\u0308 c + k c y c + N \u2211 n =1 k p,n [ y c \u2212 \u03ben sin \u03c8 n + ( \u03b7n \u2212 u c ) cos \u03c8 n ] = 0 (10) I c R 2 c u\u0308 c + N \u2211 n =1 k p,n ( \u2212x c sin \u03c8 n + y c cos \u03c8 n + u c \u2212 \u03b7n ) = T in R c (11) where k c is the carrier bearing stiffness in the transnational direction, k p,n is the n th planet bearing stiffness, and T in is the torque caused by the wind which is discussed in Section 3 . 2.3. Ring gear The ring gear is the outer fixed internal gear which meshes with planetary gears. The ring gear is modeled as a smooth thin-walled ring connected to the frame through a certain number of bolt connections. Each bolt connection is modeled as three springs, namely k sc,r , k sc, \u03c4 , and k sc, \u03b8 ( Fig. 3 a). Curved beam elements are used to discretize the ring with a typical element shown in Fig. 3 b. The stiffness and mass matrices of the ring are created by curved beam theory, and the meshes between the ring-planet pairs are considered as moving loads. 2.3.1. Element stiffness and mass matrix The mass and stiffness matrices of an element are calculated based on [24] . Here, we directly give the final equations, and details can be found in the reference paper and in Appendix B . [ k l ] = E\u0304 I l R 3 r [ D ] [ B ] \u22121 l = 1 , 2 , . . . L (12) [ m l ] = \u03c1R r [ B ] \u22121 T \u222b \u03b82 \u03b81 ( [ H ] T [ ] [ H ] d\u03b8 ) [ B ] \u22121 l = 1 , 2 , ", " The mean and variance of the displacement of the ring gear\u2019 node, planet gear, sun gear, and carrier are compared with the MCS results. In this section, for brevity, only the response of one specific direction of each component is presented in following. The results generated by the MCS for node 3 were identical to results of the linearized model ( Fig. 5 ). The mean and variance of displacement responses of nodes 3 at tangential direction are shown in Fig. 5 . The positions of the nodes are shown in Fig. 3 a where T c is the period of the carrier. The figure shows that both the gear meshing frequency and the planet frequency are presented in the response. Three peaks show the planet frequency in one carrier period. In addition, based on the obtained results from the linearized model and MCS, both methods agree well. It should be noted that the number of elements affects the phase difference in response of the nodes, and in this study, the phase difference between every two adjacent nodes was 1 18 carrier period" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002210_j.robot.2020.103670-Figure21-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002210_j.robot.2020.103670-Figure21-1.png", "caption": "Fig. 21. Contact avoidance between the robot\u2019s front wheels and the step. (a) The robot\u2019s front wheels under the step edge (0 \u2264 y2 \u2264 h), (b) The robot\u2019s front wheels above the step edge (h < y2).", "texts": [ " 1 B Bf s d 4 b r 0 i t s B 0 F v \u2225 r f 4 i h r v B 0 w c \u2225 \u221a h t i 5 e \u00b5 a u 5 A r t v w b a o s b When the robot climbs a step the height of which is h, substitute y3 = h for (14), cos\u03c61 = e1(e1 \u2212 h) + WBf \u221a 2 \u00b7 e1 \u00b7 h + WBf 2 \u2212 h2 e12 + WBf 2 (15) Using (15), the system can tilt the robot and control its angle ufficiently to put the wheels on the step when the vehicles etect the step height. .3. Requirement to avoid situations where the vehicles cannot move ecause the robot\u2019s front wheel axes are lower than the step height Fig. 21(a) is a model showing the robot and a step when the obot\u2019s front wheels are lower than the step. Here, B 0p\u2206x = [\u2206x 0]T is the position vector from O to O\u2032 in \u03a3B (see Figs. 19 and 21 (a)). B0ps is a vector in 0\u03a3B, whose height s the same as that of the robot\u2019s front wheel axes (y2), and L is he horizontal distance from the robot\u2019s rear wheel axes to the tep wall. ps = [L y2]T (16) rom Fig. 21(a), the requirement to avoid situations where the ehicles are unable to move is as described below. B 0ps \u2212 (B0p\u2206x + Bq2)\u2225 > rBf (0 < x2 < L, 0 < y2 < h) (17) (7), (10) and (16) are substituted for (17),\u221a (L \u2212 \u2206x \u2212 x2)2 > rBf (18) (18) describes the requirement to avoid contact between the obot\u2019s front wheels and the step when the height of the robot\u2019s ront wheel axes is lower than the step height. .4. Requirement to avoid situations where robot\u2019s movements are mpossible because the robot\u2019s front wheel axes are higher than step eight Fig. 21(b) is a model showing the robot and step when the obot\u2019s front wheels are higher than the step. B0pse is the position ector of the step edge in 0\u03a3B. pse = [L h]T (19) here, h is the step height. In this case, the requirement to avoid ontact between front wheels and the step is as given below. B 0pse \u2212 (B0p\u2206x + Bq2)\u2225 > rBf (0 < x2 < L, h \u2264 y2) (20) (7), (10) and (19) are substituted for (20), (L \u2212 x2 \u2212 \u2206x)2 + (h \u2212 y2)2 > rBf (21) (21) is the requirement to avoid contact between the robot\u2019s front wheels and the step when the height of the robot\u2019s front wheel axes are higher than the step height. Comparing with (18) and (21)(Fig. 21(a) and (b)), Eq. (18) shows it is more difficult for the vehicles to avoid contact between the front wheels and the step. In this step-climbing method, cos0 \u03c6123 \u2264 cos\u03c6123 and cos0 \u03c612 \u2264 cos\u03c612 when the robot lifts its front wheels. From (5), the maximum value of \u2206x, \u2206xmax, is as described below. \u2206xmax = lLB(1 \u2212 cos\u03c61) + hLB sin\u03c61 \u2265 \u2206x (22) Thus, B 0p\u2206x(max) : [\u2206xmax 0]T is given below. Bp : 0 \u2206x(max) [ \u2206xmax 0 ] = [ lLB(1 \u2212 cos\u03c61) + hLB sin\u03c61 0 ] (23) In cases where \u2206x is \u2206xmax, and when L\u2212\u2206xmax \u2212 x2 > 0, (7), (10) and (22) are substituted for (18), L > (WBf \u2212 lLB) cos\u03c61 + (RB \u2212 rBf + hLB) sin\u03c61 +lLB + rBf (24) Using (15) and (24), the system is capable of determining the eight of the front wheels in order to place them on the step, and he distance from the step required to lift the robot\u2019s front wheels n order to avoid creating an immovable state" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0003134_1.1714891-Figure11-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0003134_1.1714891-Figure11-1.png", "caption": "FIG. 11.", "texts": [ " (40) Cn Dn C D 438 w2r=~[1-cOS (Y1r)] =~ sin2 (r1l\"), (48) Ik n Ik 2n or wr = (I:)t sin (~:) r= 1,2, ... (n-1). (49) The possible natural angular frequencies of the system are given by (49). JOURNAL OF APPLIED PHYSICS [This article is copyrighted as indicated in the article. Reuse of AIP content is subject to the terms at: http://scitation.aip.org/termsconditions. Downloaded to ] IP: 131.111.164.128 On: Sat, 20 Dec 2014 02:03:32 b. Natural Frequency of Shaft Fixed at Both Ends Let us now consider the system of Fig. 11. In this case the two ends of the system are fastened to rigid supports as shown. We can write 82 =0, and hence { T1} [1 O][An Bnl[ 1 0]{T2} (50) 81 = kl2 1 en Dn kl2 1 \u00b0 . Carrying out the matrix multiplication, we obtain: Since the left end of the system is fixed, 81 = \u00b0 and hence the right member of (51) must vanish. Substituting the values of An, B n, en, and Dn from (35) into the resulting expression, we obtain: cosh (an) +sinh (a) sinh (an) 4 cosh 2 (aI2) or sinh (an) cosh2 (aI2) +------- sinh (a) 0, (52) spaced, free at the left end and fixed at the right as shown", " In this case we have 82 =0 and we can write: { Tl} [An Bn][ 1 O]{ T2} (56) 81 = en Dn kl2 1 0' Carrying out the matrix multiplication, we have: (57) Since T 1 =0, we must have ( An+~Bn) = cosh (a!1) sinh (a) sinh (an) + 0. 2 cosh2 (aI2) (58) This may be written in the form (2n+1)a (a) cosh 2 cosh \"2 0. (59) cosh 2 (aI2) Hence cosh [(2n+1)aI2J=0. (60) [sinh (n+ 1)alsinh (a) ] = 0. (53) This leads to The solution of the transcendental equation (53) is a r =nrj/(n+1), r=1, 2, 3, .. \u00b7n. (54) Substituting these values of a into (36), we obtain for the natural frequencies of the system of Fig. 11. c. Natural Frequencies of Shaft Fixed at One End The system of Fig. 12 represents a uniform shaft carrying n identical disks equidistantly VOLUME 13, JULY, 1942 (2r-1)7I'j ar , r= 1,2, .. \u00b7n. (61) (2n+1) Then from (36) we obtain 2 [(2r-1)7I'] wr = (Ik)! SIO 2(2n+1) , r= 1, 2, .. \u00b7n. (62) d. Natural Frequencies of General Terminal Conditions In the simple cases that we have considered, the frequency equations have been particularly simple transcendental equations whose solution was easily written down" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000123_s1560354719050071-Figure8-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000123_s1560354719050071-Figure8-1.png", "caption": "Fig. 8. Trajectories for K3 = 3 on the Poisson sphere, its involute ( \u03d5 = \u03c0 4 + arctan \u03b31 \u03b32 ) and the trajectory of", "texts": [ " 6): \u2014 four saddles (\u03b81, \u03b82, \u03b43, \u03b44); \u2014 two stable foci (\u03b83, \u03b84); \u2014 two unstable nodes (\u03b41, \u03b42); \u2014 two slow stable foci (\u03935, \u03936). We note that, in this case, the unstable and stable manifolds of equilibrium points \u03b43, \u03b44 and the central manifold at \u03b41, \u03b42 form a closed contour. Let \u03ba2 < 1 < \u03ba1. Then the equilibrium points \u03b41, \u03b42, \u03b43 and \u03b44 disappear, and an unstable limit cycle arises from the closed contour (see Fig. 7). If 1 < \u03ba2 < \u03ba1, then of all the equilibrium points the system (4.1) has only \u03935 and \u03936 and the unstable limit cycle (see Fig. 8). A detailed classification of phase portraits on the Poisson sphere and of trajectories of the contact point depending on the system parameters remains an open problem. REGULAR AND CHAOTIC DYNAMICS Vol. 24 No. 5 2019 trajectory of the contact point with the fixed initial conditions \u03b1(0) = (1, 0, 0), \u03b2(0) = ( 0, 3 \u221a 11 50 ,\u2212 49 50 ) , \u03b3(0) = ( 0, 49 50 , 3 \u221a 11 50 ) , X(0) = 0, Y (0) = 0. trajectory of the contact point with the fixed initial conditions \u03b1(0) = (1, 0, 0), \u03b2(0) = ( 0, \u221a 79 40 ,\u2212 39 40 ) , \u03b3(0) = ( 0, 39 40 , \u221a 79 40 ) , X(0) = 0, Y (0) = 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001322_j.epsr.2020.106409-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001322_j.epsr.2020.106409-Figure3-1.png", "caption": "Fig. 3. Simplified model of SWT coupled to PMSG and a front-end diode rectifier block.", "texts": [ " A general schematic of the proposed model covering the previous explained parts is illustrated in Fig. 2. It clearly shows the MPPT process to attain the MPP of the small wind turbine. The intelligent modular MLP approach and simplified model of WECS, blue rectangles, for solving the MPPT process will be introduced in section 2.2. Ignoring the shaft stiffness and damping effect of the SWT, and considering only the inertia, J, of the PMSG and the SWT, then the wind turbine coupled to PMSG along with the three-phase diode rectifier (Block A in Fig. 1) can be simplified, as shown in Fig. 3. The gain Kg couples the mechanical-electrical system and defines the stable state response of the idc. According to Ziegler-Nichols [25] and neglecting the time-delay in Fig. 3, the transfer function, from the rotational speed to the output current of the three-phase diode rectifier, can be approximated in the frequency domain as a first-order model: = = + G s s I s K s ( ) ( ) ( ) 1i m dc g m dc (1) where \u03c4 is the time constant, which is directly proportional to the amount of the moment of inertia, J, of the small wind turbine in Fig. 3. It is important to consider that an accurate computation of idc is essential for the MPPT process. Based on equation (1), the dynamic discrete-time model can be deduced with the zero-order hold (ZOH) method, as follows: = + = + G z Z e s K s K e z K z e z ( ) 1 1 1i sT g g T g T / 1 1 / 1m dc a a a (2) Applying the difference equations method from [29] in (2), then the expression to compute idc based on \u2126m can be obtained by: = + +i i k K e k K k i k e[ ] [ 1] [ 1] [ 1]dc dc g T m g m dc T/ /a a (3) where Ta represents the sample time of the MPPT algorithm, [k] and [k1] are the current and the previous sampling time, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002002_012076-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002002_012076-Figure7-1.png", "caption": "Figure 7. Loading during analysis", "texts": [ " Considering the change in mesh size deformation was observed and there was no noticeable change in deformation observed in 5 iterations over a range of 20mm to 30mm element size, hence that concludes that the results are converged. The following types are used to infer regarding the convergence and accuracy, in this analysis, hence the H type method of convergence is been used. The calculated forces like bump force were applied by the road to the tire of the wheel with a force of 2500N and the analysis is carried out in the no-slip condition as shown in figure 7. According to the simulation study mentioned, the analysis was carried out in ANSYS workbench and the following results were obtained. The individual analysis was carried out for the existing model and the proposed model with element size was 5mm to get the most efficient results in the stipulated amount of time. The Von-Mises stress, strain and deformation plots are as shown in figure 8, 9, 10 for air package as material for existing tire. Figure 11, 12 and 13 explain the von mises stress, strain and deformation plots for polyethylene as a material" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001673_csde50874.2020.9411585-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001673_csde50874.2020.9411585-Figure1-1.png", "caption": "Fig. 1. The ith individual with an omni-directional detecting sensor situated at (xi, yi) with detection range of rd.", "texts": [ " Thus, xi := (xi, yi) \u2208 R 2 is the configuration vector for the ith individual and x := (x1,x2,x3, ...,xn) \u2208 R 2n becomes the configuration vector for n individuals with the initial conditions vector denoted by x0 := (x1(0),x2(0),x3(0), ...,xn(0)) \u2208 R 2n. Definition 2.1: The ith individual is a point mass residing in a disk with center (xi, yi) and radius ri > 0. It is described as the set Bi := {(z1, z2) \u2208 R 2 : (z1 \u2212 xi) 2 + (z2 \u2212 yi) 2 \u2264 r2i }. (1) The ith individual has an omni-directional detecting sensor situated at (xi, yi) with detection range of rd as shown in Figure 1. There will be communication between the ith individual and jth if and only if they are in the detection range of each other. This means that the behaviour of the ith individual is influenced by its neighbours only. There is no communication between those individuals, which are not in the detection range of each other. Thus, the motion of those individuals will not be influenced by each other. At t \u2265 0, let (vi(t), wi(t)) := (x\u2032 i(t), y \u2032 i(t)) be the instantaneous velocities of the ith individual" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002523_tim.2021.3062194-Figure10-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002523_tim.2021.3062194-Figure10-1.png", "caption": "Fig. 10. Typical bearing faults. (a) Outer race fault. (b) Inner race fault. (c) Roller fault.", "texts": [ " In this research, defect characters are premanufactured to the bearings. The experimental parameters (including the bearing parameters) are shown in Table II. Due to the hardware limitations and the safety requirements in the laboratory, the rotating speed is set as 480 rpm in this experiment. By using the electric sparks machining, the failures of bearings are simulated by implanting damage into the outer race, inner race, and roller. The fault depth and width were kept at 0.3 and 1.2 mm, respectively, for all types of faults in this experiment. As shown in Fig. 10, four bearing states (three faults and normal state) are considered in this experiment. Fault types and fault dimensions are listed in Table III. Fig. 11 presents three typical time-domain waveforms for different operation states. Ideally, repetitive impacts can be observed from the acquired time-domain waveform corresponding to the rotation of the rolling elements past the damage. The modulation also occurs at the time period corresponding to every shaft revolution. Accordingly, by building the relationship between the repetitive impacts and the frequency of a rotating component, fault identification can be achieved" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001761_tasc.2020.2968043-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001761_tasc.2020.2968043-Figure1-1.png", "caption": "Fig. 1. CMG topology. (a) Conventional model. (b) Proposed model.", "texts": [ " Therefore, it is necessary to research the CMG with simple structure, high torque density and steady transmission. The purpose of this paper is to propose a novel CMG with unequal Halbach arrays, which is capable of improving the torque density and reducing the torque ripple. The topology of CMG with unequal Halbach arrays will be introduced in Section II. In Section III, the performances of the proposed CMG will be computed by the finite element analysis (FEA). Finally, several conclusions are given in Section IV. Fig. 1(a) shows the topology of conventional CMG. The CMG is mainly composed of the following components, there are two rotors with surfaced mounted PMs, and it has a stationary ring with a number of ferromagnetic segments and epoxy resin. The stationary ring is sandwiched between the two PM rotors, which takes charge of modulating the magnetic fields in two air-gaps. Fig. 1(b) shows the proposed CMG, which has three eccentric and unequal small PMs at each pole of the inner rotor, and sinusoidal magnetization is adopted. On the outer rotor, PMs are arranged in regular Halbach arrays, and each pole has two small PMs. The structure of eccentric and unequal small PMs is shown in Fig. 2. Fig. 2 shows the structure of the PMs on the inner rotor of the proposed model. The PM of each pole is divided into three small pieces, and the PM is an eccentric structure with unequal width" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001800_978-3-030-40410-9-Figure4.10-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001800_978-3-030-40410-9-Figure4.10-1.png", "caption": "Fig. 4.10 Operative ranges and specifications of LabVolt 5150", "texts": [ "C12 a1.S1 + a2.S12 0 \u23a4 \u23a6. (4.8) Also, the end-effector orientation relative to the base frame is R0 2 = \u23a1 \u23a3C12 \u2212S12 0 S12 C12 0 0 0 1 \u23a4 \u23a6. (4.9) Example 4.2 (LabVolt 5150 manipulator) Figure 4.8 shows LabVolt 5150 manipulator (Ghafil and J\u00e1rmai 2019), which is a 5DOF, four-link educational robot. The joint limits of the robot actuators, as well as link specifications, are given in Fig. 4.9. To start modelling the forward kinematics equations of this robot, zero-configuration was chosen to be as in Fig. 4.10. The same way as Example 4.1 has been followed to assign six-coordinate frames starting from the base point to the end-effector central point. Note that frame 3 and frame 4 coincide at the same point. Table 4.2 reveals the spatial parameters of the four links of the manipulator. Referring to Fig. 4.10, we have substituted the unknowns in Table 4.2 as follows: d1 = 255.55mm a2 = a3 = 190mm 78 4 Manipulator Kinematics d5 = 115mm The fourHTMs can be determined by substituting each link parameter in Eq. (4.4). The set of HTMs will be H 0 1 = \u23a1 \u23a2\u23a2\u23a3 C1 0 S1 0 S1 0 \u2212C1 0 0 1 0 d1 0 0 0 1 \u23a4 \u23a5\u23a5\u23a6, (4.10) 4.4 Denavit\u2013 Hartenberg (DH) Convention 79 H 4 5 = \u23a1 \u23a2\u23a2\u23a3 C5 \u2212S5 0 0 S5 C5 0 0 0 0 1 d5 0 0 0 1 \u23a4 \u23a5\u23a5\u23a6, (4.14) H 0 5 = H 0 1 .H 1 2 .H 2 3 .H 3 4 .H 4 5 . (4.15) Equation (4.15) is 4\u00d7 4 matrix holding the orientation and position vector of the end-effector with respect to the base frame, as revealed in Eqs" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000660_1.4035560-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000660_1.4035560-Figure3-1.png", "caption": "Figure 3- Plenum assembly with fixture for CD measurements of engineered-porous cavities", "texts": [ "org/about-asme/terms-of-use Ealy 6 GTP-16-1356 Figure 1-(a) Test article with dimensions, (b) Cut plane showing internal features*, (C) Internal lattice structure* *Dark shaded surfaces exposed to fluid* Figure 2- Preliminary print of inconel test article with inverted build orientation A state-of-the-art NSI X5000 x-ray CT system was used to scan the test article. Both 225kV and 450kV x-ray tubes with microfocus and flat panel detector were used, providing image resolutions of 60\u03bcm and 100\u03bcm, respectively. CAD model comparison of the volume reconstruction from the 450kV scan is performed. A conditioned plenum was manufactured to handle the test article and deliver uniform, steady flow. Details of the experimental rig can be found in Figure 3. A large volume of compressed air is supplied to the plenum, providing steady flow over 1\u2264 \ud835\udc43\ud835\udc45 \u22641.3. Flow rate is controlled via in-line pressure regulator and measured with an Omega FMA-1613A mass flow meter, providing flow rates with a maximum uncertainty of \u00b14.7L/min. Multiple static pressures are measured along the flow path to ensure flow uniformity and to quantify \ud835\udc431. These static pressures are sampled with an Omega HHP242-138D, certain to \u00b134.47 Pa. Flow exits into ambient laboratory environment in all cases" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001812_1077546320912109-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001812_1077546320912109-Figure3-1.png", "caption": "Figure 3. (a) Outer race, (b) inner race, and (c) roller defect model for a cylindrical roller bearing.", "texts": [ " In this section, mathematical simulation of local defects (spalls) in races and rollers is given. For a given size (s) of the localized defect, the angular width (\u03a6r) will vary with the location of the defect, viz., inner race, outer race, or roller. As the rollers start repetitive hitting of the races in the load zone, stresses are developed, which causes the initiation of deformation at the subsurface level, and gradually spread to surfaces and then grow. 2.3.1. Outer race defect. A small rectangular spall is considered in the loading direction at the outer race (Figure 3(a)). As the outer race is stationary, the defect position will not change with speed and whenever the rollers pass through a defect, a high impact impulsive force is generated because of the sudden loss and gain of contacts between rollers and races. The position of the jth roller is \u03b8j \u00bc 2\u03c0\u00f0j 1\u00de NR \u00fe \u03c9caget (10) The deformation of the jth roller while crossing the defect zone is defined as follows. When 0 < \u03b8j < \u00f0S=R\u00de, an additional deflection is added to the rolling element, i.e. \u0394j \u00bc Rr Rr cosfr (11) where fr \u00bc \u00f0S=2Rr\u00de (S = defect width and R = outer race radius, Rr = roller radius). The ball pass frequency (BPFO) generated because of the defect at the outer race is BPFO \u00bc Nr 2 \u03c9shaft 1 Rr Rp (12) 2.3.2. Inner race defect. Unlike the case for the outer race, the inner race defect (Figure 3(b)) is rotating at shaft speed. So, the relative speed is generated between the defect and rollers, as rollers will be rotated with cage speed. The contact position can be calculated as \u03b8j \u00bc 2\u03c0\u00f0j 1\u00de NR \u00fe \u03c9caget \u03c9shaftt (13) An additional deflection \u00f0\u0394j\u00de is added when the roller position is at the defect angle, i.e.\u03c9shaftt < \u03b8j <\u03c9shaftt \u00fe \u00f0S=R\u00de. The ball pass frequency (BPFI) generated because of the inner race defect is BPFI \u00bc Nr 2 \u03c9shaft 1\u00fe Rr Rp (14) 2.3.3. Roller defect. The defect present in the roller (Figure 3(c)) will rotate at a speed of roller which will get impact twice by individual raceways. So, the position of the defect at each time instant can be calculated as \u03b8j \u00bc 2\u03c0\u00f0j 1\u00de NR \u00fe \u03c9Rt (15) where the speed of roller\u03c9R \u00bc \u00f0Rp=2Rr\u00de\u03c9shaft\u00f01 \u00f0Rr=Rp\u00de\u00de. Additional deflection \u00f0\u0394j\u00de is added when the position of the roller satisfies the following condition. \u03c9Rt < \u03b8j <\u03c9Rt \u00fe \u00f0S=R\u00de (when the defect interacts with the outer race). \u03c9Rt < \u03b8j <\u03c9Rt \u00fe \u00f0S=r\u00de (when the defect interacts with the inner race). The ball spin frequency (BSF) generated because of the defect at the roller is BSF \u00bc Rp 2Rr \u03c9shaft 1 Rr Rp 2 " ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001844_s12541-020-00333-9-Figure8-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001844_s12541-020-00333-9-Figure8-1.png", "caption": "Fig. 8 Von-Mises stress in flexspline", "texts": [ " The finite model consists of 56,266 elements and is shown in Fig.\u00a07. The material settings used in the finite element model are shown as Table\u00a04. The the bottom of the flexspline and inner hole of the wave generator are fully restrained. The contact type of flexspline, wave generator and circular spline are represented by CONTAC185 in ANSYS. The applied torque between the circular spline and flexspline is 25 Nm. The results indicate that the stress is concentrated at both ends of the teeth, especially at the back end of the teeth (Fig.\u00a08). It can be seen, that, when aided by the wave generator, the stress distribution at the root of the gear tooth is unevenly distributed in the direction of tooth thickness. Stress builds up at the front and back end of the teeth. 1 3 To verify the validity of the FEM model, we compared the simulation results with the calculated results using a Hertz contact model. Hertz contact theory is used to analyze strain and stress distribution of two objects during compressive contact. It uses three assumptions: (1) small deformation in the contact area; (2) elliptical contact area; (3) the contact object with distributed vertical pressure can be regarded as elastic half space" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001616_978-3-030-17747-8-Figure19-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001616_978-3-030-17747-8-Figure19-1.png", "caption": "Fig. 19 C1: force-displacement curves and deformed configurations for GAM", "texts": [], "surrounding_texts": [ "Buckling of axially compressed cylinders on substrate was studied numerically in [51, 49], where small geometric imperfections and small perturbation force were used, respectively, to trigger the buckling. In our analyses, geometry is perfect and no perturbation force is applied. In [49] a coefficient CS = Es/E f ( R/t f )3/2 is proposed, with the critical value CS,cri t \u2248 0.88. Here, subscripts f and s refer to film (i.e. cylinder) and substrate, respectively. For systems with CS < CS,cri t , only axisymmetric wrinkling mode occurs, whereas for CS > CS,cri t the transition from axisymmetric to diamond-like mode is expected. Figure 18 shows two considered data sets; C1 and C2. Compressive axial displacement uy is imposed at both ends according to presented load function. BAM and GAM with \u03c1\u221e = 0.9 and EMC are used. Figures 19 and 20 show reaction force at x = 0 with respect to axial and radial displacements, together with GAMdeformed configurations. Both cylinders initially produce axisymmetric pattern, for all time-stepping schemes. For C1, all schemes capture the transition to the diamond-like pattern. For C2, the transition is captured by GAM and EMC. We may conclude that dissipation has stronger influence on results for cylinders with lower CS factors. Theoretical axisymmetric buckling force can be computed as, see [49] fcr,el = E f [ 1 p20 + t2f p 2 0 4c2R2 + 3RE\u0304s 2c2t f E\u0304 f p0 ] t f [ N mm ] , (24) 408 M. Lavrenc\u030cic\u030c and B. Brank Case C1 C2 fcr,el [N/mm] 0, 060 1, 27 fcr [N/mm] ( fcr fcr,el [%] ) 0, 056(93%) 1, 23(97%) where c = \u221a 3 ( 1 \u2212 \u03bd2 f ) , E\u0304s = Es 1\u2212\u03bd2 s , E\u0304 f = E f 1\u2212\u03bd2 f and p0 is a critical wave number in the axial direction, obtained by solving \u22122 + t2f 6(1\u2212\u03bd2)R2 p 4 0 \u2212 3R 6(1\u2212\u03bd2)t f E\u0304s E\u0304 f p0 = 0. (25) Numerical results match well with the theoretical critical force, see Table 7, where fcr is computed force, which is the same for all schemes." ] }, { "image_filename": "designv11_5_0002356_978-3-030-50460-1-Figure27.3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002356_978-3-030-50460-1-Figure27.3-1.png", "caption": "Fig. 27.3 Generation of the random polycentric arch: (a) location of C2 on the line joining the previous sector\u2019s end point and its C1; (b) the new arc sector is created according to the random amplitude of its angle of embrace and to the length of its radius.", "texts": [ " In order to achieve a simulation of the constructive method closer to the real one, the modelling procedure of the polycentric arches was addressed with a probabilistic approach by means of input random variables, which allow to obtain asymmetrical and irregular 2D random arches, starting from a nominal deterministic geometry. Five arcs are generated according to the values of the aforementioned parameters and assembled sequentially maintaining the tangent continuity all along the polycentric arch. The centre of every sector was forced to be on the line connecting the end point of the previous arc with its centre, ensuring the same tangent between adjacent arc sectors (Figure 27.3). The circular arch has been selected as nominal geometry and it is defined by the following geometrical parameters{ \u03b1 = 180\u25e6 \u2212 2\u03b2 = 157.5\u25e6 t/rn = 0.15 (27.1) being \u03b1 the angle of embrace, \u03b2 the angle at the springer, rn the radius and t the thickness. In the proposed model the angles, respectively of embrace (\u03b1\u0303i) and at the springer (\u03b2\u0303i), and the radius of each single sector (r\u0303i) are defined as random variables of whom the values are extracted starting by uniform probability density functions, referred to the nominal features for the mean values and to a priori tolerance values (\u03b6,\u03c7,\u03b5) for the lateral bounds delimiters (Figure 27" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0003137_1.1752269-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0003137_1.1752269-Figure6-1.png", "caption": "FIG. 6. Schematic diagram of the pendulum system.", "texts": [ " The simplicity of the result permits its use, but it is necessary to develop the accurate method and from that to show that the formula (S') may be safely used for relative measurements. The purpose of this section is This article is copyrighted as indicated in the article. Reuse of AIP content is subject to the terms at: http://scitationnew.aip.org/termsconditions. Downloaded to IP: 35.46.82.124 On: Wed, 03 Dec 2014 21:07:45 thus to consider the mechanical construction of the pendulum and to develop the accurate expression for the coefficient of friction. Only those approximations will be made which leave the results accurate to within 1 percent. Fig. 6 shows the mechanical system of the pendulum. The pendulum rod ON is rigidly attached to the axle A which slides on the spheres Band C. The mass M of the pendulum is considered concentrated at N. 2() is the angle which the contact points of the spheres and the axle make with the center of the axle O. Fl and F2 are the forces which act upon the two spheres, F is the force in the direction ON, and a is the angle ON makes with the vertical, OP. a is the length, ON, of the pendulum and r is the radius of the axle", " Solving (9) and (10) for (F1+F2), we find: F\u00b7 (cos a\u00b1p\u00b7sin a) (11) The friction moment R about 0 is, however, (12) Setting the value of (F1+F2) from (11) in (12), dissipated in a half-period is f \"'Y+l R\u00b7da, \"'Y (14) where 'Y is the number of the half-period, the 1\"'Y+1 term p\u00b7sin a\u00b7da may be neglected in com- \"'Y 1 \"'Y+l parison to cos a\u00b7 da when the system is \"'Y used to measure the coefficients of friction of oils. We may then use for our approximate friction moment: p\u00b7r\u00b7F cos a R=--\u00b7--. (1 +p2) cos () (15) The expression (13) may be easily integrated, but it complicates the problem without adding to the accuracy. Fig. 6 shows that the force F is the sum of the component of the weight along ON and the centrifugal force: ( da)2 F=M\u00b7g\u00b7cos a+M\u00b7a\u00b7 dt . (16) we have From the energy equation of the pendulum, p \u2022 r . F . (cos a \u00b1 p \u2022 sin a) R= (13) !\u00b7Ma2. (::Y +M\u00b7g\u00b7a\u00b7 (cos a'Y-cos a) =0 (17) Since p is small (\"-'0.1) and since the energy where a-y is the maximum amplitude of the 'Yth This article is copyrighted as indicated in the article. Reuse of AIP content is subject to the terms at: http://scitationnew.aip.org/termsconditions" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000765_1.4043367-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000765_1.4043367-Figure6-1.png", "caption": "Figure 6: Tooth root crack geometric measurements. (a) front view; (b) isometric view.", "texts": [ " The mesh stiffness reflects the toggling between single- and double-tooth engagements. It can be observed that the stiffness is reduced with the increase of the fault dimensions. For the cases, CT 04\u2013MT 08, a shift is observed in the graph as a result of a decrease in the contact length along the involute profile of the damaged tooth. A crack was simulated in the root of a single tooth. We assumed the crack exists through the whole tooth width, with a constant length, lc, and a crack inclination angle, \u03b1c. The geometry is presented in Figure 6. Severity was set by increasing the crack length lc by discrete amounts (lc = 2, 3, 4, 5mm). The crack inclination angle was set to \u03b1c = 75.5\u25e6. The effect of described tooth root fault on the dynamic response is expressed by reduction of the tooth stiffness without changing the tooth profile along the contact area. A ce pt ed M an us cr ip t N ot C op ye di te d Downloaded From: https://mechanicaldesign.asmedigitalcollection.asme.org on 04/07/2019 Terms of Use: http://www.asme.org/about-asme/terms-of-use 8 2", " However, the gear tooth stiffness is changed due to the influence of the crack. When the tooth crack is present, the cross-section height of the faulty tooth, Hx, can be calculated as [17]: Hx = { 2y, x \u2264 xA yB\u2212yA (xB\u2212xA)2 (x\u2212 xA)2 + yA + y, x > xA (1) where (x, y) are the coordinates of the tooth profile, (xB, yB) are the coordinates of the tooth tip point B, and (xA, yA) are the coordinates of the crack tip point A, which can be calculated as follows: xA = xC \u2212 lc cos\u03b1c (2) yA = yC \u2212 lc sin\u03b1c (3) where (xC, yC) are the coordinates of the initial crack point C (see Figure 6). Here, a parabolic curve between the crack tip and the tooth profile tip is chosen as the limiting line for reducing the tooth thickness, as proposed by Mohammed et al. [18]. The geometric parameters of the fault define the gear mesh stiffness alteration as a function of the mesh angle. The effects of the four tooth root crack cases on mesh stiffness are shown in Figure 7. It can be seen that the stiffness is reduced with the increase in the crack length. A single stage spur-gear transmission was used in this study" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002700_s12541-021-00510-4-Figure14-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002700_s12541-021-00510-4-Figure14-1.png", "caption": "Fig. 14 Tooth profile overlapping interference inspection by the inverse method", "texts": [ " Assuming that the gear and the ring gear are normally engaged, the gear rotates at 1 , whereas the ring gear rotates at 2 . Now, taking the center of the fixed coordinate system Sg as the rotation center and the ring gear node circle as the rotation line, the gear and the ring gear simultaneously rotate at 2 in the reverse direction. Due to the same rotation angle, the relative position relationship remains the same. If interference occurs, the gear tooth tip enters the ring gear tooth profile (Fig.\u00a014). However, the critical point for interference lies at position H when the gear rotates around \u22152(Fig.\u00a010). In order to simplify the calculation, the check angle was set as 1 \u2208 [0,\u2212 ] , and the calculation method is presented in Fig.\u00a015. (1) The initial value was set as 1= 0 , and the calculation step was \u0394 1. (2) The calculation step (clockwise rotation) was subtracted. (3) The gear tooth profile after the rotation was calculated. 1 3 (4) The gear was reversed by taking O2 as the center and the ring gear node circle as the rotation line" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001952_s00170-020-05540-2-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001952_s00170-020-05540-2-Figure3-1.png", "caption": "Fig. 3 Front view (left) and top view (right) of Kennametal Stellite Excalibur Torch", "texts": [ " It is integrated with a 3-axis motion linear drive table, which allows for the production of more complex AM designs [13]. The integrated system is able to control most of the parameters in this process, such as movement direction, shield gas, powder gas and center gas flow rates, powder feed rate, current, voltage, and coordinated travel speed. The commercial PTA system is a Kennametal Starweld 400 PTA Welding system, with a Kennametal Stellite Excalibur Torch. An image of the torch is presented in Fig. 3 where the nozzle of the center gas, powder gas, and shield gas are presented alongside the non-consumable tungsten electrode. It is important to note that the torch provides a shielding atmosphere around the powder being deposited; however, it is localized to the area directly beneath the torch. The PTA-AM geometry is created by defining a set of parameters, from the aforementioned controllable parameters, by inputting them into the machine\u2019s control pendant. The control pendant is the interface between the user and the PTA-AM system" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001644_j.conbuildmat.2020.121961-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001644_j.conbuildmat.2020.121961-Figure2-1.png", "caption": "Fig. 2. Steel wire length relationship.", "texts": [ " The spiral strand cable consists of one central circular straight wire and three outer layers helically wound in opposite directions over one another around the central wire, as shown in Fig. 1. This section assumes a locked coil wire rope with a nominal diameter of 22 mm as an example to establish a refined finite element model. The parameters of the model are listed in Table 1. When the rotation angle reaches 360 , that is, the side wire rotates one circle from the starting point of the screw, the corresponding cable length is the lay distance P (shown in Fig. 2): P \u00bc 2pRi tanai \u00f01\u00de where ai is the twist angle of the i-layer steel wire. Riis the radius of the layer where the steel wire is located, Ri \u00bc r1 \u00fe 2r2 \u00fe \u00fe 2ri 1 \u00fe ri. ri is the radius of each layer of steel wire, and pi is the length within the unit lay length of the i-layer steel wire. In this study, a refined model that considers the interaction between steel wires is established, and thus a solid element is used for modelling. The steel wire model was built using the Rhinoceros software. The geometry of the core was obtained by linear z-axis extrusion, and the helical wires were generated by extrusion of circular areas along the helical curves corresponding to the centre axis of the wire" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001469_tmag.2020.3019821-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001469_tmag.2020.3019821-Figure7-1.png", "caption": "Fig. 7. No-load flux distribution and flux density. (a) NSNS. (b) DF-NSNS. (c) N/Fe/N/Fe. (d) DF- N/Fe/N/Fe.", "texts": [ " The optimized parameters of four machines are listed in TABLE 1. ms and mr are the number of stator and rotor phases, l is the stack length, Dso is the stator outer diameter, g is the air gap length, hm is the magnet thickness, ksp is the split ratio, ksso and krso are the stator and rotor slot opening ratio, kstt is the stator tooth width to pole pitch ratio, ksyst and kryrt are stator yoke to stator The comparison of no-load flux distribution and flux density of four optimized machines is shown in Fig. 7. For N/Fe/N/Fe and DF-N/Fe/N/Fe with ferromagnetic pole shoes, it can be seen that the pole flux leakage is smaller than the NSNS and DF-NSNS counterpart. Moreover, for DF-NSNS and DF-N/Fe/N/Fe with two sets of windings, it can be noted that they have thicker stator tooth and thinner rotor yoke compared to the NSNS and N/Fe/N/Fe, since more flux goes through one stator tooth owing to the additional rotor armature windings and smaller rotor yoke thickness saves space to place the rotor armature windings" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000123_s1560354719050071-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000123_s1560354719050071-Figure4-1.png", "caption": "Fig. 4. The system on the plane.", "texts": [ " 5 2019 Suppose that the frame and the rotors execute the following motion relative to the shell: \u2014 the frame rotates with angular velocity \u03a9(t), given as a function of time, about the axis of dynamical symmetry of the shell. If the shell is homogeneous, then any straight line passing through the geometric center of the shell can be chosen as the axis of rotation of the frame; \u2014 the rotors rotate (relative to the frame) with angular velocity \u03c6\u0307i(t), given as a function of time, about its axis of dynamical symmetry ni. We define two coordinate systems (see Fig. 4): \u2014 an inertial coordinate system OXY Z with origin at some point of the plane and with the axis OZ perpendicular to it. \u2014 a noninertial coordinate system Cx1x2x3, which is attached to the frame, so that the origin C coincides with the center of mass of the system. A distinctive feature of this system is the fact that the mass distribution remains constant in the coordinate system Cx1x2x3. This is due to the fact that the rotors and the shell in this system rotate in a prescribed manner about their symmetry axes (in contrast to the coordinate system attached to the shell)" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002175_icra40945.2020.9197401-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002175_icra40945.2020.9197401-Figure5-1.png", "caption": "Figure 5. (a) Pouring DragonSkin 10 silicone into the mold of the first layer of the FRA, (b) fiber wrapping, (c) pouring DragonSkin 10 into the mold of the second layer of the FRA, (d) completed FRA, (e) pouring DragonSkin 10 into pressure chamber mold, (f) integration of the cured pressure chamber with the alignment pins and a pressure pipe, (g) assembly of the SCS parts, (h) pouring DragonSkin 30 silicone into the mold of encapsulation layer, (i) completed encapsulation layer, (j) gluing FRAs, the SCS and the encapsulation layer, (k-l) capping and pressure pipe integration.", "texts": [ " The fabrication consists of four steps: i) fabrication of the FRAs; ii) preparation of the SCS and its encapsulation layer; iii) integration; iv) capping. After fabrication, the manipulator is properly assembled as shown in Fig. 4. A. FRA fabrication Each FRA is fabricated in 3 steps: i) first silicone layer fabrication; ii) fiber winding; iii) second silicone layer fabrication. The A and B parts of the DragonSkin 10 were mixed equally by weight and degassed to eliminate entrapped air bubbles. The mixture was poured into the first mold (Fig. 5a) for the first layer of the FRA. When the silicone is cured 697 Authorized licensed use limited to: University of New South Wales. Downloaded on September 20,2020 at 15:57:05 UTC from IEEE Xplore. Restrictions apply. completely, 0.3 mm inextensible nylon fiber was wound with 1.5 mm pitch (Fig. 5b). Then, it was placed in the second mold and the second layer of the DragonSkin 10 was poured (Fig. 5c). The process was repeated one more time to fabricate the second FRA (Fig. 5d). B. SCS fabrication To fabricate pressure chambers for the SCS, the prepared DragonSkin 10 was poured into the pressure chamber molds and a straight mold was placed at the middle for the inflation channel (Fig. 5e). Two alignment pins and a pressure pipe were glued to the cured chamber by silicone glue (Fig. 5f). These processes were repeated two more times to fabricate three identical pressure chambers for joints. The completed chambers were placed into the appearing rigid compartments by the integration of the rigid links (Fig. 5g). To fabricate the encapsulation layer, the prepared DragonSkin 30 was poured into the encapsulation layer molds and a second mold part was placed in the center to leave a space for the SCS (Fig. 5h). After the layer is cured (Fig. 5i), the SCS is placed into the dedicated opening. The separately fabricated FRAs (Fig. 5d), the SCS (Fig. 5g), and the encapsulation layer (Fig. 5i) were assembled (Fig. 5j). As the final step, the prepared DragonSkin 30 was poured into a capping mold, then both sides of the assembled manipulator were dipped into the mold one by one. The pressure pipes for the FRA actuation were inserted and fixed by silicone glue (Fig. 5k-l). The manipulator was characterized in terms of compliance, stiffness capability and workspace. All tests were performed in the x-y plane, with reference to Fig. 6, by compensating the gravity. For clarification, the pressurized regions are named as ri (i = 1, 2, 3, 4, 5). While the r1 and r2 represent the FRAs, the r3, r4 and r5 represent the joints of the SCS from proximal to distal part. To actuate the manipulator, only one (r1) of the two FRAs is pressurized at a time. On the other hand, the joints of the SCS are pressurized together with the same pressure to demonstrate the maximum stiffness contribution of the SCS to the manipulator" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001403_j.ymssp.2020.107075-Figure10-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001403_j.ymssp.2020.107075-Figure10-1.png", "caption": "Fig. 10. Rolling ball bearing with double half-inner rings and pre-damage.", "texts": [ " OirOb can be calculated through law of cosines, thus the vector rbr can be obtained when the pre-damage is added. As described in [23], the deformations at the fault edges can affect the contact characteristics between the ball and races. However, this issue is ignored in this paper. The experimental setup is showed in Fig. 9. The rotor is supported by an angular contact ball bearing (7205AC) and a rolling ball bearing with double half-inner rings (Table 3) in which the pre-damage is on the inner-right half ring showed in Fig. 10. Fig. 11 illustrates the supporting structure of the rotor system as well as force transfer path. The vibration responses appear when the rotor rotates under different speeds and axial forces. The axial forces were applied through the axial loading device and measured by the force sensor. The vibration responses were measured by Kistler accelerometer 8702B100M1 (sensitivity: 50 mV/lm, measuring range: \u00b1100 g, response bandwidth: 54.0 kHz). The vibration signals were measured with high frequency range so that the full spectrum can be examined and then it is found that the resonance frequencies due to bearing defects were dominant in frequency range (3000 Hz\u20135000 Hz)", " The frequency ratio is applied to reveal the severity of bearing skidding by means of the comparison of actual frequency ratio fr0 and the calculated value fr. Besides, a condition indicator g of bearing skidding is also introduced as g \u00bc f BPFI f 0BPFI f BPFI 100% \u00f028\u00de In general, it is indicated that the bearing skidding may appear when the condition indicator g is nonzero or there are deviations between f 0BPFI and f BPFI. Fig. 12 shows the time-domain waveform and envelope spectrum of vibration response from experiment results. The impact signal (caused by the pre-damage of half inner ring showed in Fig. 10) can be clearly observed from the time- domain waveform. The rotating frequency and bearing pass frequency of inner ring and corresponding double frequency can be easily obtained from the envelope spectrum of vibration response. Then the frequency ratio fr0 and condition indicator g in actual operating conditions can be calculated to compare with the simulation results. As is demonstrated Figs. 13 and 14, the simulation and experimental results of frequency ratio, and condition indicator are respectively in reasonable agreements, which validates the proposed dynamic model of rolling ball bearing with double half-inner rings", " Therefore, it is an effective way to prevent three-point abnormal contact of the bearing when the axial load is sufficient and stable in the actual operation of the bearing. The experimental scheme is based on rotor test-bed similar to the experimental setup showed in Fig. 9, where the difference is that the radial loading device is applied in this section and the structural diagram is seen in Fig. 23. From this figure, the bearing applied in this experiment (Table 3), with pre-damage on inner-right half ring showed in Fig. 10, is used to withstand radial and axial load. The same method (illustrated in Section 3) is applied to analyze the vibration signals. In addition, the contact force of the ball and the ring in the contact ellipse cannot be measured in the actual operating condition. Therefore, the vibration acceleration amplitude of the bearing pass frequency of inner ring, obtained in this experimental analysis, is used to predict the abnormal contact phenomenon, that is, the occurrence of the bearing pass frequency of inner ring means that the three-point contact appears" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002470_j.ijfatigue.2021.106154-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002470_j.ijfatigue.2021.106154-Figure1-1.png", "caption": "Fig. 1. The geometry and local finite element meshes of the tested fatigue specimens in the present work. (a) The fatigue specimens containing a small recasting zone. (b) The schematically illustrated recasting zone with finite element meshes.", "texts": [ " The wrought nickel-based superalloy Inconel 718 (IN718) manufactured by ThyssenKrupp, Germany was investigated in the present work, which was solution-treated following the standard ASTM B637 [28]. The specific nominal chemical compositions of the recasting and wrought IN718 are listed in Table 1. Plenty of Nb-enriched precipitates distribute in the recasting zone, which is considered as the Laves phases [29]. Recast IN718 in Table 1 represents the recasting material of Inconel 718 excluding the Laves phase. The fatigue specimens depicted in Fig. 1(a) were designed and tested to identify the differences in the fatigue performance between the recast-contained components and the popular base IN718. Partial fatigue specimens were processed by the conventional 500 W millisecond Nd:YAG laser machine to implement a recasting zone in the center region. The geometry of the recasting zone depends on laser manufacturing parameters, especially the average power and the pulse width [30]. Hence, after a series of preliminary attempts, the laser processing parameters were set as the average power P = 500 W, the pulse width d = 30 ms, the defocus L = 0, the pulse frequency f = 10 Hz and the duty cycle 100%", " Since fatigue life mainly depends on the peak and valley values, the validity and accuracy of the constitutive model are further verified. In summary, the constitutive model proposed in the present study provides a sound description of the cyclic plastic behavior of the base material and a reasonable assessment of the fatigue performance of the recasting material. To quantify the effect of material degradation on the fatigue performance of laser manufactured specimens, fatigue tests of the specimens illustrated in Fig. 1 were conducted under the load-controlled uniaxial tensile loading condition with the loading ratio R\u03c3 = 0.1. Although the whole specimens were under uniaxial, the recasting material in the specimens was under a multiaxial stress state due to nonuniform material property. The recasting material is much weaker than the base material, so the fatigue failure started in the recasting zone. The fracture of all recast-contained specimens is located on the symmetric plane of the recasting zone, which is perpendicular to the fatigue loading direction", " However, it must be clear that material inhomogeneity makes the central part of the specimen in the multiaxial stress state, which is coupled with the constraint effect [40] and gradient plasticity [41,42]. Hence, to further elucidate the failure mechanism of the recastcontained specimen, it is quite crucial to identify the local stress and strain states in the recasting zone. Finite element computations of the recast-contained specimen under different uniaxial tensile stress amplitudes with the ratio R\u03c3 = 0.1 were performed. The constitutive model discussed in the previous sections was used in the computations. The finite element meshes of the recasting zone are illustrated in Fig. 1(b). To identify the failure mechanism of the recast-contained specimen, the stress and strain distributions along the depth direction on the center axis of the recasting zone under stabilized peak loadings are shown in Fig. 10. y represents the loading direction, z stands for the depth direction of the recasting zone, and x conforms to the right-hand rule. The coordinates are illustrated in Figs. 1 and 2. Under the present uniaxial tension-compression fatigue tests, the stress and strain distributions T" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001761_tasc.2020.2968043-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001761_tasc.2020.2968043-Figure3-1.png", "caption": "Fig. 3. Magnetic flux lines distribution in proposed CMG.", "texts": [ " The magnetization direction in each PM segment is given by \u2192 M = Mr \u2192 r +M\u03b8 \u2192 \u03b8 (1) where Mr and M\u03b8 are Mr = \u221e\u2211 n=1,3,5 Mrn(n) cos[np(\u03b8 \u2212 \u03b80)] (2) M\u03b8 = \u221e\u2211 n=1,3,5 M\u03b8n(n) sin[np(\u03b8 \u2212 \u03b80)] (3) where Mrn and M\u03b8n are Mrn(n) = 4Br \u03bc0\u03c0 1 n sin (n\u03c0 2m ) \u00b7 { 1 + m\u2211 k=2 cos [ (k \u2212 1)\u03c0 m ] cos [ (k \u2212 1)n\u03c0 m ]} (4) M\u03b8n(n) = 4Br \u03bc0\u03c0 1 n sin (n\u03c0 2m ) \u00b7 { m\u2211 k=2 sin [ (k \u2212 1)\u03c0 m ] sin [ (k \u2212 1)n\u03c0 m ]} (5) where m is the number of PM segments per pole, p is the number of pole pairs, K is the kth piece of PM segments per pole, \u03b8 is the angle between the x-axis and the centerline of each PM segment, and Br is the remanence of the PMs. In order to verify the performances of CMG with unequal Halbach arrays, the FEM-based model is established and analyzed. The main geometrical parameters are listed in Table I. Fig. 3 shows the magnetic field distribution of the proposed CMG calculated by FEA. Fig. 4 shows the radial and tangential magnetic flux density waveforms in the inner air-gap of the CMG with unequal Halbach arrays. It can be found that the radial magnetic flux density waveforms in the inner air-gap of the proposed CMG are closer to a sinusoidal waveform. Fig. 5 shows the harmonic spectra of the inner air gap flux density. It can be seen from Fig. 5 that some harmonic magnetic density values are greatly suppressed" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002035_j.measurement.2020.108224-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002035_j.measurement.2020.108224-Figure6-1.png", "caption": "Fig. 6. Three-dimensional model of centrifugal fan.", "texts": [], "surrounding_texts": [ "According to the Table 7, the average stress is 55.4169 MPa at the rated speed. According to S-N curve of Fig. 5d, the damage of the fan is too small to make the blade appear fatigue state when it runs at uniform speed. In this paper, the overload acceleration of continuing start-stop experiment is done in fatigue testing [20]. The experimental parameters and results are shown in Table 8 and stress curves in a period of time on the 55th day are shown in Fig. 10. Two sets of strain data are unavailable due to measurement channel failure. Zero drift occurs in one of the channels like what is shown in curve 4 of lower part of Fig. 10, but the size and shape of the curve are the same as others. Therefore, it is also a normal signal for the following analysis. Based on these parameters and results, the blade fatigue is analyzed and verified from three aspects as follows. Firstly, according to the previous S-N mixed cyclic load strengthening theory, the formula can be obtained as Eq. (2). r1 is original stress and r2 is the overload condition stress. The stable operation stage of the last 5 test cycles are shown in Fig. 10. Zero Fig. 9. Position of resistance strain gauge. drift occurs in one of the channels as is shown in curve 4, but the time correspondence and shape of this curve are the same as others. Thus, it is also a normal signal and the average maximum value is 146.26 MPa among all curves. According to Fig. 5d, N1 is slightly greater than the fatigue limit 107 of steel. The overload coefficient m \u00bc 2:64 is close to 2.6 that is maximum allowable overload percentage. Because the blade material is steel, k \u00bc 6:5 7 is the recommended value. N2 \u00bc 6:06 109. Compared N1 and N2, the cycles number of accelerated fatigue test meets the demand of fatigue limit, so the cycles number under rated speed greatly exceeds the fatigue limit. N2 \u00bc N1 r1 r2 k \u00bc N1 mk \u00f02\u00de Secondly, according to the use requirements, the times of startstop is 36,000 per year (100 times a day and 360 days a year) when the motor vehicle fan runs at 1440 rpm. The expected service life of the fan is 297,720 times for the service life is expected to be 8.27 years according to the requirements of service conditions Table 7 Average test results of blade stress (MPa) at different speeds (r/min). 1440 1600 1800 2000 2200 55.4169 81.0751 92.0646 121.1321 138.3195 Table 8 Fatigue test parameters. Working condition Parameters per cycle Steady speed Operation time/cycles Maximum stress Number Continuous start-stop 5 s acceleration, 4 s steady state, 5 s deceleration and 1 s stop 2200 (r/min) 55 days/316,800 146.26 MPa 4 (36,000*8.27 = 297,720). The times of start-stop has exceeded the designed even if rotational speed is increased by 52.7%. This shows that the fan meets the expected service life requirements. At last, through the above experiments, it is confirmed that the blade is verified by simulation and test. In fatigue experiment, due to the limited measuring points, the fatigue phenomenon of other n to blades after fatigue test. parts may be missed. The final step is to check whether fatigue will occur in other parts. The blades are metallographically examined in accordance with the procedure shown in Fig. 11. In addition to randomly selecting 6 blades from the above 12 measuring points, the other 6 blades are randomly selected for measurement. No cracks are found in the stress concentration (ear root) of the blade after careful inspection. So, no fatigue is picked up through metallographic examination. Through the mutual confirmation of the above three methods, fan blades have passed strict fatigue test. It meets relevant industry standard and service life requirements." ] }, { "image_filename": "designv11_5_0002707_s0263574721000795-Figure9-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002707_s0263574721000795-Figure9-1.png", "caption": "Figure 9. Influence of a robots encounter with terrain on the CPG output.", "texts": [ " The position and velocity of the CoM, relative to the origin of the inverted pendulum, are given by x(t) = x0 cosh (kt) + x\u03070 1 k sinh (kt) (18) x\u0307(t) = x0k sinh (kt) + x\u03070 cosh (kt) (19) where k = \u221a g/h, g is the gravitational acceleration, h is the height of the CoM above the ground, and x0 and x\u03070 are the position and the velocity of the CoM relative to the origin of the inverted pendulum at t = 0, respectively. Feedback loops are designed to make the CoM move forward along the slope and to reduce the step length during up-slope walking to maintain motion stability. In this work, the body attitude state \u03b8pitch is used as input sensory information to the trajectory generator to generate slope-adaptive trajectories. Figure 9 shows that according to the presented feedback path, the robot adopts the output of its various layers as the standard timing modulated signal when crossing the terrain. Thus, using the output signals of the pattern formation layer, according to the state of the robot for an upward and downward slope, an online compensation strategy is designed as follows: CoMx_new = \u23a7\u23a8 \u23a9 { CoMx + Feed, \u03b8pitch > 0, PF_u\u03b8 = 2 CoMx \u2212 Feed, \u03b8pitch < 0, PF_u\u03b8 = 2 CoMx, PF_u\u03b8 = 0 (20) Considering that the main change the robot detects during the uphill and downhill movement is the body attitude angle in the y direction and that the attitude angle in the x direction changes only minimally, the design strategy of the feedback via CoM trajectory is as follows: Feed = Kfeed \u00d7 \u03b8pitch_DC, Kfeed > 0 (21) where Kfeed is the feedback gain of the feedback term in the CPG model, and \u03b8pitch_DC is the filtered body posture angle" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002230_ecce44975.2020.9236388-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002230_ecce44975.2020.9236388-Figure3-1.png", "caption": "Fig. 3: (a) A CAD model of the test bench with the temperature measuring point Tr. (b) The stator front view and the considered temperature measuring points Tuv, Tvw, Twu, Tbi, and Tair. (c) The back plate view with the positions of Tf and Tbo.", "texts": [ " Regarding Rig, it could be estimated by either approximating the equivalent length of the interface gaps between stator yoke and stator fixture or, again, by means of a dedicated DC test. For the considered machine the interface gap has been estimated to be approximately 30 \u00b5m. In order to verify the analytically derived thermal model, a test setup featuring an accurate measurement of the machine\u2019s temperatures is essential. In general, the transient temperature characteristics depend on the dynamic load profile that the machine is exposed to. Consequently, the experimental setup shown in Fig. 3a was designed for both the operation of the machine at different load points by controlling speed and torque, and a simultaneous acquisition of different temperatures within the machine. Its main components include the device under test (A), a hysteresis brake to apply various load points (B), a sensor to acquire speed and torque (C), and the power electronics for DC/AC conversion featuring a controller (D) for facilitating a field-oriented control of the machine to adjust the load point operation. Table II summarizes the key figures of the utilized torque-speed sensor and hysteresis brake. The transient temperatures of the individual motor components are determined with sensors at various positions of interest. The internal structure of the motor, as well as its back with an attached mounting flange are illustrated in Fig. 3b and 3c, respectively. For characterizing and monitoring the main thermal paths from different heat sources through the machine\u2019s components to the ambient, the following temperatures are acquired by thermal sensors: \u2022 Tuv, Tvw, Twu at the surface of one coil per phase; \u2022 Tair at the tip of one pole shoe of the stator yoke; \u2022 Tbi on the outer (static) ring of the inner ball bearing; \u2022 Tbo on the outer (static) ring of the outer ball bearing; \u2022 Tf on the mounting flange made of aluminum; \u2022 Tr on the outer side of the rotor. Their spatial positions inside the motor are illustrated in Fig. 3. On the one hand, all temperatures except Tr can be measured with conventional thermocouples of type K, which are connected to the relevant surface by using a thermally conductive adhesive. Instead, the outer rotor temperature has to be measured by a contactless measuring method. For this purpose, a thermographic camera is used. Because of the given topology, the magnet\u2019s temperature cannot be directly measured. In Fig. 4, sample data of one exemplary measurement cycle are presented. A stepwise random but reasonably bounded change of the load characterized by the respective torque and speed is considered" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002356_978-3-030-50460-1-Figure17.5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002356_978-3-030-50460-1-Figure17.5-1.png", "caption": "Fig. 17.5 Comparison of prescribed distributed loads (left halft) respectively displacements (right half) and consecutive structural response", "texts": [ " The external work in the static problem can be denoted as follows Wext,stat = 1 2 \"uT \"\u0302 F = \u22121 2 \"\u0302uT \"FR; (17.20) For a single point load or point displacement, the external energy as well as the stress and strain fields of both analysis modes would match for concordant prescribed loads and displacements in the linear elastic regime. It is to be noted that the distribution of local stresses and strains can be significantly different between the modes for distributed loads or displacements, due to the differences in local stiffness of the structure, as is illustrated in Fig. 17.5. The nonlinear analysis should validate the performance of the structure by overstraining in a disproportionate worst-case scenario. For this purpose, a preferably equivalent LBC is applied to both analysis modes but with an overloading characteristic in the nonlinear analysis. Initially, a time-dependent LBC is chosen such that the energy from external forces, and the strain energy of both analysis modes match in the linear elastic deformation regime. For this purpose, a prescribed displacement is applied to increase linearly with time (with a constant, but sufficiently small velocity to match the quasi-static condition)" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000403_0954405416640171-Figure12-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000403_0954405416640171-Figure12-1.png", "caption": "Figure 12. Tooth flank form error after simulation.", "texts": [ " The result shows that new tooth flank has good uniformity and smoothness, and on the fillet area, it also has some obvious changes on strain energy and can be consistent with the actual state. Tooth flank form error after optimization In previous modeling of the spiral bevel gear, the B-spline fitting method for the initial discrete points is one of the most main means of tooth flank construction. It was mentioned by Gabiccini et al and Artoni et al.37,38 Recently, Lin and Fong39 have conducted a detailed study. After the B-spline fitting with 37 3 33 control points, the normal error range is reduced to 100mm. As shown in Figure 12, it represents the tooth flank form error on concave side of gear model after the above simulation CNC-milling. With calculations between the initial data points and theoretical points, the average error can reach 264mm, the maximum error is 348mm and the minimum is 71.1mm. The distribution of data points is almost in the messy, unevenness and poor continuity mainly because of the cutting feed speed, the feed amount and sampling error and so on. With the NURBS uniform fitting of the spherical involute tooth profile and the entire tooth flank, the residual normal error is obtained, as shown in Figure 13" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002426_j.engfailanal.2020.105211-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002426_j.engfailanal.2020.105211-Figure7-1.png", "caption": "Fig. 7. Relationship between external raceway contact load and applied load.", "texts": [ "11 ni (9) Kn is the contact stiffness coefficient between the roller and the raceway, which is determined by the geometric parameters of the bearing and can be obtained from the following formula [22] Kn = 6.24 \u00d7 104l0.82 e D0.11 w [1 + c0.9 \u03bc cos(\u03b1v \u2212 \u03b1\u03bc)] \u2212 1.11 (10) le is the effective contact length of the roller, which can be determined by YANG et al [22]. le = lw/cos(0.5(\u03b1v \u2212 \u03b1\u03bc)) (11) c\u03bc = sin(\u03b1v + \u03b1f ) sin(\u03b1\u03bc + \u03b1f ) (12) \u03b1\u03bc is the contact angle between the roller and the inner raceway. \u03b1f is the contact angle between the roller-guiding flange and the guide edge of the inner ring [26]. The relation between the contact load of the outer raceway and the applied load is shown in Fig. 7. The components of the contact load in the radial and axial directions are as follows Qri = Qcos\u03b1vcos\u03d5i (13) L.-H. Zhao et al. Engineering Failure Analysis 122 (2021) 105211 Qai = Qsin\u03b1v (14) Above analysis only consider the axial displacement and radial displacement of bearing. For TRB, there is also an angular misalignment between the inner and outer rings, even if both bearing centers are alignments, roller will be pressure to the big guard to produce skewed torque. The concave curvature of the outer raceway prevents the roller from being skewed, and this resistance and subsequent deformation will change the load distribution of roller - raceway and roller - inner ring guide edge" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002700_s12541-021-00510-4-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002700_s12541-021-00510-4-Figure2-1.png", "caption": "Fig. 2 Coordinate system", "texts": [ " Finally, a real conjugated straight-line internal gear pair was machined and tested to verify the reliability of the proposed design method. Figure\u00a01 displays the typical structure of a linear conjugate internal gear pump, which is mainly composed of an upper shell, a lower shell, a gear, a gear ring, a transmission shaft, a seal ring, and other components. The gear, the ring gear, and the crescent are the key components to complete oil suction and drainage, and their cross-sections are exhibited in Fig.\u00a02. During operation, the motor drives the gear on the transmission shaft, consequently, the gear meshes with the ring gear and starts to rotate. At the oil suction port, the gear and the gear ring are removed, and the cavity volume in the oil suction area increases to form a vacuum to complete oil absorption. The oil between gear teeth is transferred to the oil outlet through the sealed transition cavity formed by the crescent plate. At the oil outlet, the gear and the ring gear become engaged, and the cavity volume decreases to remove oil. As the axial structure of the gear ring remains consistent, the numerical model could be established based on the cross-section shown in Fig.\u00a02. For the convenience of the numerical model, three coordinate systems were set up (Fig.\u00a03). The moving coordinate system S1(O1x1y1) , which is fixed with the internal gear 1 3 and rotates with it, takes the center of the gear as the origin. The line passing point P, the coincidence of the two pitch circles, is the Y-axis, and the X-axis is perpendicular to the Y-axis. The coordinate system S2(O2x2y2) takes the center of the ring gear as the origin, and the line passing point P is the Y-axis. The coordinate Sg(O2xgyg) is the same as S2(O2x2y2) ; however, the coordinate S2(O2x2y2) is fixed with the ring gear and rotated with it, whereas the coordinate Sg(O2xgyg) is stationary" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001738_j.mechmachtheory.2019.103629-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001738_j.mechmachtheory.2019.103629-Figure2-1.png", "caption": "Fig. 2. The studied manipulator in different configurations.", "texts": [ " Compared to the previous version the proposed solution features increased robustness and simplicity in the generation of motion in horizontal plane. In the proposed design, pairs of linkages are connected to diagonally opposite joint positions of the rhombus; however, any other arrangement that fully constrain the manipulated platform when all four linkages are locked is possible. The selected arrangement has the advantages of decoupling planar translations of the manipulated platform from the other two DOFs. As illustrated in Fig. 2 , the manipulated platform is a rhombus linkage, composed of four fixed-length links (green, yellow, blue and purple) connected by revolute joints. As can be seen in Fig. 2 (a), the purple platform section controls the rotation angle of the end-effector. The red platform is connected to the purple link via a prismatic joint and to the green link by a helical joint. By keeping the purple section fixed while modifying the shape of the rhombus, the vertical position of the manipulated platform is changed. Fig. 2 (b) and (c) demonstrate the maximum and minimum range of vertical translation. As can be seen, the vertical range is small; however, it is sufficient for the applications such as cutting applications as mentioned earlier. Fig. 2 (c) and (d) demonstrate a 90 \u25e6 rotation of the manipulated platform with constant position. The main advantage of the proposed design is that infinite rotation of the manipulated platform is possible. The mechanism is targeting applications requiring large translations in horizontal plane; small range of vertical translations, and infinite rotation around the vertical axis. The advantage of the proposed mechanism over other Sch\u00f6nflies motion generator manipulators such as Quadrupteron [39] or 4-DOF Delta parallel manipulator [5,44] could be listed as: \u2022 This mechanism has larger workspace in horizontal plane" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000250_iros40897.2019.8968050-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000250_iros40897.2019.8968050-Figure2-1.png", "caption": "Fig. 2. Description of the biasing of the CdTe detector and the interaction of gamma-ray photons at different distances of the electrodes of the detector.", "texts": [ " Although this dependency might be partially corrected with proper calibration techniques, the low slope of the shaped signal at low threshold voltage levels introduces large jitter in the trigger signal, which cannot be corrected. Note that the longer the shaper\u2019s peak time and the higher the order of the integrator, the larger the jitter of the trigger independently of the threshold voltage reference. The architecture of Fig. 1 offers limited time resolution when the shaper\u2019s peak time is designed of few s. Based on the proposed design for PET scanner [9], the photon flux is parallel to the CdTe detector\u2019s electrodes as illustrated in Fig. 2. In this particular case, the probability for every photon to interact close to the cathode (case a) is equal to the probability to interact close to the anode (case c) or half way (case b), and it is independent of the energy of the photon. Therefore, the degradation of the collection efficiency due to the limited mobility of holes in cadmium telluride detectors plays a fundamental role in the design of the electronics and the biasing conditions. 0018-9499 \u00a9 2015 IEEE. Translations and content mining are permitted for academic research only", ", the distance to the cathode/anode of the photon interaction in the proposed detector design) in 2-mm thick CdTe detectors concludes that the total charge of electrons and holes must be integrated to achieve the energy of 511 keV with 1% resolution at FWHM [10]. For CdTe detector at room temperature, the ratio of the electrons mobility ( cm V s) to that of holes mobility ( cm V s) is around 11 [11], [12]. Considering that the mobility ratio increases inversely with temperature, the timings of the pixel front-end electronics have been designed to provide an energy resolution of 0.1% at 511 keV regardless the total charge collection time, i.e., cases a, b, or c of Fig. 2. II. VIP-PIX ASIC AND PIXEL ARCHITECTURE The front-end electronics of every pixel has been designed and optimized for positron emission tomography scanners based on the multiple stacking of pixelated Cadmium Telluride detectors such as the described in [9]. Fig. 3 shows the block diagram of the pixel front-end electronics. As an alternative to the high energy-resolution front-end architecture shown in Fig. 1, we connected the unfiltered signal coming from the pre-amplifier to the discriminator to provide a fast trigger with small time-walk and less dependent on the depth of interaction" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001588_j.matchemphys.2020.124022-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001588_j.matchemphys.2020.124022-Figure4-1.png", "caption": "Fig. 4. Speculation on the formation of crater and striation morphology with 107 W/cm2 and 108 W/cm2.", "texts": [ " The edges gradually rise to form a crater morphology. With 107 W/cm2, the surface formed the crater morphology. With 108 W/cm2, the surface appeared stripes. Some scholars have explained the formation mechanism of striation morphology as pyrolysis [21], photolysis [22], and oxidative deposition [23]. Based on the formation of crater morphology, forming the striation morphology is inferred. As E increases, the amount of gasification increases, and the recoil pressure increases, forming a deeper crater. As shown in Fig. 4, the actual morphology obtained is different from the ideal morphology. The edges in the Y direction are not completely solidified and are obscured by the next pulse spot, resulting in a smaller convex. The edges in the X direction are completely solidified before the next processing. The X-direction has increased because of the cumulative effect of multiple processing. Therefore, the striation morphology is formed. B. Liu et al. Materials Chemistry and Physics 259 (2021) 124022 As shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000187_s10846-019-01113-y-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000187_s10846-019-01113-y-Figure7-1.png", "caption": "Fig. 7 a A special case of misjudgment and its solution b Modification of Radius (n = 12)", "texts": [ " Self-collision and collision against stations can be solved by comparing the minimum distance between corresponding axes of cylinders with the sum of radii, as is shown in Fig. 6. The collision margin is the minimum distance after the deduction of radii. It is worth mentioning that these solutions are conservative and introduce additional allowance. This allowance is beneficial for prediction of self-collision and collision against obstacles because the radii of links are much smaller than their length. Yet, space stations have much bigger radius causing the allowance unacceptable. Figure 7a demonstrates a special case. Here, the smaller cylinder representing the link clearly does not collide with the larger one representing the station. But the distance between two axes is smaller than the sum of radii. In order to solve this problem, the envelope of station should be used in analysis. The premise of this case is that the projection of link\u2019s axis is partially or completely outside the station\u2019s axis. The station\u2019s envelope can be discretized by a group of segments as is shown in Fig. 7a. And due to the discretization, the radius of link should also be modified larger for compensation. The worst situation where the link and the station are tangent at the middle of two discrete segments, depicted in Fig. 7b, should be considered. This modification can be deduced from triangular law of cosines. rlink modify 2 = rstat ion 2 + (rstat ion + rlink) 2 \u22122 cos(\u03c0/n)rstat ion(rstat ion + rlink) (20) Herein, rlink modify is the modified radius. rstat ion/rlink is the radius ratio. n stands for the number of segments. In this paper, n is set to be 30 empirically. Distances between the axis of a link and the group of segments can be calculated. The collision margin is the minimum among these distances after the deduction of the modified link\u2019s radius" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001717_jmes_jour_1959_001_016_02-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001717_jmes_jour_1959_001_016_02-Figure7-1.png", "caption": "Fig. 7. Components in which the taroidal element may play a part", "texts": [ "3= 1 dp/dO = 0 at O = f 4 2 (that is, essentially unit F with zero me). In Fig. 6 similar curves are shown for a O = 0 and + ~ / 2 Vol I No 2 19S9 at UNIV OF VIRGINIA on June 5, 2016jms.sagepub.comDownloaded from 120 C. E. TURNER toroidal element between limits 8 = 0 and +m/2. In all cases the behaviour of the shell is obviously very markedly dependent upon the value of p. A common mechanical engineering component in which the toroidal element may play an important part is the flatplate expansion bellows, Fig. 7. This device is often used to absorb the thermal expansion of a pipeline conveying hot fluids or gases, and the meridional radius of m a t u r e of the root or crest of the convolution is normally so much smaller than the radius of revolution that the approximation q+cos 8 -rr 7 is quite acceptable. The stresses in such a bellows subjected to an axial tensile load are shown in Fig. 8. In this solution the symmetry of the problem leaves two constants for each toroidal part to be determined, together with two for the bending and two for the stretching of the flat-plate parts* by matching the boundary conditions at the inner and outer edge of the plate" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001934_j.mechmachtheory.2020.103895-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001934_j.mechmachtheory.2020.103895-Figure1-1.png", "caption": "Fig. 1. Leg of the Delta robot.", "texts": [ " [1] : n O = n B \u2212 n J \u2212 n R \u2212 n F (1) The previous expression can be easily verified remembering that joints and rigid connections couple two boundary nodes and express the nodal displacements of one boundary node in terms of those of the other: it means that only one of the two boundary nodes can be selected as an independent node. Finally, each fixed node can be directly removed from the computation of Eq. (1) . To better explain this result, the Delta robot [55,56] , whose one of the three legs is shown in Fig. 1 , will be used as a demonstrative example considering three models with decreasing complexity, evaluated in terms of the number of dofs . The first model 1 shown in Fig. 2 is a fully flexible model in which all links are deformable. In model 2 the short sides of the parallelogram are rigid while in model 3 only the forearms D and arms P are flexible while the remaining parts are rigid. All boundary nodes have been marked with blank circles. The end-effector node E has been added to the set of boundary nodes", " In this section, the proposed method is applied to describe the elastodynamic behavior of the Ragnar robot developed by the Blue WorkForce company ( https://bwf3.tindwork.dk/ ). The Ragnar robot is an LMPM designed to accomplish Sch\u00f6nflies motions. By replacing the moving platform, it is possible to transform the Ragnar robot into a pure translational robot [66] . In its translational version for pick-and-place tasks, displayed in Fig. 3 , the four legs of the Ragnar robot have the same topology of the Delta robot shown in Fig. 1 . The geometric parameters have been reported in Fig. 3 . The reader is referred to [66\u201369] for the complete list of geo- metric and inertial parameters, here recalled in Tables 1 and 2 for convenience. As reported in Wu et al. [66] , the actuator stiffness has been set to k a = 50 , 0 0 0 (Nm/rad). Changing the joints of the MP, for example, to model the 4-DoF Ragnar robot, would involve only a change of the H matrices for those joints without modifying the stiffness/inertia matrices of the MP", " The submatrices H B i and G B i refer to the i th-leg and are defined as H B i = \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 h R b i 0 0 0 0 0 0 h R b i . . . 0 0 . . . 0 0 . . . 0 0 . . . h R \u03c0i . . . 0 0 . . . h R \u03c0i . . . 0 0 . . . 0 0 . . . 0 0 . . . h R \u03c0i 0 0 . . . h R \u03c0i 0 . . . 0 \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 , G B i = \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 O 6 . . . O 6 \u02dc 1 \u02dc 1 \u02dc 1 \u02dc 1 \u02dc 1 \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 (33) where O 6 is the 6 \u00d7 6 dimensional zero-matrix. By introducing the 6 \u00d7 6 dimensional identity-matrix 1 6 and referring to the unit vectors e b i and e \u03c0i of the i th leg displayed in Fig. 1 , we define h R b i = [ 0 e b i ] , h R \u03c0i = [ 0 e \u03c0i ] , \u02dc 1 = [ 1 6 1 6 ] (34) and H B v = \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 0 . . . 0 h R b 1 . . . 0 0 . . . 0 0 h R b 2 . . . 0 0 . . . 0 0 . . . h R b 3 0 0 . . . 0 0 . . . h R b \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 , G B v = \u23a1 \u23a2 \u23a3 \u02dc 1 \u02dc 1 \u02dc 1 \u02dc 1 \u23a4 \u23a5 \u23a6 (35) 4 These matrices are used to derive the three transformation matrices of the reduced models presented in Section 3 . In order to take into account the actuator stiffness k a applied at the base revolute joint of each leg, the expression of B appearing in Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002336_indicon49873.2020.9342485-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002336_indicon49873.2020.9342485-Figure1-1.png", "caption": "Fig. 1: Quadrotor illustration with inertial frames and 6 DOF coordinates. \ud835\udf14\ud835\udc56 is the angular speed and Fi is the propeller force of the ith rotor.", "texts": [], "surrounding_texts": [ "978-1-7281-6916-3/20/$31.00 \u00a92020 IEEE\nProportional Integral Derivative (PID), a classical controller, and Linear Quadratic Regulator (LQR), an optimal controller, acting on a quadrotor model to control its translational and rotational movements. The complicated dynamic model of a quadrotor based on Newtonian and Euler's laws and fundamental physics principles is simplified to derive a linear mathematical model. The derivation gives the equations that govern a quadrotor's motion concerning the body frame and the inertial frame in the state-space form. This linearized model is utilized to develop the controllers. The performance of the PID control and the optimal regulator established on LQR is compared in various conditions subjected to the Quadrotor model.\nKeywords\u2014 Proportional Integral Derivative, Linear\nQuadratic Regulator, Quadrotor, Simulink\nI. INTRODUCTION\nQuadrotors (or unmanned aerial vehicles powered by four rotors) are finding increasing applications in fields ranging from video surveillance to emergency response. They are required to fly in unknown environments and essentially with accurate control. Consequently, the controllers for quadrotors have been studied, researched, and improved over the period. Most controllers used on quadrotors utilize classical control, such as proportionalintegral-derivative (PID) control.\nHowever, classical control theory encompasses mostly linear time-invariant single-input single-output systems. The basis of control for these systems depends on how their behavior is modified using a feedback loop. On the other hand, the Linear Quadratic Regulator (LQR), a type of optimal control whose theory focuses on mathematical optimization of an objective cost function for a dynamic system and multi-input and multi-output systems and is expected to be more robust for a quadrotor. LQR focuses on non-linear models rather than the classical linear equation approach of PID. PID controllers' main drawback is that the actual system requires its linearization in every test and new condition. However, for LQR control, this step is eliminated, and system equations can be directly fed to the controller and acquire the desired response.\nShahida et al. [1] had explored how PID controllers gave stability by bringing the closed-loop poles to the more negative side of the s-plane and compared to LQR controllers. Hayrettin et al. [2] tackled quadrotor control's\nproblem by proposing an adaptive controller that is a hybrid of both PID and LQR controllers. Meera et al. [3] presented results on a PID controller coupled with a Kalman filter being applied to UAVs. Lucas et al. [4] worked on using LQR control to tune a PID controller, combining the control techniques to control the quadrotor.\nExisting research in the field has stated how PID has effectively been implemented for control in simulations and even real-world scenarios. However, when extended to scenarios with considerable disturbances to a quadrotor's flight or during starting up from no load, the PID controller gains have to be tuned continuously. This motivated to study and apply optimal control to comprehend it\u2019s effect on such uncertain circumstances.\nThis study aims to apply both classical PID controller and optimal controller based on LQR for the quadrotor translational and rotational motions and compare their performances. This work can be extended to other unmanned aerial vehicles such as fixed-wing aircraft and vertical takeoff or landing (VTOL) vehicles, which opens up a wide avenue of applications.\nThe following sections presents the modeling and simulation, followed by the results. Section II details the Quadrotor Model, i.e., the quadrotor's physical model and layouts the dynamic equations governing its motion. These equations are linearized, and a state-space model is developed in Section III. The IV and V sections follow up with a comprehensive study and observations of the PID and LQR controllers. The results section contains various cases for flight simulation using the two controllers. These are compared, and the conclusion is presented.\nII. QUADROTOR MODEL\nThe position of a quadrotor in space can be specified using two reference frames: fixed and mobile. The fixed (inertial) coordinate system is one where Newton\u2019s laws are valid, and it is with respect to the Earth. The second frame is with respect to the quadrotor frame center, where the flight controller is positioned. The quadrotor has four Degrees of Freedom, namely Thrust, Roll, Pitch, and Yaw. The quadrotor's position and movement are controlled by varying the speeds and rotation direction, of the four motors (two diagonally opposite motors spin clockwise and the other two motors spin anti-clockwise). However, the quadrotor has four inputs (4 motors) and six outputs (three translational and 20 20 IE EE\n1 7t\nh In\ndi a\nC ou\nnc il\nIn te\nrn at\nio na\nl C on\nfe re\nnc e\n(I N\nD IC\nO N\n) | 9\n78 -1\n-7 28\n1- 69\n16 -3\n/2 0/\n$3 1.\n00 \u00a9\n20 20\nIE EE\n| D\nO I:\n10 .1\n10 9/\nIN D\nIC O\nN 49\nAuthorized licensed use limited to: Raytheon Technologies. Downloaded on May 20,2021 at 09:56:48 UTC from IEEE Xplore. Restrictions apply.", "three rotational motions along the three axes), and hence, it is an underactuated nonlinear complex system.\nThe Euler angles \u03c6 \u2208 [\u2212\u03c0, \u03c0], \u03b8 \u2208 [\u03c0/2, \u03c0/2], \ud835\udf13 \u2208 [\u2212\u03c0, \u03c0] are used to give the orientation of the mobile reference frame with respect to the inertial frame. These angles help form a rotation matrix that is used to represent rotation/orientation in the same way position is used to represent displacement vectors. The non-linear equations of the quadrotor system are given as [5]:\n?\u0308? = (\ud835\udc50\ud835\udc5c\ud835\udc60\ud835\udf11\ud835\udc60\ud835\udc56\ud835\udc5b\ud835\udf03\ud835\udc50\ud835\udc5c\ud835\udc60\ud835\udf13 + \ud835\udc60\ud835\udc56\ud835\udc5b\ud835\udf11\ud835\udc60\ud835\udc56\ud835\udc5b\ud835\udf13) \ud835\udc39\n\ud835\udc5a\n(1)\n?\u0308? = (\ud835\udc50\ud835\udc5c\ud835\udc60\ud835\udf11\ud835\udc60\ud835\udc56\ud835\udc5b\ud835\udf03\ud835\udc60\ud835\udc56\ud835\udc5b\ud835\udf13 \u2212 \ud835\udc60\ud835\udc56\ud835\udc5b\ud835\udf11\ud835\udc50\ud835\udc5c\ud835\udc60\ud835\udf13) \ud835\udc39\n\ud835\udc5a\n(2)\n?\u0308? = \u2212\ud835\udc54 + (\ud835\udc50\ud835\udc5c\ud835\udc60\ud835\udf11\ud835\udc50\ud835\udc5c\ud835\udc60\ud835\udf03) \ud835\udc39\n\ud835\udc5a\n(3)\n?\u0307? = \ud835\udc3c\ud835\udc66\ud835\udc66 \u2212 \ud835\udc3c\ud835\udc67\ud835\udc67\n\ud835\udc3c\ud835\udc65\ud835\udc65\n\ud835\udc5e\ud835\udc5f \u2212 \ud835\udc3d\ud835\udc5f\n\ud835\udc3c\ud835\udc65\ud835\udc65\n\ud835\udc5e\ud835\udf14 + \ud835\udc622\n\ud835\udc3c\ud835\udc65\ud835\udc65\n(4)\n?\u0307? = \ud835\udc3c\ud835\udc67\ud835\udc67 \u2212 \ud835\udc3c\ud835\udc65\ud835\udc65\n\ud835\udc3c\ud835\udc66\ud835\udc66\n\ud835\udc5d\ud835\udc5f \u2212 \ud835\udc3d\ud835\udc5f\n\ud835\udc3c\ud835\udc66\ud835\udc66\n\ud835\udc5d\ud835\udf14 + \ud835\udc623\n\ud835\udc3c\ud835\udc66\ud835\udc66\n(5)\n?\u0307? = \ud835\udc3c\ud835\udc65\ud835\udc65 \u2212 \ud835\udc3c\ud835\udc66\ud835\udc66\n\ud835\udc3c\ud835\udc67\ud835\udc67 \ud835\udc5d\ud835\udc5e +\n\ud835\udc624 \ud835\udc3c\ud835\udc67\ud835\udc67\n(6)\nwhere Jr is the rotor inertia and p, q, and, r represents the angular velocities with respect to roll, pitch and yaw motions respectively. The input forces are given as:\n\ud835\udc621 = \ud835\udc3e\ud835\udc53 \u2217 (\ud835\udf141 2 + \ud835\udf142 2 + \ud835\udf143 2 + \ud835\udf144 2)\n(7)\n\ud835\udc622 = \ud835\udc3e\ud835\udc53 \u2217 (\ud835\udf144 2 \u2212 \ud835\udf142 2)\n(8)\n\ud835\udc623 = \ud835\udc3e\ud835\udc53 \u2217 (\ud835\udf141 2 \u2212 \ud835\udf143 2)\n(9)\n\ud835\udc624 = \ud835\udc3e\ud835\udc5a \u2217 (\ud835\udf141 2 \u2212 \ud835\udf142 2 + \ud835\udf143 2 \u2212 \ud835\udf144 2)\n(10)\nwhere \ud835\udc3e\ud835\udc53 is the thrust factor and \ud835\udc3e\ud835\udc5a is the drag factor whose value depends on propeller size and air conditions in which the quadrotor is flying, \ud835\udf14\ud835\udc56 is the angular speed of the ith rotor. The quadrotor parameters considered for simulation are given in table 1.\nThe LQR optimization problem requires a linear state-space system as the A and B matrices along with the Q and R\noptimal control matrices are necessary for computing the full state feedback K matrix. The following section focusses on modelling the non-linear quadrotor equations as a linear state-space system.\nIII. STATE-SPACE MODELING\nThe quadrotor state-space system consists of 12 state variables, four input variables, and four output variables. The state variables represent the absolute quadrotor orientation in space consisting of linear and rotational coordinates and their respective velocities. The input variables consist of the four quadrotor motions, namely thrust, roll, pitch, and yaw motion. The output consists of the required state variables for stability analysis of the quadrotor, which are the vertical displacement (z), roll (\ud835\udf11), pitch (\ud835\udf03), and yaw (\ud835\udf13) angle displacements.\n?\u0307? = \ud835\udc34\ud835\udc4b + \ud835\udc35\ud835\udc48\n(11)\n\ud835\udc4c = \ud835\udc36\ud835\udc4b + \ud835\udc37\ud835\udc48\n(12)\n\ud835\udc4b\ud835\udc47 = [\ud835\udc65 \ud835\udc66 \ud835\udc67 \ud835\udf11 \ud835\udf03 \ud835\udf13 ?\u0307? ?\u0307? ?\u0307? \ud835\udc5d \ud835\udc5e \ud835\udc5f]\n(13)\n\ud835\udc48\ud835\udc47 = [\ud835\udc621 \ud835\udc622 \ud835\udc623 \ud835\udc624]\n(14)\n\ud835\udc4c\ud835\udc47 = [\ud835\udc67 \ud835\udf11 \ud835\udf03 \ud835\udf13]\n(15)\nwhere A is a 12\u00d712 state matrix and B is a 12\u00d74 input matrix which are given by [5]:\n\ud835\udc34 = [ 0 0 0 1 0 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 0 0 0 \ud835\udc621 \ud835\udc5a 0 0 0 0 0 0 0 0 0 0 \u2212 \ud835\udc621 \ud835\udc5a 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 1 0 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0]\n\ud835\udc35 = [ 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 1 \ud835\udc5a 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 1 \ud835\udc3c\ud835\udc65\ud835\udc65 0 0 0 0 1 \ud835\udc3c\ud835\udc66\ud835\udc66 0 0 0 0 1\n\ud835\udc3c\ud835\udc67\ud835\udc67]\n(16)\n(17)\nC matrix is a 4\u00d712 matrix with each column having a unity multiplier to scrape out the 4 state parameters from X, as mentioned above. D is a 4\u00d74 null matrix. Once the statespace equations are obtained, the controller is designed for the quadrotor system. The following section describes the action of a PID controller on the Quadrotor system.\nAuthorized licensed use limited to: Raytheon Technologies. Downloaded on May 20,2021 at 09:56:48 UTC from IEEE Xplore. Restrictions apply.", "IV. PROPORTIONAL INTEGRAL DERIVATIVE CONTROL\nThe PID controller is one of the standard classic models in control theory. The aim is to minimize the error by tuning the proportional, integral, and the derivative coefficients used in the controller equation. Error is defined as the difference between the setpoint value and the actual controller output at any particular instance. The proportional term increases/decreases the error by multiplying with a proportionality constant. The derivative term is used to estimate the controller's future response based on the error with respect to time. The derivative constant also accounts for the damping factor and needs to be tuned to minimize damping and obtain a smooth response. The integral term in the controller sums up all the past values, and the integral coefficient is used for assigning the weight to the integral term to obtain the desired output response in the shortest possible time. Thus, all three coefficients need to be tuned to obtain the setpoint value of the control system.\nThe PID control equation uses the \ud835\udc3e\ud835\udc5d, \ud835\udc3e\ud835\udc56, and, \ud835\udc3e\ud835\udc51 parameters and the error value to compute the controller response. The control equation is given below:\n\ud835\udc48(\ud835\udc61) = \ud835\udc3e\ud835\udc5d\ud835\udc52(\ud835\udc61) + \ud835\udc3e\ud835\udc56 \u222b\ud835\udc52(\ud835\udc61)\ud835\udc51\ud835\udc61\n\ud835\udc61\n0\n+ \ud835\udc3e\ud835\udc51\n\ud835\udc51\ud835\udc52(\ud835\udc61)\n\ud835\udc51\ud835\udc61\n(18)\n\ud835\udc3e\ud835\udc51 and \ud835\udc3e\ud835\udc56 are defined as:\n\ud835\udc3e\ud835\udc51 = \ud835\udc3e\ud835\udc5d\ud835\udc47\ud835\udc51 & \ud835\udc3e\ud835\udc56 = \ud835\udc3e\ud835\udc5d/\ud835\udc47\ud835\udc56 (19)\nwhere \ud835\udc47\ud835\udc51 and \ud835\udc47\ud835\udc56 are the time periods for integral and derivative actions. The error value is defined as:\n\ud835\udc52(\ud835\udc61) = \ud835\udc60\ud835\udc52\ud835\udc61\ud835\udc5d\ud835\udc5c\ud835\udc56\ud835\udc5b\ud835\udc61 \ud835\udc63\ud835\udc4e\ud835\udc59\ud835\udc62\ud835\udc52 \u2212 \ud835\udc4e\ud835\udc50\ud835\udc61\ud835\udc62\ud835\udc4e\ud835\udc59 \ud835\udc63\ud835\udc4e\ud835\udc59\ud835\udc62\ud835\udc52 (20)\nIn the Quadrotor system, the state parameters need to be controlled so that it remains stable during the flight period. A small deviation in the quadrotor's setpoint values would create large variations in motion, which would lead to deviation from its desired trajectory and making it difficult for the user to control. Even if the main control lies in the hand of the user flying the quadrotor, unavoidable disturbances during the flight time will create instability and malfunction of the system. Thus, there is a need for an internal flight controller, which would autonomously adjust the rotor to a stable position. The four inputs thrust, roll, pitch, and yaw motion control the six-state parameters depending on the action. Thrust controls the \ud835\udc67 parameter, roll controls the \ud835\udc65 and the \ud835\udf03 parameters, pitch controls \ud835\udc66 and \ud835\udf11 parameter and yaw controls the \ud835\udf13 angle [6]. Thus, the six-\nstate parameters require 6 PID controllers with 1-2-2-1 controllers for thrust, roll, pitch, and yaw motion control, respectively. For roll and pitch motion, a cascaded PID loop is designed to have an inner and an outer feedback loop. The inner loop needs to be faster than the outer loop for an efficient response of the controller. Slowing the outer loop increases the overshoot but reduces the setting time for the rotational coordinates and their velocity and increases the time for the corresponding translational coordinate to attain the set point. The tuned PID controller parameters for optimal control are given in table 2 below.\nAfter the successful implementation of the PID controller, the optimal LQR controller is designed for the linearized quadrotor model.\nV. LINEAR QUADRATIC REGULATOR\nAs compared to the classic PID controller, the LQR controller involves a lot of mathematical computations to calculate the full state feedback matrix K. LQR controller uses an optimal control algorithm to minimize the cost function defined by the system equations. The cost function involves the state parameters and the input parameters to the system along with the Q and R matrices. The overall cost function needs to be minimum for optimal LQR solution. The Q and R matrices represent the weights assigned to the state parameters and the input parameters. By varying the values of the two matrices, the total value of the cost function can be adjusted according to the desired output. The two main quantities that need to be optimized for the quadrotor model are the power consumption and response speed. For a faster response of the controller, the Q matrix values need to be changed, whereas for minimizing the power consumption while achieving the desired setpoint without focusing on the time of response, the R matrix values need to be adjusted [8].\nFig.2: The closed loop quadrotor system with PID controller for stabilization\nFig.3: The closed loop quadrotor system with LQR controller for stabilization\nAuthorized licensed use limited to: Raytheon Technologies. Downloaded on May 20,2021 at 09:56:48 UTC from IEEE Xplore. Restrictions apply." ] }, { "image_filename": "designv11_5_0001575_j.measurement.2020.108685-Figure16-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001575_j.measurement.2020.108685-Figure16-1.png", "caption": "Fig. 16. Planet gear with a broken tooth.", "texts": [ " Therefore, the output voltage of the system can be treated as the strain of the ring gear. From the perspective of fault diagnosis, the absolute magnitude of the strain has less significance, and what we expect is to find the difference between the normal strain and the abnormal strain. Accordingly, the replacement of strain with output voltage is reasonable. 4.1 Low speed condition In this condition, the input speed of the planetary gearbox is set to be about 70 RPM. (1) Tooth-broken fault of planet gear The planet gear with a broken tooth is shown in Fig. 16. In this condition, only one planet gear is used in the gearbox. It has been indicated in Section 2 that when the faulty tooth of planet gear meshes with the ring gear tooth near the measurement point, the tooth fault feature results from the variation of the ratio of ring-planet mesh force distributed on each tooth of the ring gear. Accordingly, the tooth fault feature of planet gear is supposed to be seen even in the one-planet system. 16 Before using the strategy in Section 3, the raw strain (voltage) signal is low pass filtered first to remove the quasi-static component caused by the bulk temperature of the ring gear and then is resampled to the angular domain based on the phase signal offered by the Hall sensor on the axle of carrier" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001800_978-3-030-40410-9-Figure6.5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001800_978-3-030-40410-9-Figure6.5-1.png", "caption": "Fig. 6.5 Circular obstacles detection", "texts": [ " The simple way is to apply a static penalty on the solution vector that forms a graph which touches an obstacle. This penalty is enough to get rid of that solution and force the optimization algorithm to propose another feasible one. We will show how to detect a collision path and apply the penalty in the following example. In this section, collision detection between path and obstacles will be described in detail. Assume that we have a path consisting of successive points pi, i = 1 \u2026 n where n is the total number of points on the path as shown in Fig. 6.5. Also, there are Oj obstacles with radius rj, j = 1\u2026 k where k is the total number of obstacles in the workspace. A collision will occur when the distance between pi and the centre of the obstacle Oj is less than the radius rj. 6.5 Collision Detection 131 132 6 Path and Trajectory Planning d j = \u221a( pxi \u2212 Ox j )2 + ( pyi \u2212 Oyj )2 , (6.2) d j < r j , (6.3) where x and y denote the x-axis and y-axis of pi and Oj, respectively. For rectangular obstacles, any point pi on the path should not be contained within the space occupied by the rectangular obstacle" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002614_j.compositesa.2021.106449-Figure10-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002614_j.compositesa.2021.106449-Figure10-1.png", "caption": "Fig. 10. Process of determining tow-tow intersections.", "texts": [ " The intersected tows are computed by determining whether it contains the intersected mesh, obtained by projecting down current tow points on the mesh assembly of preceding tows using a ray-tracing algorithm. The ray-tracing algorithm which determines tow intersections, however, will detect edge-edge intersection along the entire length of parallel tows. To avoid these spurious intersections and ignore very small tow interpenetrations, a small tolerance (1/6 tow width) is set by using inner tow points only (red dots shown in Fig. 10) to detect tow intersection. The tolerance is essentially computed by the distance between the outermost inner points and edge points, which can be further optimized by using bias or increasing the number of transverse tow points. The general process of determining the tow-tow intersection edges is presented in Fig. 10. For every identified intersected tow, only the X. Li et al. Composites Part A 147 (2021) 106449 interpolation functions at edges are used to compute the intersections with current tow (7 interpolation curves). The intersection edge is represented by a series of determined intersections, calculated by a \u2018fsolve\u2019 function in the SciPy library. This function is used to approximately determine the intersections between high order interpolation curves and works for both linear and nonlinear cases. A small numerical tolerance 1x10\u2212 6 was used in the current study", " These and other complex intersection cases are treated separately and require a more general zoffset algorithm with a slight drop in geometric fidelity of overlap features. A compromise method, presented in detail in Section 3.6, is employed to describe the defect zone. The simple and complex intersection cases can be distinguished by checking the constituents and characteristics of intersection edges. For the simple intersection case (Fig. 11a), two intersection edges can be defined for each intersected tow, and each edge will have a computable number of interpolated intersection points (see Fig. 10). If these intersection edges cross they are removed as per Fig. 12. Other cases where the intersection edge is not computable, for example, there can be 0 (Fig. 11b) or less than 7 internal intersections (Fig. 11d) along an edge, are also removed. For the remaining simple tow intersection edges, the intersection points are duplicated to create a prismatic and regular towdrop around the overlapped region, according to the procedure shown in Fig. 12. The point duplication parameters are determined based on a desired tow-drop angle and the tow thickness" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000536_j.mechmachtheory.2016.08.008-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000536_j.mechmachtheory.2016.08.008-Figure7-1.png", "caption": "Fig. 7. The RCCC linkage and local Cartesian reference frames definition according to Denavit convention.", "texts": [ " 13.2. Kinematic analysis of the RCCC four-bar linkage Thanks to the Principle of Transference, the steps of this kinematic analysis are similar to those required for the spherical four-bar linkage. A recent kinematic analysis of the RCCC linkage was reported by Figliolini et al. [48]. Once the definitions of Eq. (79b) are taken into account, the analysis equations are the dual counterparts of those already established for the spherical case. A pictorial view of the spatial linkage is shown in Fig. 7. The dual transform is now introduced \u03c6 \u03c6 \u03c6 \u03c6 = \u00b1 \u2212 \u2212 ( ) \u03b8 \u03b8 \u03b8 \u03b8 \u2212 \u2212 \u2212 \u2212 \u23a1 \u23a3 \u23a2 \u23a2 \u23a2 \u23a2 \u23a4 \u23a6 \u23a5 \u23a5 \u23a5 \u23a5 Q i i e cos 2 e sin 2 e sin 2 e cos 2 , 97 k i k i k i k i k /2 /2 /2 /2 k k k k its dual loop closure equation being expressed as = ( ) Q Q Q Q I . 981 2 3 4 The velocity and acceleration relations are obtained upon differentiating both sides of (98) with respect to time. In particular, the relations below are derived when Eqs. (92) and (95) are taken into account \u03b8 \u03b8 \u03b8 \u03b8\u0307 + \u0307 + \u0307 + \u0307 = ( ) H Q HQ Q Q Q Q HQ Q Q Q Q HQ 0 991 2 1 2 3 4 3 1 2 3 4 4 1 2 3 4 with 0 denoting the 2 2 zero matrix" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001644_j.conbuildmat.2020.121961-Figure14-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001644_j.conbuildmat.2020.121961-Figure14-1.png", "caption": "Fig. 14. Tensile force of a strand wire.", "texts": [ " Therefore, as the lay angle increases, the axial strain of the locked coil wire rope increases. In the process of wire twisting, if the cable lay length is large, steel wire skipping will occur. Therefore, the cable lay length should be increased within a certain range. Benndorf [16] was the first to work out a partition of the wire rope tensile force in wire tensile force. The calculation of tensile stress of the spiral strand in the current study is based on the research of Feyrer [17]. The geometric relationship in the process of steel wire stretching is shown in Fig. 14. and Fig. 15.. A partition of the tensile force in wire tensile forces must be calculated [17]. The wire tensile force component in the strand axe direction (neglecting the small shear force) is Si \u00bc Ficosai Ui \u00bc Fisinai : \u00f02\u00de The tensile force of the wire layer i is Fi \u00bc Dli li EiAi; \u00f03\u00de where li is the wire length, Dli is the wire elongation, Ei is the elasticity modulus, and Ai is the cross-section of a wire in wire layer i. The extension of the wire is ei \u00bc Dli li : \u00f04\u00de With lS respect to the length of the strand, the length of the wire is li \u00bc lS cosai ; \u00f05\u00de Therefore, when the failures of higher classification are neglected, the wire elongation is Dli \u00bc \u00f0DlS Dui tanai\u00decosai: \u00f06\u00de The contraction of the winding radius, that is, the circumference in relation to the wire extension that transverses the contraction ratio, can also be designated as \u2018Poisson\u2019s ratio\u2019 of the wire helix as follows: t \u00bc Dui=ui Dli=li ; \u00f07\u00de Dui \u00bc t ui Dli=li; \u00f08\u00de Dui \u00bc t ui Dli=li; \u00f09\u00de Dui \u00bc t li sinai Dli=li: \u00f010\u00de On this basis, the elongation of a wire in the wire layer i is: Dli \u00bc DlS cosa 1\u00fe tsin2a : \u00f011\u00de This equation, together with Eqs" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002603_s00170-021-07057-8-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002603_s00170-021-07057-8-Figure3-1.png", "caption": "Fig. 3 Sets of holes in different forms, for which parameter research is carried out. a Cylindrical, model 1. b Wide oval, model 2. c Drop, model 3. d Hexagonal, model 4. e Narrow oval, model 5. f Pentagonal, model 6", "texts": [ " Catia V5 \u00a9 Part Design module was used in the design process. The study aims to reduce the production-induced sag in the upper part of the geometry in microchannels (CCC) produced without support in the DMLS production system, with production parameters. In this direction, the research process begins with the production of hole sets designed in cylindrical, wide oval, drop, hexagonal, narrow oval, and pentagon forms in the DMLS method. Five different scales of each microchannel form are included in the hole sets (Fig. 3). In the first model, in a cylindrical shape (Fig. 3a), hole diameters were determined as 1 mm, 2 mm, 4 mm, 6 mm, and 8 mm, respectively. In other models (Fig. 3b, c, d, e, f), the holes in different forms scaled equally with the hole circumference in the cylindrical shape will be evaluated within the scope of these forms\u2019 usability research micro-coolers. After creating the CAD data of the hole sets consisting of six different geometries and five different scales, the amodel in the cylindrical form was first produced by the DMLS method with nine different production parameter sets. Parameter sets to reduce sagging in the internal channel geometry in non-support productions have been investigated", " By decreasing the laser power value from 340 to 100 W in the \u201cp6\u201d and \u201cp8\u201d parameter studies, in the \u201cp2\u201d and \u201cp3\u201d parameter trials, the thickness parameter was determined as 0.03 mm and 0.09 mm instead of 0.06 mm. The parameter defined in the scanning direction is determined as X instead of X-Y. In the \u201cp4\u201d parameter test, the value of overlap with inskin was determined as 0.04 mm instead of 0.02 mm; for the \u201cp1\u201d parameter, the minimum length value was chosen as 0.4 mm instead of 0.2 mm (Table 4). In the first stage of the study, the SEM images of the samples produced in nine parameters of five holes modeled in cylindrical form (Fig. 3a) were examined (Table 4), and the parameter with the least sagging was determined (Table 5). Ideal parameter sets were determined by SEM research of other hole sections produced in these parameters. Parameter sets are determined within the scope of hatch distance, laser power, thickness, scanning direction, overlap within the skin, and minimum length parameters. The numbers with black signs in Tables 4 and 5 indicate the changed parameter values. In this case, the hatch distance value, which is 0", " In the first stage of the study, the \u201cp6\u201d and \u201cp8\u201d parameters, which were determined ideally against the sagging problem, were applied to six different geometries where the variability was examined in the second stage of the study, and the results are shared in Table 7 and Fig. 6. The sagging rates of the holes of different sizes evaluated within the study\u2019s scope are shared graphically in Fig. 6. According to the study results, it was determined that the least deformation occurred in the production performed with the \u201cp6\u201d parameter (Table 6). For comparison purposes, \u201cp8\u201d and the standard default parameter \u201cp std\u201d parameters are also selected for the second part of the study. They are based on the production of hole sets in different forms seen in Fig. 3b\u2013f. Because it is a parameter in which the least deformation occurs and the laser power and scanning direction values changed in the parameters \u201cp6\u201d and \u201cp7\u201d are changed jointly, \u201cp8\u201d has been the parameter chosen for the second stage (Table 5). In the second stage of the study, it was determined that the \u201cp6\u201d parameter caused the least sagging in all the investigated geometries according to the SEM images obtained from the production results with \u201cp6\u201d, \u201cp8,\u201d and \u201cp std,\u201d which is the default parameter of the production system (Table 7)" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000550_s00170-016-9311-z-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000550_s00170-016-9311-z-Figure1-1.png", "caption": "Fig. 1 System diagram", "texts": [ ", v\u2032n as: v 0 1 \u00bc v1 v1k k; v 0 2 \u00bc v2\u2212 v2; v 0 1 v 0 1 v2\u2212 v2; v1h iv0 1 ; v 0 n \u00bc vn\u2212 vn; v 0 n\u22121 v 0 n\u22121\u2212\u22ef\u2212 vn; v 0 1 v 0 1 vn\u2212 vn; v 0 n\u22121 v0 n\u22121\u2212\u22ef\u2212 vn; v 0 1 v0 1 ; \u00f05\u00de Parameter Value Here, vki is equal to \u03b4xki, vector v\u2032kn is a normalized version of vki, and < , > signifies the inner product. The frequency of reorthonormalization is not critical, as long as neither the magnitude nor the orientation divergences have exceeded computer limitations. Then, the Lyapunov exponents can be calculated as \u03bbi\u2248 1 kh Xk k\u00bc1 ln v 0 k\u00f0 \u00de i ; i \u00bc 1; 2;\u22ef; n\u00f0 \u00de \u00f06\u00de Here, k is the number of total calculation steps, h is timestep size. The established coordinate system for unmanned aerial vehicles B(x, y, z) and geodetic coordinate system E(X, Y, Z) is given in Fig. 1. Assuming, quadrotor unmanned aerial vehicles are of a rigid body and four propeller axes are all perpendicular to surface of the vehicle body. Depending on the values of F1, F2, F3, and F4, four situations exist. When \u03a91 = \u03a92 = \u03a93 = \u03a94, the quadrotor unmanned aerial vehicles are rising, descending, or hovering; when \u03a92 = \u03a94 and \u03a91 \u2260 \u03a93, the system is pitching; when \u03a91 = \u03a93 and \u03a92 \u2260 \u03a94, the system is rolling; finally, when \u03a91 = \u03a93 \u2260 \u03a92 = \u03a94, the system is yawing. The dynamic model of quadrotor unmanned aerial vehicles based on Euler-Poincare equation are q\u0307 \u00bc V q\u00f0 \u00dep M q\u00f0 \u00dep\u0307 \u00fe C q; p\u00f0 \u00dep\u00fe F p; q; u\u00f0 \u00de \u00bc 0 \u00f07\u00de where p is a vector of quasi-velocities, q is the generalized coordinate vector, F(p, q, u) is a vector of externally applied generalized forces and the M(q) is the inertial matrix; V(q) is the Kinematics matrix; C(q, p) is the gyroscopic matrix" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002210_j.robot.2020.103670-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002210_j.robot.2020.103670-Figure7-1.png", "caption": "Fig. 7. Control of robot shoulder rotary motions by using the body. (a) Front view of the robot body, (b) Side view of the stopper on the robot\u2019s chest.", "texts": [ " In this paper, the length of the upper arm link (from Joint 2 shoulder) to Joint 4 (elbow)) is l2, while the length of the forearm ink (from Joint 4 to Joint 6) is l4C . The manipulator joint angles re \u221290 deg \u2264 \u03c62 \u2264 +90 deg and 0 \u2264 \u03c63 \u2264 +100 deg. All of the anipulators have hand mechanisms consisting of two fingers Joint 6) to enable them to hold items (Fig. 5). Touch sensors are et on the forearm links of each manipulator, which are covered y bumpers (Fig. 6). These touch sensors are used when the robot limbs a step (see Section 3). The robot also has a stopper mounted on the front of its body Fig. 7(a)), which makes it possible to limit the passive rotational ravel of the manipulators when pushing a wheelchair that is in he process of climbing a step (Fig. 7 (b)). This enables the robot o imitate the actions of a (Fig. 8) human transporting a heavy bject using a hand cart or wheelchair. When the robot pushes he wheelchair and assists in the step-climbing maneuver, the pper links of the manipulators press against the stopper on the obot\u2019s chest, which means it does not need to exert force around he shoulder axes. The wheelchair has a push handle mechanism on its back Fig. 9) that is equipped with a rotary shaft, thus allowing passive otation of the handle", " <2> The robot stops, and the wheelchair continues to drive orward until the accelerometer system detects the inclination t which the system is capable of placing the wheelchair\u2019s front heels on the step. <3> At this point, the wheelchair center of mass is in front f the contact point between the rear wheels and the ground. owever, it shifts to behind the contact point as the wheelchair ilt increases. The robot stopper limits the passive rotation travel bout Joint 2, thus preventing the wheelchair from tipping over ackward (Fig. 7(b)). <4> The wheelchair and the robot move forward, and the ront wheels of the wheelchair are placed on the step. w r b o w w p s w u t s 4 i s 1 H [Stage 2] <5> The wheelchair and the robot continue to move forward. <6> The rear wheels of the wheelchair come into contact ith the step. <7> The robot continues to push the wheelchair so that the ear wheels of the wheelchair keep in contact with the step and egin climbing. The robot continues to support the wheelchair in rder to prevent it from tipping over backward" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002614_j.compositesa.2021.106449-Figure11-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002614_j.compositesa.2021.106449-Figure11-1.png", "caption": "Fig. 11. Tow-tow intersections for simple case (a) regular cross intersection and complex cases: (b) transverse intersections; (c) multiple tows intersections; (d) partial intersections.", "texts": [ " The intersection edge is represented by a series of determined intersections, calculated by a \u2018fsolve\u2019 function in the SciPy library. This function is used to approximately determine the intersections between high order interpolation curves and works for both linear and nonlinear cases. A small numerical tolerance 1x10\u2212 6 was used in the current study. It should be noted that the number of points along each intersection edge depends on the number of interpolation curves of current tow, seven in this work. The tow-tow intersections can be a very simple case (see Fig. 11a), where only two tows are intersected in the middle portion without any geometric interference of other tows, creating small prismatic zones called \u201cresin-rich area\u201d. The pre-determined intersection edges are used to remap the tow points, which are utilized to precisely depict the prismatic zones for simple tow drop and overlap cases with the combination of a robust z-offset algorithm. While for other uncommon complex cases, precise geometric description of tow drops becomes extremely challenging, such as parallel or shallow angle tow intersections (Fig. 11b), multiple tows intersections (Fig. 11c), and partial intersections (Fig. 11d) near tow drops or part boundaries, etc. It should be noted that the three complex intersection cases primarily aim to demonstrate the difference and challenges of tow drop modelling with the comparison of the regular and well-intersected case (Fig. 11a). Any irregular intersection condition not categorised as regular will be handled by the general z-offset algorithm. These and other complex intersection cases are treated separately and require a more general zoffset algorithm with a slight drop in geometric fidelity of overlap features. A compromise method, presented in detail in Section 3.6, is employed to describe the defect zone. The simple and complex intersection cases can be distinguished by checking the constituents and characteristics of intersection edges. For the simple intersection case (Fig. 11a), two intersection edges can be defined for each intersected tow, and each edge will have a computable number of interpolated intersection points (see Fig. 10). If these intersection edges cross they are removed as per Fig. 12. Other cases where the intersection edge is not computable, for example, there can be 0 (Fig. 11b) or less than 7 internal intersections (Fig. 11d) along an edge, are also removed. For the remaining simple tow intersection edges, the intersection points are duplicated to create a prismatic and regular towdrop around the overlapped region, according to the procedure shown in Fig. 12. The point duplication parameters are determined based on a desired tow-drop angle and the tow thickness. The remaining tow points are re-mapped to avoid high mesh distortion, especially for oblique tow intersection angles. Another special case in the simple tow-tow intersection is if the intersection angle of the tows is very shallow, such as 5 degrees" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000539_s00170-016-9337-2-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000539_s00170-016-9337-2-Figure7-1.png", "caption": "Fig. 7 Meshes on platform surface", "texts": [ " There are handful optimization methods that can be used to solve such problems, like sequential quadratic programming (SQP), penalty and augmented Lagrangian methods, or genetic algorithms. In this research, we use C++ programming language and open-source library NLopt to solve the optimization problem. We provide an example of solving the step-over cost functional built on the blade platform surface shown in Figs. 4 and 5. The 3D meshes on the machining surface are generated by open source software Gmesh, as shown in Fig. 7. The algorithms solving nonlinear optimization problems usually have higher than O(n2) performance; although a denser mesh will have a better approximation of the surface, we avoid generating a dense mesh randomly. Instead, we set the tolerance of the meshing method according to the geometric tolerance and machining tolerance, we also set the upper bound of the number of triangles to be ten thousands (empirical value). The potential values of the vertices on the outer boundary L1 are fixed to zero, and the potential values of the vertices on the inner boundary L2 are fixed to eight" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0003067_bf02120345-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0003067_bf02120345-Figure4-1.png", "caption": "Fig. 4. The length of an element of the elastic prismatic rod replacing the helical spring in the deformed state.", "texts": [ " compressive force P from the increase of the internal spring energy E. This increase bE is equal to l 1 9 f M2dz f Q2dz b E = . ] 2a + .1 ~ ' (6) 0 \" 0 which can be computed for given deflection components v/and y with the aid of equations (2) and (4). As already stated, it is part ly supplied by the compressive force P, owing to the decrease b(A1) of the spring length. In the non-deflected state of the rod (spnng) the element dz has a length dl. The deflection causes the element to rotate as a whole through the angle ~v (see fig. 4), so that the distance at right angles between two adjacent cross-sections also equals dl. On the other hand,owing to the shear effect, the length ds of the element measured along the curved central line is ds = d//cos ~0 ~ dl (1 + 89 Furthermore ds = dz~/1 + (y,)2 m dz {i + 89 Thus the length of the spring decreases by the amount t t - - ~ ( A t ) l b(Al) = . f d l - fdz ~ ,_} f {(y,)2 __ 92} dz. (7) 0 0 0 The energy supplied by the compressive force is Pb(A1), and the amount of work W done by the components #L and q of the external load can be calculated from W = bE - - Pb(Al)" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000471_j.ijsolstr.2016.07.004-Figure20-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000471_j.ijsolstr.2016.07.004-Figure20-1.png", "caption": "Fig. 20. Elements of an annular membrane.", "texts": [ " (67) \u2013( 69 ) as \u03c31 \u03c30 = \u23a7 \u23a8 \u23a9 C 2 / r \u221a C 2 \u2212M 2 / r 2 r < R 2 C 1 a 2 / b 2 + 1 + \u221a C 2 1 + M 2 2 r > = R (67) r \u03c32 \u03c30 = { 0 r < R 2 C 1 a 2 / b 2 + 1 \u2212 \u221a C 2 1 + M 2 r 2 r > = R (68) \u00af = 3 M\u0304 8 ( 1 \u2212 a 2 b 2 )[ 1 R\u0304 2 \u2212 1 B + ln ( B A ) + 1 R\u0304 2 \u2212 8 3 ( a b )2 + 5 3 ] (69) here \u00af = 3 E\u03d5 4 \u03c30 ( 1 \u2212 a 2 b 2 ) (70) .2.1. Verification of accuracy In this sub-section, the annular membrane is analyzed with he new wrinkling model and the results are compared with he theoretical solutions obtained from Eqs. (63) to ( 65 ), ( 67 ) o (69). The finite element model is shown in Fig. 20 . Its outer adius b and inner radius a (shown in Fig. 19 ) are 1250 mm nd 500 mm, respectively. The thickness t is 1 mm. The Young\u2019s odulus E = 10 0 0 MPa and the Poisson\u2019s ratio \u03c5 =1/3. The inner dge is restrained in the radial direction and the circumferential OFs of nodes on the edge are coupled to have the same torsional otation. The outer edge is firstly pretensioned to introduce a u c t r r a m p M i a s c b p r o a p a w fl v t 6 m ( 0 8 As for the numerical examples in Section 6.2 , their geometrical and physical parameters are the same as given in Section 6" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000264_cbs46900.2019.9114526-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000264_cbs46900.2019.9114526-Figure6-1.png", "caption": "Fig. 6: WhiskEye robot platform and its operational environment with the trajectory overlaid in magenta.", "texts": [ " (1), | \u00b7 |L1 represents the L1 norm between the corresponding terms. \u03b1, \u03b2 and \u03b3 are scaling factors for the respective errors which represent the standard deviations of the raw sensory data. Finally, the returned match ID (m), is the ID of the combined template with the lowest \u03b5. In this section, we describe the operational environment and the robot platform that were used for empirical validation of the proposed ViTa-SLAM algorithm. A. Robot Platform The robot platform used for this research is called the WhiskEye (Fig. 6a) the design of which is based on a previous whiskered robot [19]. WhiskEye is composed of a Festo Robotino body, a 3 DoF neck, and a 3D printed head. The robot is ROS compatible which allows for candidate control architectures to be developed and deployed on either the physical platform or the Gazebo simulation (shown in Fig. 6b) of WhiskEye as used in this study. Mounted on the head are the visual and tactile sensors. Two monocular cameras with a resolution of 640\u00d7 480 pixels each provide a stream of RGB images with 5 frames per second. An array of artificial whiskers consisting of 24 macro-vibrissae whiskers arranged into 4 rows of 6 whiskers provides tactile information. Each whisker is equipped with a 2-axis hall effect sensor to detect 2D deflections of the whisker shaft measured at its base during, and is actuated using a small BLDC motor to reproduce the active whisking behavior observed in rodents. The tactile data from whiskers is extracted during every whisk cycle, which takes 1 second to complete. B. Operational Environment As a proof of concept, the algorithm was primarily tested in a simulated aliased environment to test visual-, tactile- and ViTa-SLAM under the challenging conditions a rodent faces in nature including: coarse vision, ill-lit tunnels, ambiguous visual and tactile environments. Fig. 6c shows the used simulated environment, a 6 \u00d7 6 m2 arena with 4 wallmounted visual and 3 tactile landmarks designed to be qualitatively equivalent to the natural environment. In this setting, the robot was made to revolve around the center of the arena, with a radius of 1 m whilst facing outwards to the walls. The following metrics were used to evaluate the performance of the proposed ViTa-SLAM against the VisualSLAM and Tactile-SLAM. 1) Localization Error Metric (LEM) The localization error metric (LEM) measures the root mean squared error (RMSE) between the true pose and the estimated pose where the error is calculated separately for position and orientation" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000020_978-3-030-12391-8_5-Figure5.3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000020_978-3-030-12391-8_5-Figure5.3-1.png", "caption": "Fig. 5.3 Mode 2 of the S4 Beam", "texts": [ " In seeking to obtain agreement between the simulations and the experimental measurements, three main variables were studied. The solver tolerances were varied to find the settings that provide the minimum solve time without sacrificing significant accuracy. The preload of the bolt was varied over multiples of the nominal value to span any uncertainty in the actual load. And finally, the measured surface contour was applied to the interfaces of the FEM to quantify the effect of small but realistic changes to the interface geometry on the results. In all models, mode 2 of the S4 Beam was used. Figure 5.3 shows that mode 2 is first order, in phase bending mode with motion in the Y-direction. When initially creating many of the models in this paper, convergence issues caused the models to take several days to solve and many did not converge at all. Computation time was also a topic of discussion in [20], which was a precursor to this work. Hence, the solver settings were explored further to seek to speed up the computations. The default solution control parameters defined in Abaqus/Standard are designed to provide reasonably optimal solution of complex problems involving combinations of nonlinearities as well as efficient solution of simpler nonlinear cases" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001800_978-3-030-40410-9-Figure1.2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001800_978-3-030-40410-9-Figure1.2-1.png", "caption": "Fig. 1.2 Watt and Stephenson mechanisms", "texts": [ " xvi About the Authors Chapter 1 Introduction The development of mechanisms has passed through different stages since the invention of the wheel; the first form of the mechanism was found in an archaeological excavation in Mesopotamia (modern Iraq), and dates to around 3500 BC (Postgate 2017) (Fig. 1.1). Many kinds of technical literature have been written, and many thoughts have been introduced from that time to develop different types of mechanisms, like Watt and Stephenson mechanisms which are shown in Fig. 1.2. Around the middle of the twentieth-century people began automating mechanisms to do a specific job, taking advantage of new opportunities due to the development of computer science. In the 1960s, these automated machines were unique devices called industrial robots (Craig 2005). Nowadays, robots play a significant role in all aspects of human life (Daneshmand et al. 2017; Spola\u00f4r and Benitti 2017) because of the human tendency to design products with low cost, high quality, and fast production rates" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002169_0142331220949366-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002169_0142331220949366-Figure1-1.png", "caption": "Figure 1. Nonlinear inverted pendulum and cart system.", "texts": [ " Furthermore, if (96) is satisfied, which implies that J (t)\\0, it remains to conclude that the closed-loop system (80)\u2013(82) with local nonlinear models is robustly stable respecting the H\u2018 performance level \u00a7c and the maximized admissible Lipschitz constant g. This completes the proof. Illustrative example In this section, we consider the inverted pendulum benchmark system (Sami and Patton, 2013), which consists of a movable carriage with a degree of freedom on which a pendulum is mounted and can rotate freely in the direction of travel (see Figure 1). The cart is driven by a motor which exerts a force through a belt transmission. Nonlinear inverted pendulum and cart system modelling A state-space model for the nonlinear pendulum and inverted carriage system can be expressed as _x1 t\u00f0 \u00de= x2 t\u00f0 \u00de, _x2 t\u00f0 \u00de= g sin x1 t\u00f0 \u00de\u00f0 \u00de m l a x2 2 t\u00f0 \u00de sin 2x1 t\u00f0 \u00de\u00f0 \u00de 2 ba cos x1 t\u00f0 \u00de\u00f0 \u00de x4 t\u00f0 \u00de 4 l 3 m l a cos x1 t\u00f0 \u00de\u00f0 \u00de2 a cos x1 t\u00f0 \u00de\u00f0 \u00de u t\u00f0 \u00de fc\u00f0 \u00de 4 l 3 m l a cos x1 t\u00f0 \u00de\u00f0 \u00de2 , _x3 t\u00f0 \u00de= x4 t\u00f0 \u00de, _x4 t\u00f0 \u00de= m g a sin 2x1 t\u00f0 \u00de\u00f0 \u00de 2 + 4 m l a 3 x2 2 t\u00f0 \u00de sin x1 t\u00f0 \u00de\u00f0 \u00de b a x4 t\u00f0 \u00de 4 3 m a cos x1 t\u00f0 \u00de\u00f0 \u00de2 4 a 3 u t\u00f0 \u00de fc\u00f0 \u00de 4 3 m a cos x1 t\u00f0 \u00de\u00f0 \u00de2 : 8>>>>>>>>>>< >>>>>>>>>>: where x1(t), x2(t), x3(t) and x4(t) represent, respectively, pendulum angle position, cart position, pendulum angular velocity and cart speed" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000654_smc.2016.7844562-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000654_smc.2016.7844562-Figure3-1.png", "caption": "Fig. 3. Forces and torques acting on the system", "texts": [ " If we assume that the local and global coordinate systems are perfectly aligned for (\u03c6, \u03b8, \u03c8) = (0, 0, 0), then any reference frame {o} vector po = [ pox poy poz ]T can be expressed in terms of the reference frame {g} vector pg = [ pgx pgy pgz ]T as pg = Rg op o, (4) where Rg o is the rotation matrix describing the total rotation between the local and global coordinate systems. Finally, if we assume the ZY X convention then the rotation matrix is given by Rg o = R(\u03c6, \u03b8, \u03c8) = R(Z,\u03c8)R(Y, \u03b8)R(X,\u03c6) = c\u03c8 c\u03b8 c\u03c8 s\u03b8 s\u03c6 \u2212 s\u03c8 c\u03c6 c\u03c8 s\u03b8 c\u03c6 + s\u03c8 s\u03c6 s\u03c8 c\u03b8 s\u03c8 s\u03b8 s\u03c6 + c\u03c8 c\u03c6 s\u03c8 s\u03b8 c\u03c6 \u2212 c\u03c8 s\u03c6 \u2212s\u03b8 c\u03b8 s\u03c6 c\u03b8 c\u03c6 . (5) The forces and torques acting on the system are shown in Fig. 3. There are only two types of forces acting on the system, i.e. the thrust T and the gravitational force G. By using a static approximation the rotor thrust is as follows [14]: T = 8\u2211 i=1 Fi = b 8\u2211 i=1 \u21262 i , (6) where b [Ns2/rad2] is the rotor thrust constant and \u2126i [rad/s] and Fi [N ] are the angular velocity and the thrust force applied SMC_2016 002183 by the i-th rotor, respectively. If we denote the octorotor mass with mo then the gravitational force can be expressed as G = mog, where g \u2248 9.81 [m/s2] is the gravitional acceleration of Earth. The thrust T acts in the Z direction of the local coordinate system and therefore can be written as the vector T o = [ 0 0 T ]T , while the gravitational force G acts in the ZB direction of the global coordinate system and can be written in the same manner as the vector Gg = [ 0 0 G ]T . Let us denote the torques acting around the X , Y and Z axes of the local coordinate system as \u03c4x, \u03c4y and \u03c4z , respectively (see Fig. 3). Then, we assume that the distances from the octorotor center of mass to the center of mass of all its rotors are equal and denoted by l. Now, the torques around the X and Y axes can be formulated as: \u03c4x = l ( F1 + \u221a 2 2 F2 + \u221a 2 2 F8 \u2212 F5 \u2212 \u221a 2 2 F4 \u2212 \u221a 2 2 F6 ) = bl ( \u21262 1 + \u221a 2 2 \u21262 2 + \u221a 2 2 \u21262 8 \u2212 \u21262 5 \u2212 \u221a 2 2 \u21262 4 \u2212 \u221a 2 2 \u21262 6 ) , \u03c4y = l ( F7 + \u221a 2 2 F6 + \u221a 2 2 F8 \u2212 F3 \u2212 \u221a 2 2 F2 \u2212 \u221a 2 2 F4 ) = bl ( \u21262 7 + \u221a 2 2 \u21262 6 + \u221a 2 2 \u21262 8 \u2212 \u21262 3 \u2212 \u221a 2 2 \u21262 2 \u2212 \u221a 2 2 \u21262 4 ) . The angular motion of any rotor is causing a drag moment that is opposite to the direction of the motion according to Newton\u2019s third law" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000718_978-1-4471-7278-9-Figure16.18-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000718_978-1-4471-7278-9-Figure16.18-1.png", "caption": "Fig. 16.18 Covering and routing solution for the Orbiter", "texts": [ " Using the Nearest Neighbor routing algorithm, the robot can work in one of two modes: sequentially covering and then routing when the operation is fully planned, or spontaneously covering and routing when the operation is improvised. In either case, the heuristics provide a satisfactory solution, if not optimal. The result shows that the total of 88 diamond-shape workspaces are needed for covering the whole area and the total operation time is 561 seconds. Figure 16.17 shows the Diamond covering solution and Fig. 16.18 shows the covering and routing solution. Navigation itself is a simple task given a big empty space. However, when the environment is complex, the navigation becomes a very complicated problem. With instinctual navigation algorithms such as wall-following and Hansel-and-Gretel, we found that simple algorithms can solve sophisticated problems such as solving maze puzzles and cover a service area with multiple constraints. Despite their simplicity, they can solve some hard problems such as the Traveling Workstation Problem that requires simultaneous covering and routing" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002163_is48319.2020.9199967-Figure19-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002163_is48319.2020.9199967-Figure19-1.png", "caption": "Figure 19. Experimental environments (bumpy walls)", "texts": [], "surrounding_texts": [ "We placed the proposed robot on an aerial ladder and a bumpy wall. In these experiments, the robot was manually controlled. Fig. 18 and 20 to 22 show the results. From these results, we confirmed that the complex behavior to complete these tasks is realized by pulling and loosening the strings by the two motors. We conclude that we can control many degrees of freedom of the soft body and produce intelligent behavior by the robot\u2019s dynamics No. 1 No. 2 No. 3 No. 4 Figure 20. Realized behavior (Bumpy walls Pattern 1). No. 1 No. 2 No. 3 No. 4 468 V. CONCLUSION In this paper, we focused on the physical properties of the soft arm and developed a ladder-climbing robot inspired by an octopus. We conducted experiments to demonstrate the effectiveness of the developed robot and confirmed that adaptive intelligent behavior could be realized by a twodimensional control input. Complex movements with many degrees of freedom were produced by the interaction between the soft arm and the environment. In our future work, we will apply machine learning to the soft robot to acquire the control input and demonstrate the effectiveness of the soft body with machine learning. ACKNOWLEDGEMENT This research was partially supported by the Japan Society for the Promotion of Science through the Grant-in-Aid for Scientific Research (18K11445). REFERENCES [1] K. Ito, S. Hagimori, \u201cFlexible manipulator inspired by octopus: development of softarms using sponge and experiment for grasping various objects\u201d, \u201c Artificial Life and Robotics, Vol. 22, issue 3 pp. 283\u2013 288, 2017 [2] K. Ito, R. Aoyagi, and Y. Homma, \u201cTAOYAKA-III: A Six-Legged Robot Capable of Climbing Various Columnar Objects,\u201d J. Robot. Mechatron., Vol.31, No.1, pp. 78-87, 2019. [3] T. Mukai, K. Ito,\u201dFlexible manipulator inspired by Octopi: Comparative study of pushing and pulling mechanisms in realizing intelligent behavior\u201d, Proc. of the Twenty-fourth International Symposium on Artificial Life and Robotics (AROB 24th), pp. 393\u2013396, 2019 [4] K. Ito, Y. Aso and K. Aihara, \u201cMulti-legged robot for rough terrain: SHINAYAKA-L VI\u201d, Proc. of the International Conference on Advanced Mechatronic systems (ICAMech 2019)\u201d, pp. 32, 2019 [5] K. Ito, H. Maruyama, \u201cSemi-autonomous serially connected multicrawler robot for search and rescue\u201d, \u201cAdvanced Robotics\u201d, Vol. 30, issue 7, pp. 489-503, 2016. [6] A. Saito, K. Nagayama, K. Ito, T. Oomichi, S. Ashizawa, and F. Matsuno, \u201cSemi-Autonomous Multi-Legged Robot with Suckers to Climb a Wall\u201d, \u201cJournal of Robotics and Mechatronics\u201d, Vol. 30, No.1, pp.24-32, 2018. [7] H. Yoneda, K. Sekiyama, Y. Hasegawa and T. Fukuda, \u201cVertical Ladder Climbing Motion with Posture Control Considering Gravitation Momentum for MultiLocomotion Robot\u201d, \u201cThe Japan Society of Mechanical Engineers\u201d, Vol. 75, No. 751, pp. 12-19, 2009. [8] M. Kanazawa, S. Nozawa, Y. Kakiuchi, Y. Kanemoto, M. Kuroda, K. Okada, M. Inada and T. Yoshiike, \u201cRobust Vertical Ladder Climbing and Transitioning between Ladder and Catwalk for Humanoid Robots\u201d, Proc. of the 2015 IEEE International Conference on Intelligent Robots and Systems (IROS), pp. 2202-2209, 2015. [9] T. Takemori, M. Tanaka and F. Matuno, \u201cLadder Climbing with a Snake Robot\u201d, Proc. of the 2018 IEEE International Conference on Intelligent Robots and Systems (IROS), pp. 8140-8145, 2018. [10] S. Fujii, K. Inoue, T. Takubo, Y. Mae and T. Arai, \u201cLadder Climbing Control for Limb Mechanism Robot \u201cASTERISK\u201d\u201d, Proc. of the 2008 IEEE International Conference on Robotics and Automation (ICRA), pp. 3052-3057, 2008. [11] D. Rus, M. Tolley, \u201cDesign, fabrication and control of soft robots\u201d, Nature, Vol.521, pp.467-475, 2015 [12] G. Sumbre, Y. Gutfreund, G. Fiorito, T. Flash and B. Hochner, \u201cControl of Octopus Arm Extension by a Peripheral Motor Program\u201d, Science, Vol. 293, No. 5536, pp. 1845-1848, 2001. [13] Y. Gutfreund, T. Flash, G. Fiorito and B. Hochner, \u201cPatterns of Arm Muscle Activation Involved in Octopus Reaching Movements\u201d, \u201cThe Journal of Neuroscience\u201d, Vol. 18, No. 15, pp. 5976-5987, 1998. Authorized licensed use limited to: Carleton University. Downloaded on November 04,2020 at 13:12:36 UTC from IEEE Xplore. Restrictions apply." ] }, { "image_filename": "designv11_5_0001397_1350650120942329-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001397_1350650120942329-Figure3-1.png", "caption": "Figure 3. Illustration of offset oil jet stream distance S and elevation angle for the equivalent helical gears.", "texts": [ " The x0, y0, z0 denotes the nozzle exit position; the elevation angle represents the angle between the jet flow and the z direction, which is always restricted to /24 4 ; and the azimuth angle \u2019 represents the angle between the jet stream projection line on the xO0y plane and the x direction, which is restricted to 4 \u20194 2 , and the point O0 is defined as gear coordinates origin. To identify the constraints for the oil jet orientations, the equivalent helical cylindrical gear of the spiral bevel gear is adopted as shown in Figure 3. To investigate spiral bevel gear meshing at the impingement point, the following method is proposed and used: first, the spherical involute profile36,37 near the impingement point is projected into the back cone; then two fan-shaped helical gears are attained by unfolding the conical surface, while the jet streamline is also projected into this extended surface. Obviously, the meshing at the impingement point for spiral bevel gears is similar to fan-shaped bevel gears, thus the latter can be approximately substituted for the former. As shown in Figure 3, the pitch radii and teeth numbers of the equivalent bevel gears are rIv \u00bc RI e tan p cos2 I , RI v \u00bc RI e tan g cos2 I NI pv \u00bc Np cos p cos3 I , NI gv \u00bc Ng cos g cos3 I 8< : \u00f01\u00de where Re is the outer cone distance; , are the reference cone angle and helical angle, respectively; the subscript p, g denote the pinion and gear, respectively; the subscript v represents the equivalent cylindrical gear; the superscript I, II denotes the impingement point. Here, sin I \u00bc 1 D0 \u00bdRI e \u00fe Rem RI e D0 sin m Rem\u00f0 \u00de where D0 is the milling cutter nominal diameter; the subscript m denotes middle point on meshing width" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002828_j.jii.2021.100265-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002828_j.jii.2021.100265-Figure3-1.png", "caption": "Fig. 3. (a) The clashed layers (in red) and other layers (in yellow), (b) the original 3D model, and (c) the processed 3D model of the part with layer number.", "texts": [ " Note that this robot motion planning method is used only on the current deposition layer. The collision between the robot/torch and other layers of the part is not considered at this stage. However, compared to the welding process, the volume of the part keeps growing dynamically in the WAAM deposition process, which may lead to more collision occurrences. Once robot motions for each layer have been planned, the collision between the robot and other layers of the same part can be tested. As shown in Fig. 3(a), layers coloured in red indicate collisions existed between the robot and/or torch while depositing a layer underneath. This information is stored and used to generate an initial collision matrix C, which is used as a heuristic to guide the later stages of the deposition sequence planning (Section 5). To assist with explanations over the following sections, a sample part, shown in Fig. 3(b), is provided. Firstly, the deposition path of each layer is generated from its 3D CAD model through a multi-direction WAAM planning algorithm (details can be found in [15]), as shown in Fig. 3(c). The individual layers are numbered categorically relating to their local sub-volume. Then, the collision matrix C is generated. This matrix is square, its dimensions equal to the number of layers of the part to be manufactured. As shown in Fig. 4, rows of the C represent weld paths of each layer; columns are used to represent a geometric collision model of each deposited layer. The Fig. 2. The overall workflow of the algorithm. L. Yuan et al. Journal of Industrial Information Integration xxx (xxxx) xxx welding simulation feature of the robotic AOLP engine is then used to fill the elements of C", " The data encoded in C is used in later stages to identify the collision between the robot/tool and the part. It provides a rapid look-up table for potential collision, which is useful as the part\u2019s geometric model changes as more layers are deposited. Further details of matrix C are explained as below, (i) The first column of C is the deposition sequence determined according to the sub-volume numbers. (ii) Column naming conventions in C are listed, for example, as a,b. In our system, the 3D geometric model of the part is divided into subvolumes, as shown in Fig. 3(c). \u201ca\u201d denotes the sub-volume number and \u201cb\u201d represents the specific layer number within its sub-volume. (iii) In matrix C, a \u20180\u2019 element denotes the corresponding deposition is collision-free, and \u201c \u00d7 \u201d means that some form of collision has been detected for that particular weld path. (iv) The elements in the bottom-left triangle region of matrix C represent the collision information obtained during the deposition process. For a particular layer, all \u201c \u00d7 \u201d in the bottom left triangle means collision occurs between the specific layer and the layers deposited before it" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000286_b978-0-08-102884-1.00001-7-Figure1.15-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000286_b978-0-08-102884-1.00001-7-Figure1.15-1.png", "caption": "FIG. 1.15 The all-electric-aircraft concept. The possible location of electrically powered actuators, drives and related functions within a civil aircraft are shown. The location of power generation is highlighted in bold italics, the APU (auxiliary power unit) is normally only run on the ground. The actuators for the flying surfaces can either be powered by hydraulics or electrical power. The hydraulic pumps can be either directly coupled to the engines or driven by an electric motor.", "texts": [ " As with automotive applications, the aerospace industry is faced with challenge of reducing operating costs, particularly fuel, and reducing environmental impact. In addressing this challenge, designers are increasingly turning to the concept of the moreelectric-aircraft and the all-electric-aircraft. In this approach systems are replaced with an electrically powered equivalent. While this will increase the aircraft\u2019s electrical power requirements, there is an overall saving in weight and increase in efficiency (Jones, 2002). Fig. 1.15 shows the key components and load centres for the all-electric-aircraft. Two main areas are being addressed, cabin air supply and actuators. In the majority of passenger aircraft, the cabin air supply is bled from the engines, this has a detrimental impact on engine performance as in most conditions, conventional pneumatic systems withdraw more power than needed, causing excess energy to be dumped. In bleed-less technology, no HP air is extracted from the engines, allowing more efficient thrust production and engine operations" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001772_j.compstruct.2020.111986-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001772_j.compstruct.2020.111986-Figure1-1.png", "caption": "Fig. 1. Structural diagram of tow-placement head (TPH) attached on an industrial robot.", "texts": [ " With the curvature correction method, the gap-error rate of the variable stiffness laminate can be well controlled and the uniformity of the thickness of the laminate is well ensured. Finally, a practical case is investigated to demonstrate the effectiveness of the proposed trajectory planning method. 2. The requirements and constrains of fiber tow placement AFP technologies have been widely used in the aerospace structural components, such as aircraft wings, fuselage, control surfaces, etc. In our lab, the AFP procedure is completed by a 6-DOF industrial robot equipped with a tow-placement head (TPH) on the end-effector, as shown in Fig. 1. The TPH is equipped with a heat-pressure device for fixing and curing a fiber tow on to a mold. The heat-pressure device including a heating member, a pressure member, a fixing frame and a roller. The roller is surfaced with silicone rubber with excellent wear resistance and heat resistance. The individual tows, are fed through a tension system to the tow-placement head. The path width is determined by the number of individual tows, which is two in our machine. In practical applications, mainly three widths of tows: 3" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001800_978-3-030-40410-9-Figure3.8-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001800_978-3-030-40410-9-Figure3.8-1.png", "caption": "Fig. 3.8 Directions of the roll, pitch, and yaw angles", "texts": [ " Other successive rotations about the principal axes of frame o1x1y1z1 called roll, pitch, and yaw angles\u2014denoted as (\u03c6, \u03b8, \u03c8), respectively\u2014can be used to specify the rotation matrix. We will describe the three successive rotations as follows. The first rotation is roll rotation about the z1-axis through \u03c6 angle, then pitch rotation about the y1-axis by an angle \u03b8 , and yaw about the x1-axis by an angle \u03c8 , as revealed in Fig. 3.7. Since the rotations are made relative to frame o1x1y1z1, as illustrated in Fig. 3.8, the transformation matrix will be RXY Z = \u23a1 \u23a3C\u03c6 \u2212S\u03c6 0 S\u03c6 C\u03c6 0 0 0 1 \u23a4 \u23a6 \u23a1 \u23a3 C\u03b8 0 S\u03b8 0 1 0 \u2212S\u03b8 0 C\u03b8 \u23a4 \u23a6 \u23a1 \u23a3 1 0 0 0 C\u03c8 \u2212S\u03c8 0 S\u03c8 C\u03c8 \u23a4 \u23a6 = \u23a1 \u23a3C\u03c6C\u03b8 \u2212S\u03c6C\u03c8 + C\u03c6S\u03b8 S\u03c8 S\u03c6S\u03c8 + C\u03c6S\u03b8C\u03c8 S\u03c6C\u03b8 C\u03c6C\u03c8 + S\u03c6S\u03b8 S\u03c8 \u2212C\u03c6S\u03c8 + S\u03c6S\u03b8C\u03c8 S\u03b8 C\u03b8 S\u03c8 C\u03b8C\u03c8 \u23a4 \u23a6. (3.25) Example 3.2 In general, we can write a program using MATLAB to calculate the Euler transformation matrix from the basic three angles. We have considered that the rotation will be in degrees, so we have used the function cosd( ) and sind( ). If the rotations are in radian, then it is proper to use the functions cos( ) and sin( )" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001443_rpj-11-2019-0287-Figure8-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001443_rpj-11-2019-0287-Figure8-1.png", "caption": "Figure 8 Microstructural observation of the samples", "texts": [ " Prior to the observation of microstructure, all as-built samples were polished using a Wirtz Buehler (Dusseldorf, Germany) polishing machine and then chemically etched. The etchant was the modified Fry\u2019s reagent that was prepared by a mixture of 0.2 g CuCl2, 10mL HCl, 30mL H2O and 10mL HNO3 at the room temperature and was uniformly applied to the surface of the samples, retained for 60 s, then rinsed with alcohol and blown dry. The surface observed for all samples was the supporting surface, as shown in Figure 8. The microstructure of the samples was observed by a LEICADM2500M (Wetzlar, Germany) optical microscope, Keyence VHX-100 (Osaka, Japan) digital microscope and JEOL JSM-5500 (Tokyo, Japan) scanning electronmicroscope (SEM). The dimensional error of the as-built samples with and without support structures at different building angles is depicted in Figure 9. The horizontal axis represents different building angles, and the vertical axis represents the dimensional errors. It can be seen that when the building angles are 45 and 60\u00b0, with support structures, the dimensional error of the samples is small, indicating that the dimensional accuracy of the samples is high" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001839_012015-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001839_012015-Figure7-1.png", "caption": "Figure 7. Tribotester", "texts": [ " The printed specimens were examined by three tests which are tribology, surface structure (optical microscope) and surface roughness. The tribological measurement was performed at tribology laboratory in Szent Istv\u00e1n University under the laboratory conditions. The first step in the measurement process is to connect the measuring circuit, which consists of a computer, Spider 8 measuring converter, a tribotester, and an inverter. Spider 8 is a strain gauge measurement device for measuring load, force, and wear. The indispensable part of the measuring system is the tribotester, which is in this case compatible with polymers. Figure 7 displays the structure of the used PLINT TE 77 tribotester (High-Frequency Tribotest), employed by Zsidai and Kal\u00e1cska in previous research [18]. Before starting the measurement, some parameters must be defined which are given in Table 3. A computerised microscope (ZEISS brand) with a high quality camera and four magnification lenses (10x, 20x, 50x and 100x) was used to study the surface structure of the samples before and after the tribology test (displayed in figure 8). Surface roughness of steel counterpart, Ra [\u03bcm] 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0003099_312-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0003099_312-Figure1-1.png", "caption": "Fig. 1. Principle of two-ball machine. Motion at point of contact is combined rolling and sliding as in gear tooth engagement. Severe loads cause surface breakdown.", "texts": [ " This temperature for failure will then be the measure of the lubricating ability of the oil. This contribntion deals with the examination of an expression based on this concept relating the operating variables of a test machine. A P P A R A T U S The apparatus employed was the two-ball machine, the idea of which originated with Dr. H. E. Merritt. The machine was supplied as a prototype by David Brown & Sons (Huddersfield) Ltd., but has subsequently been considerably developed for research purposes at the Thornton Research Centre. The principle is illustrated in Fig. 1. The two steel balls of 1 i i n diameter are mounted at the respective ends of two vertical spindles, SUPPLEMENT No. 1 35 the axes of which are $in apart. Both spindles are capable of rotation, in the same or opposite direction as required and also at different speeds. Conditions of combined sliding and rolling in the contact are thus produced. Fig. 2 shows the apparatus. It consists essentially of two parts, coupled by a horizontal hinge so that the upper part can be tilted back clear of the lower" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001612_j.oceaneng.2020.108224-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001612_j.oceaneng.2020.108224-Figure7-1.png", "caption": "Fig. 7. The plan graph of the towing tank.", "texts": [ " 6(c) is the position of the sensor installed. The heave displacement data are measured using an ultrasonic sensor, and the pitch angle is measured using the posture sensor. The acceleration data of the heave and pitch are measured using accelerometers, which are installed in the bow, stern, and center of the trimaran. The accelerometers are installed inside of the trimaran. Fig. 6(d) and (e) show the composition of the actuator. The towing tank has a wave damper region, and the plan graph of the towing tank is shown in Fig. 7. For an irregular wave, the conditions of the trimaran\u2019s model tests are listed in Table 4. Hs represents the significant wave height. The trimaran\u2019s velocities were 2.93 m/s and 6.51 m/s, which are equivalent to 18 knots and 40 knots, respectively, because the trimaran is a 1:10 model. The displacement and acceleration data of the heave and pitch were measured using an ultrasonic sensor, an accelerometer, and a posture sensor. The acceleration sensor was installed at the center, bow, and stern of the vessel" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001423_s42835-020-00443-4-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001423_s42835-020-00443-4-Figure3-1.png", "caption": "Fig. 3 The framework of LOS navigation strategy", "texts": [ " Continuous function N( ):\u211c \u2192 \u211c is called a Nussbaum function if it satisfies the following conditions [39]: (12)u = sat(vp) = \u23a7 \u23aa\u23a8\u23aa\u23a9 sgn(vp)uM , \ufffd\ufffd\ufffdvp \ufffd\ufffd\ufffd \u2265 uM vp, \ufffd\ufffd\ufffdvp \ufffd\ufffd\ufffd < uM (13)g(vp) = uM \u00d7 tanh(vp / u) = uM evp\u2215uM \u2212 e\u2212vp\u2215uM evp\u2215uM + e\u2212vp\u2215uM (14) \u23a7\u23aa\u23aa\u23a8\u23aa\u23aa\u23a9 lim k\u2192\u221e sup 1 k \u222b k 0 N( )d = \u221e lim k\u2192\u221e inf 1 k \u222b k 0 N( )d = \u2212\u221e 1 3 Lemma 1 VN(t) and (t) are smooth functions on [0, tf ] , where VN(t) \u2265 0 and \u2200t \u2208 [0, tf ] . For a Nussbaum function N( ) , the following inequality is valid. where 0 and C are nonzero constants. The certification process of Lemma 1 can refer to paper [40]. The function of LOS navigation strategy is to change the reference path into a controllable nominal course (target course). The framework of LOS navigation strategy is shown in Fig.\u00a03. where is the course of USV, (x, y) is the USV\u2019s actual position, (xp( ), yp( )) is the target position, is an independent variable. = a tan 2(v, u) is the sideslip angle, where u is surge velocity and v is sway velocity. U = \u221a u2 + v2 represents the speed of USV. xe and ye stand for the along-track error and the cross-tracking error respectively. p = a tan 2(x\ufffd p ( ), y\ufffd p ( )) is a rotation angle, where y\ufffd p ( ) = y and x\ufffd p ( ) = x . The motion model of USV can be expressed as Define the following error variables" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002525_taes.2021.3061795-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002525_taes.2021.3061795-Figure7-1.png", "caption": "Fig. 7. Outdoor autonomous flight hardware: (a) the custom built quadrotor at the Flight Dynamics and Control Lab at GWU (b) components of the base module", "texts": [ " Restrictions apply. 0018-9251 (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. represented by rotation matrices. It shows the convergence of the attitude within few milliseconds. Next, outdoor autonomous flight tests were performed utilizing the proposed estimator. 2 Another quadrotor UAV built at the Flight Dynamics and Control Lab at GWU (see Figure 7- (a)) was used as the testing platform. The rover module with the Jetson TX2 computing module was attached to the UAV. In addition to the IMU and the RTK GPS sensors, the Jetson TX2 at the rover module is connected to MikroKopter BLCtrl v2 ESC units over I2C, to control four 700 kV T-Motor brushless DC motors with 10\u00d7 4.7 propellers. The states estimated by the proposed method were used to control the UAV using the geometric controller proposed in [21]. The measurement error covariance estimated by the sensors are used as the sensor noise values in the estimator" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000814_b978-0-12-814401-5.00008-6-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000814_b978-0-12-814401-5.00008-6-Figure6-1.png", "caption": "Fig. 6 Schematic drawing of a SAW delay line. (Reproduced from Di Pietrantonio, F., et al., 2016. Sens. Actuators B Chem. 226, 1\u20136.)", "texts": [ " In SAWdevices, the acoustic wave propagates on the surface of the piezoelectric film and is generated by interdigital transducers (IDTs) fabricated on the same surface. Two main configurations are available for SAW devices, depending on the pattern of the IDTs: resonator or delay line. In SAW resonators, the input and output IDTs are fabricated between two reflectors in order to create a resonant cavity for the surface wave (Fig. 5). The reflectors are not used in SAW delay lines where the input and output IDTs are positioned at a characteristic distance (Fig. 6). In this configuration, the input IDT is used to generate the wave, and the output IDT is employed for the detection of the propagating wave. In particular, the distance between the fingers and the lateral dimension of the single finger are proportional to the wavelength and are used to fix the working frequency of the device. The typical dimension of a SAW device is in the order of some square millimeters and is typically fabricated, not only on piezoelectric thin films but also especially on crystalline substrate such as quartz, lithium niobate (LiNbO3), lithium tantalate (LiTaO3), and langasite with different cut and orientation" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002269_lra.2020.3039732-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002269_lra.2020.3039732-Figure4-1.png", "caption": "Fig. 4. Surgical FOV control with RCM constraints.", "texts": [ " NDI FocusT = NDI End T \u00b7 \u23a1 \u23a2\u23a2\u23a2\u23a3 1 0 0 0 0 1 0 0 0 0 1 R 0 0 0 1 \u23a4 \u23a5\u23a5\u23a5\u23a6 (1) We keep the orientation of End FocusT at the focal point unchanged and move its origin position OFocus to the instrument tip OIns (NDI Ins T (1 : 3, 4) is the origin point of NDI Ins T ) NDI NewFocusT = [ NDI FocusR NDI Ins T (1 : 3, 4) 0 1 ] (2) Then, we keep the orientation of the coordinate system unchanged and move the origin of the obtained coordinate system to point O1 End; then we obtain: NDI O1 End T = NDI NewFocusT \u00b7 \u23a1 \u23a2\u23a2\u23a2\u23a3 1 0 0 0 0 1 0 0 0 0 1 \u2212R 0 0 0 1 \u23a4 \u23a5\u23a5\u23a5\u23a6 (3) At this time, the Z-axis vector of NDI O1 End T and the Z-axis vector of NDI Ins T can be extracted as follows: NDIzO1 End = NDI O1 End T (1 : 3, 4) (4) NDIzIns = NDI Ins T (1 : 3, 4) (5) Based on the two extracted Z-axis vectors, the current spatial angle \u03b81 between the endoscope and the instrument and the rotation axis NDIzaxis(kx, ky, kz) from the endoscope to the instrument can be calculated as follows: \u03b81 = acos \u239b \u239d NDIzIns \u00b7 NDIzO1 End \u2016NDIzIns\u2016 \u00b7 \u2225\u2225\u2225NDIzO1 End \u2225\u2225\u2225 \u239e \u23a0 (6) NDIzaxis = NDIzIns \u00d7 NDIzO1 End \u2016NDIzIns\u2016 \u00b7 \u2225\u2225\u2225NDIzO1 End \u2225\u2225\u2225 (7) According to the space required between the surgeon\u2019s hand and the end-effector of the robot, we can set the desired tracking angle \u03b82, and then the desired endoscope pose can be obtained as: NDI O2 End T = RK(\u03b8) \u00b7 NDI O1 End T (8) RK(\u03b8) = \u23a1 \u23a2\u23a2\u23a2\u23a3 kxkxv\u03b8 + c\u03b8 kxkyv\u03b8 \u2212 kzs\u03b8 kxkzv\u03b8 + kys\u03b8 0 kxkyv\u03b8 + kzs\u03b8 kykyv\u03b8 + c\u03b8 kykzv\u03b8 \u2212 kxs\u03b8 0 kxkzv\u03b8 \u2212 kys\u03b8 kykzv\u03b8 + kxs\u03b8 kzkzv\u03b8 + c\u03b8 0 0 0 0 1 \u23a4 \u23a5\u23a5\u23a5\u23a6 (9) where \u03b8 = \u03b82 \u2212 \u03b81,c\u03b8 = cos\u03b8,s\u03b8 = sin\u03b8,v\u03b8 = 1\u2212 cos\u03b8, and RK(\u03b8) is the equivalent axis angle rotation matrix [38]. When the current and desired endoscope poses are known, the robot drives the endoscope to move from the current pose to the desired pose, which can realize surgical FOV control without RCM constraints (for the tracking motion outside the nasal cavity, the sixth DOF of the robot [26] does not participate in motion). Authorized licensed use limited to: Auckland University of Technology. Downloaded on December 20,2020 at 02:17:57 UTC from IEEE Xplore. Restrictions apply. Fig. 4 shows the schematic diagram of surgical FOV control based on tracking motion with RCM constraints. For surgical FOV control without RCM constraints, the desired position of the endoscope tip is determined by the position of the instrument tip and the focal length of the endoscope, and the orientation of the endoscope is determined by the preset tracking angle. For surgical FOV control with RCM constraints, the desired position of the endoscope tip is still determined by the position of the instrument tip and the focal length of the endoscope" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000100_ro-man46459.2019.8956305-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000100_ro-man46459.2019.8956305-Figure1-1.png", "caption": "Fig. 1. ROV Minerva with kinematics notations. It has one lateral thruster, two longitudinal thrusters, and two vertical thrusters. Photo: Johanna Jarnegren.", "texts": [ " Detection of actuator faults and faults in navigation sensors are considered as a combined problem in this brief, and experimental data are used to assess the occurrence of different fault types. Actuator reconfiguration possibilities are rather limited on the vessel considered and are not within the scope of this brief. This brief is organized as follows. Section II describes the model of the ROV used in the experiment and its failure modes. Section III, presents the switched HMM, the PF and the navigation system design. Results from ROV sea trials are presented in Section IV. The ROV Minerva [18], as shown in Fig. 1, is a SUB-fighter 7500 ROV. It is powered from and communicates with a surface supply vessel through a 600-m umbilical cable. Minerva is equipped with five thrusters and various navigation sensors. A hydroacoustic positioning reference (HPR) system, is used to measure the position of the ROV relative to a transducer on the surface vessel. A Doppler velocity log (DVL) is installed to measure the ROV velocity. An inertial measurement unit (IMU) provides turn rate and heading measurements. Depth is provided by the HPR and also by a pressure gauge" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002356_978-3-030-50460-1-Figure20.1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002356_978-3-030-50460-1-Figure20.1-1.png", "caption": "Fig. 20.1 Geometry of the tooth", "texts": [ " In the present paper, gear-mesh stiffness for healthy and spalled gear have been calculated using the potential energy method. This gear-mesh stiffness values were utilized to obtain the dynamic response of the system using a 6-DoF dynamic model. Later time domain statistical features were calculated for the early indication of spall fault. In this paper, gear-mesh stiffness is evaluated based on the potential energy method proposed by Yang and Lin (1987) and further refined by Tian et al (2004) for the calculation of the gear-mesh stiffness. Figure 20.1 shows the geometry of the single tooth used for the calculation of stiffness. The total mesh stiffness can be obtained by the equation (20.1) for single pair of teeth in contact and equation (20.2) for the two pairs of teeth in contact. Kt = 1 1 Khp + 1 Kap + 1 Kbp + 1 Ksp + 1 Kag + 1 Kbg + 1 Ksg (20.1) 366 Handikherkar, Phalle Kt = 2\u2211 i=1 1 1 Khpi + 1 Kapi + 1 Kbpi + 1 Kspi + 1 Kagi + 1 Kbgi + 1 Ksgi (20.2) The expressions of Hertzian stiffness, axial compressive stiffness, shear stiffness and bending stiffness components are shown in equations (20" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002166_s10846-020-01259-0-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002166_s10846-020-01259-0-Figure6-1.png", "caption": "Fig. 6 NED coordinate system", "texts": [ " In addition, the yaw angle measured by this method is obtained by reference to the geomagnetic vector, while the geomagnetic north direction and the geographic north direction are not coincident. In this paper, if not specified, the north direction of the geographic coordinate system is along the north direction of the geomagnetism, and the yaw angle is measured relative to the magnetic north. The geographic coordinate system used in this paper is the NED (North-East-Down) coordinate system, as shown in Fig. 6, and the X-axis of the coordinate system is in the horizontal plane pointing in the north direction, the Z-axis is pointing vertically downwards to the earth, while the Y-axis is determined according to the right-hand rule. The acceleration of gravity in the geographic coordinate system is expressed as Gn = [ 0 0 g ]T . The gravity acceleration in the sensor coordinate system obtained from accelerometer measurements can be expressed as Gb =[ ax ay az ]T , and Gb = Cb nG n, Cb n is the transformation matrix as Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002236_j.isatra.2020.10.053-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002236_j.isatra.2020.10.053-Figure2-1.png", "caption": "Figure 2: Mechanical assembly with the detailed design of single link.", "texts": [ " It is placed over the cover. Support of the spring bar: This structure is mounted over the upper cover. It maintains the bar in the predefined place, but it does not interfere with the rotational movement. Joint: This part connects two spring bars that are rotated 90\u25e6 with respect to each other. This configuration provides a reduction of the link size. Union: It connects the main structure in the link with the joint. It provides space for the instrumentation wire. he sections of the link are shown in Figure 2. The A actuators are placed internally and the spring bar laced into the joint to guarantee the union. ll the parts forming each link in the RS were nufactured by using 3D printing technology. The nufacturing of these elements was completed in a kerbot R\u00a9 printer. The printing process was carried t in 1.75mm of diameter ABS material with a diamr of 0.4mm of the nozzle. igure 2 shows the holder of the link in the main ucture. That section contains the actuators. The metry of this element is important considering the t that the position of the actuators must be equidist each other to standardize the SMA spring design to reduce disturbances that may affect the moveme at each joint" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002166_s10846-020-01259-0-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002166_s10846-020-01259-0-Figure2-1.png", "caption": "Fig. 2 Schematic diagram of the reference block and the measurement block", "texts": [ " One is the V-frame with GB4972-85 standard in China used as the reference block, and the main working surface of the reference block is its 90\u25e6 V-shaped groove surface. The parallelism of the V-groove to the bottom surface, the symmetry of the side faces, and the perpendicularity of the end faces are all guaranteed with accuracy. The other tool is a cuboid with inner hollow used as the measurement block. The cuboid is machined by using wire-electrode cutting to ensure the perpendicularity between the surfaces. The schematic diagram of the reference block and the measurement block is shown in Fig. 2. Besides that, the input applied to the accelerometer of each measurement attitude is gravity acceleration. Since the accelerometer measures the specific force, the vertical downward gravity acceleration is equivalent to acceleration input in the opposite direction. Therefore, the equivalent input generated by the gravitational acceleration expressed in the reference coordinate system (Geographic Coordinate System) is [ 0 0 g ]T (g \u2248 \u22129.8m/s2). And the input vector in the input matrix is represented under the sensor coordinate system, so the input has to perform the corresponding coordinate transformation", " The calibration of the gyroscope can be determined by applying a rotation vector and then measuring the output. However, it is difficult to accurately load a rotation vector onto a certain axis of the gyroscope in practice due to the error in the process of mounting the chip on the PCB and mounting the PCB on the measuring block. Therefore, the gyroscope is also calibrated by multiple loading methods. The tools used to calibrate the gyroscope is the same as that used to calibrate the accelerometer (Fig.2). During the experiment, the reference block is placed on the rotating platform with its rotation vector perpendicular to the plane of the turntable, as in Fig. 5. Since the Z axis of the reference block is perpendicular to the bottom surface, the rotation vector coincides with the Z axis of the reference block. That is, the coordinates of the input rotation vector in the reference block coordinate system are expressed as[ 0 0 \u03c9L ]T , where, \u03c9L is the value of the input rotation vector. The measuring block can constrain all its three rotational DOFs (Degrees of freedom) through its fitting with Vgroove" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000427_0954410016643978-Figure15-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000427_0954410016643978-Figure15-1.png", "caption": "Figure 15. Simulative casing bolted joints structure (mm).", "texts": [ " The axial elastic modulus E021 of the yellow part is defined based on the fractal contact theory,37 the details are shown in Appendix 2. Subsequently, the axial elastic modulus E022 of the green part is obtained by equations (36), (41), (42), and the shear modulus of the partitioned thin-layer elements is G021 \u00bc E021 2 \u00f01\u00fe \u00de , G022 \u00bc E022 2 \u00f01\u00fe \u00de \u00f043\u00de at RMIT University Library on May 19, 2016pig.sagepub.comDownloaded from Mode experiment of simulative casing bolted joints structure As shown in Figure 15, it is the dimensions of the simulative casing bolted joints structure, including two thin-walled cylinders connected by 12 M20 bolts. The corresponding material parameters of the casings are given in Table 1. Each bolt is applied by the same preload of 100, 80, 60, 40, and 20N.m, respectively. The natural frequencies and the mode shapes of the structure are measured by the hammering method, and there are 240 points for measuring the mode shapes. One acceleration sensor is placed in the structure" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001644_j.conbuildmat.2020.121961-Figure8-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001644_j.conbuildmat.2020.121961-Figure8-1.png", "caption": "Fig. 8. Plastic development process of locked coil wire rope.", "texts": [], "surrounding_texts": [ "Fig. 7 shows the stress development process of the locked coil wire rope with boundary \u2161. To facilitate the observation of the phenomenon, this study intercepts half of the model for display, and hides the loading beam part. It can be observed from the figure that when the axial strain reaches 0.002, the stress distribution of the section becomes uniform. Owing to the spatial arrangement of the steel wires, the axial stiffness of the steel wires of each layer decreases with an increase in the lay angle, and thus the stress of the central steel wire is large. With the increase in axial load, owing to the mutual extrusion of steel wires, the maximum stress appears at the contact part of the steel wires. The central steel wire yielded at first, and the axial stiffness of the cable decreased. After the yield of the centre wire, stress redistribution begins, and the stress of the Z-shaped steel wire increases. When the axial strain reached 0.017, the Z-shaped steel wire failed and the axial bearing capacity decreased significantly. When the axial strain reaches 0.02, the cable fails, which is consistent with the development law in the load\u2013strain curve. Fig. 6. Stress cloud diag" ] }, { "image_filename": "designv11_5_0000888_j.promfg.2019.06.223-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000888_j.promfg.2019.06.223-Figure4-1.png", "caption": "Fig. 4. Modeling of position-based spatial constraint.", "texts": [ " According to the posture and position of a CF\u2019s corresponding actuator, one or more subregions are set to be the work region(s). To avoid collisions, any CFs located in these regions are not deposited concurrently with the central CF. In the example above, R7 is the work region for the blue CF because its corresponding robotic arm occupies part of this sub-region during deposition, as shown in Fig. 3(b). Since the red CF to be deposited by another actuator lies in R7, these two CFs cannot be deposited concurrently. In another example shown in Fig. 4(c), the work regions for the blue CF could be R5, R6 and R7 when it is deposited by a SCARA. Based on this model, for two CFs, CF1 and CF2, they are firstly given safety envelops and work regions. Subsequently, interference of work regions, i.e., whether CF1 lies in any work region of CF2 or vice versa, will be checked. These two CFs can be deposited concurrently if there is no interference. Position-based spatial constraint exists when the actuators have to follow a position order. A typical example is the composite X-Y stage with multiple actuators", " 5(a), each end-effector (actuator) can move independently, but they must follow a position order in the X-axis and cannot get across one another. In a typical coordinate system, layers of fabrication materials are deposited in the X-Y plane and stacked along the Z-axis. To model position-based spatial constraint, each actuator is given a position index to indicate its order in the Xand Y-axis. A larger X index value indicates the actuator is on the right side of those with smaller values in the X-axis, and a larger Y index indicates it is above those with smaller values in the Y-axis. For the composite X-Y stage in Fig. 4(a), the three end-effectors can be given position indices X(-1), X(0), and X(1) to indicate their position order in the X-axis respectively. Since they do not have a restricted position order in the Y-axis, their indices are all set to be Y(0), as in Fig. 4(b). Based on this model, for two CFs, CF1 and CF2, the proposed approach will check whether the positions of CF1 and CF2 match the position indices of their actuators. Specifically, Yi Cai et al. / Procedia Manufacturing 34 (2019) 584\u2013593 587 4 Author name / Procedia Manufacturing 00 (2019) 000\u2013000 it examines whether the CF whose actuator has a larger X index is located on the right side in X-axis, and that with a larger Y index is on the upper side in Y-axis. For example, if the X index of CF1\u2019s corresponding actuator is larger than that of CF2\u2019s, it means CF1\u2019s actuator should be on the right of CF2\u2019s" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001980_1099636220931479-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001980_1099636220931479-Figure2-1.png", "caption": "Figure 2. The 2/1 stacking lay-up of FMLs.", "texts": [ " This was followed by the cooling process on the annealed aluminium until the ambient temperature was reached. The physical and tensile properties of annealed aluminium 5052 are summarised in Table 1. The preparation of FMLs was then initiated by the mechanical surface treatment of the aluminium using 80-grit size silicon carbide abrasive paper to improve the surface roughness. The increase of surface roughness of aluminium had been identified that can improve the adhesive bonding between aluminium and composite [29]. FMLs with 2/1 lay-up as elucidated in Figure 2 were fabricated by stacking the aluminium layers to the composite core with the placement of 0.06mm thick PP adhesive films (Collano adhesives) at the metal-composite interfaces. FMLs were then hot compressed at a temperature of 170 C and pressure of 1MPa for 8minutes. Finally, FMLs were subjected to rapid cooling until the ambient temperature and taken out from the hot press machine for visual inspection of any defects. The sample preparation scheme for FMLs is depicted in Figure 3. FMLs are defined based on the metal volume fraction (MVF), which can be determined by using equation (1)" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001644_j.conbuildmat.2020.121961-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001644_j.conbuildmat.2020.121961-Figure3-1.png", "caption": "Fig. 3. Refined finite element model of locked coil wire rope.", "texts": [ " The figure shows the model of a Zshaped steel wire, which can restore the true shape of the steel wire to the greatest extent. The steel wire model was imported into the ABAQUS/Explicit finite element software, and then assembled Table 1 Geometrical data of the 22 mm locked coil wire rope. Strand diameter (mm) Wire diameter (mm) Strength (MPa) Shape Number Lay length pi (mm) Lay angle, ai ( ) 22 O1 3 1670 round 1 200 \u2013 O2 2.95 1670 round 6 200.9 5.4 O3 3 1670 round 12 203.5 10.8 Z1 3.5 1570 Z 22 208.3 16.65 to obtain the model of the locked coil wire rope, as shown in Fig. 3. The von Mises yield criterion was assumed. The material model used was MAT_PLASTIC_KINEMATIC with von Mises yield criterion and bilinear isotropic strain hardening available in ABAQUS, and the Poisson\u2019s ratio is 0.3. To facilitate the calculation in engineering, the boundary conditions of cable members are usually simplified as hinged boundaries. Previous studies have shown that, in fact, the boundary of a cable member lies between the rigid and hinged joints. To study the influence of end conditions on the mechanical properties of the cable, four types of boundaries, that is, rigid joint boundary, loading beam boundary, simulated cable node boundary, and hinged boundary, are established in this study, which are named as boundary I, boundary II, boundary III, and boundary IV, as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001758_s00202-019-00915-5-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001758_s00202-019-00915-5-Figure1-1.png", "caption": "Fig. 1 General schematic diagram\u2014SRM drive system", "texts": [ ", adaptive neural fuzzy interference system (ANFIS) and fuzzy-PID are explained in Sect.\u00a03. In Sects.\u00a04 and 5, simulated results and experimental verification are discussed and the results are presented. Finally, Sect.\u00a06 concludes the findings in proposed hybrid observer performance using FLC and ANFIS control technique. The general blocks consist of SRM drive system such as DC voltage source, capacitive filter, power converter circuit, SRM motor, rotor position sensor, speed controller, gate pulses triggering unit, and its general schematic diagram is shown in Fig.\u00a01. In the conventional methods, rotor position is determined by using position sensor. With the advent of digital signal processing, it can control without any mechanical sensors such as speed and position sensors. This type of control mechanism is generally called as \u2018sensorless speed control mechanism\u2019. The sensorless speed control mechanism through hybrid observer algorithm (HOA) is explained in the following sections. The nonlinear property of the SRM drive system can demonstrate through the differential equations" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001357_012022-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001357_012022-Figure1-1.png", "caption": "Figure 1. Double Aided Nozzle [22] Figure 2. Four legged Column Frame [22]", "texts": [ "In this project, A mechanical column-type structure consisting of a four-column frame shown in figure 2 for 3D printing was set up. The 3D printer was divided into several structural parts in the frame: X-axis path, Y-axis path, Z-axis column path, print head, top stabilization system and ready-mixed concrete. This printer had a width of the X-axis and height of the Z-axis as 20m. The Y-axis length could be extended indefinitely by increasing the column number of Z axis. The Y-axis length could be extended indefinitely by increasing the column number of Z axis. A double aided nozzle was used in this project as shown in figure 1 that can extrude concrete with 15 mm diameter aggregates. The upper end of the concrete feeding system was constructed with a V- ICSICME- 2020 IOP Conf. Series: Materials Science and Engineering 814 (2020) 012022 IOP Publishing doi:10.1088/1757-899X/814/1/012022 shaped vibrating sieve. This sieve had a 15 mm pore spacing and can filter out aggregates with a diameter of more than 15 mm. The dual-assisted print head's working principle was that feed containers A and B can work together at the same time" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002018_0142331220932649-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002018_0142331220932649-Figure2-1.png", "caption": "Figure 2. The structure of the considered parallel robot.", "texts": [ " (b) In rehabilitation, surgery and some other applications, the desired impedance dynamics are designed in low-frequency range in which the acceleration in the desired impedance dynamics varies slowly. Moreover, the stiffness and damping signals are much larger than the stiffness signal in the desired impedance dynamics. Thus, the error e1 and its time derivative _e1 play a dominant role in the low-pass filtered form of e1 in (26). To illustrate the effectiveness of the proposed variable impedance control laws, simulations are taken on a five-bar parallel robot with the structure described by Figure 2, where links 1 and 3 are parallel, links 2 and 4 are parallel and have same length, and the point P is the end effector point. The matrices M(q), C(q, _q) in equation (1) are presented as M(q)= d11(q) d12(q) d21(q) d22(q) ,C(q, _q)= 0 h _q2 h _q1 0 , F( _q)=diag(d1, d2) _q \u00f032\u00de where d11(q)=m1l2 c1 +m3l2 c3 +m4l2 1 + I1 + I3, d12(q)= d21(q)= (m3l2lc3 +m4l1lc4) cos (q2 q1), d22(q)=m2l2 c2 +m3l2 2 +m4l2 c4 + I2 + I4, h= (m3l2lc3 +m4l1lc4) sin (q2 q1) \u00f033\u00de with l1, l2, l3 and l4 being lengths of the joints links, lc1, lc2, lc3 and lc4 being distances between the joints and respective centre of masses, I1, I2, I3 and I4 being inertial moments of the respective links, and m1,m2,m3 and m4 being masses of the respective links" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002356_978-3-030-50460-1-Figure17.8-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002356_978-3-030-50460-1-Figure17.8-1.png", "caption": "Fig. 17.8 CAD models for 3D printing of designs a) and b) with equal mass, thickness and clamping area", "texts": [ " For experimental validation, the two configurations of Fig. 17.7 are designed for 3D printing with equal mass and tested in tension. An AlMgSc aluminum alloy is used with approximated mechanical values described in the last section. At the current stage, the analysis does not consider that the mechanical properties for the 3D printed material can be inhomogeneous with possibly included material defects, and can depend on the cross sectional dimension. Each specimen is fixed on the left side of Fig. 17.8 and loaded axially with a uniform traverse speed of 1.0 mm/s on the other side. The applied force raised by the machine for impinging the uniform speed on the specimen is measured in the experiment. In Fig. 17.9 and the graphs of Fig. 17.10, the tensile test results are illustrated. The 17 Multimodal approach for automation of mechanical design 319 rupture areas and sequence of rupture are marked in white and by numeration. The SIMP design a) shows a collapsing failure directly when the first rupture occurs, due to the simultaneous breaking of the two marked beams" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001644_j.conbuildmat.2020.121961-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001644_j.conbuildmat.2020.121961-Figure4-1.png", "caption": "Fig. 4. End conditions.", "texts": [ " To facilitate the calculation in engineering, the boundary conditions of cable members are usually simplified as hinged boundaries. Previous studies have shown that, in fact, the boundary of a cable member lies between the rigid and hinged joints. To study the influence of end conditions on the mechanical properties of the cable, four types of boundaries, that is, rigid joint boundary, loading beam boundary, simulated cable node boundary, and hinged boundary, are established in this study, which are named as boundary I, boundary II, boundary III, and boundary IV, as shown in Fig. 4. The rigid and hinged boundary models directly apply constraints and loads to the end section. The loading beam boundary is formed by extending the appropriate length at both ends of the cable segment. The load is applied to the cable through the loading beam, which is used to simulate the mechanical behaviour of the cable segment. In engineering, the scattered steel wires are usually inserted into an anchor cup at the end of the cable and are restrained to achieve the anchoring effect. Therefore, this study proposes a simulated cable node treatment method to study the stress situation of the cable end" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000539_s00170-016-9337-2-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000539_s00170-016-9337-2-Figure4-1.png", "caption": "Fig. 4 Blade platform surface", "texts": [ " 20 and 21 are the vertices of the triangles that the step-over are larger than the allowed stepover value. In order to verify the algorithm, we made experiments and simulations using blade machining cases. Selection of parameters for comparison tool paths, the machining trials, and the comparison results are presented in the next section. In the previous sections, we described our field-based tool path generation algorithm to generate smooth tool paths for machining freeform surfaces. To verify that the algorithm can be used in complex surfaces, wemodify the real platform surface (in Fig. 4) to a test surface (with variation of curvatures) and generate tool path for it in Sect. 7.1. Real blade machining experiments are shown in Sect. 7.2 to verify this algorithm. To verify that the hybrid method can further improve the efficiency, we made simulations using VERICUT\u00ae, which is presented in Sect. 7.3. The proposed methods are developed using Visual Studio 2010 as a model of CAM Software. To further verify that the tool path generation algorithm can be used in complex surfaces, wemodify the real platform surface to a test surface (with variation of curvatures), as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002483_tpel.2021.3056287-Figure14-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002483_tpel.2021.3056287-Figure14-1.png", "caption": "Fig. 14. 2-D structure for the studied four phase SRM.", "texts": [ " Restrictions apply. as shown in the fourth column of Table I. During the GA biobjective optimization, thousands of cases need to be evaluated for one operation point. Therefore, substantial time could be saved by the proposed method, which will be verified in the next section. To verify the effectiveness of the proposed method, the dynamic model and bi-objective optimization algorithm were built in the MATLAB/Simulink environment. All the simulation tests are carried out on a four-phase 5.1 kW SRM. Fig. 14 shows the 2-D structure of the studied four-phase SRM, and Table II lists some of the related parameters. It is noted that all the phase current in the simulation and experimental tests are controlled through a widely employed asymmetrical half-bridge converter, which is shown in Fig. 15. Because the proposed method generates the reference current directly, the turn-ON and turn-OFF angles are not required. Owing to the fact that soft current chopping (SCC) method has a low current decreasing rate, the current tracking performance might be poor if the reference current decreases fast" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0003078_ac60055a026-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0003078_ac60055a026-Figure4-1.png", "caption": "Figure 4. Position of Stopcocks during Manipulations", "texts": [ " If the train has not been used that day pass the gas through the entire train, using such additional gas as necessary to remove all air. All air is removed when only microbubbles appear in the azotometer. Shut off the gas flow a t the gas holder, and test the system for leaks by lowering the leveling bulb of the azotometer. A leak will appear as a continuous stream of bubbles. A few bubbles will appear when the bulb is first lowered but will not continue if the apparatus is gas tight. Introduce the sulfamic reagent (0.1 sulfuric acid + 0.1 N sulfamic acid) as follows: (Figure 4 gives the position of the stopcocks of the reactor during this step in the determination. Part A shows the position while the capillary is being filled; part B shows the position after it is filled.) Fill the gas holder, and bring the gas in it to the pressure of the pressure regulator by manipulation of the air-lift stopcock. Heat the solution in the reaction flask hy raising a large test tube filled with water a t the 1018 A N A L Y T I C A L C H E M I S T R Y tioiling point (but not actuallg boiling) around the reaction flask. Cautiously open stopcock B (Figure 4) to the position in part C, and allow the reagent t,o enter the reaction flask. Turn stopcock I3 back to the position in B (Figure 4) and heat the reactor flask again. Now allow just sufficient carbon dioxide to pass into the reaction vessel so that the cooling of the react.or contents does not, cause a partial vacuum. A4110w to cool for 1 minute, and adjust t,he gas flow from the gas holder to a rate such that 1 bubble of gas per second enters the azotometer. It is imperative that this rate be low, or undesirably large bubbles of nitrogen will be formed in the azotometer. When most of the nitrogen has passed into the azotometer as shown by the decreased bubble size, increase the sweeping rate to approximately 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002322_tmech.2020.3047476-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002322_tmech.2020.3047476-Figure5-1.png", "caption": "Fig. 5. (a) Collision detection model of a humanoid robot based on SSCHs, and the SV model of upper arm formed by the self-motion manifold under joint limits, the SV model consists of three convex subparts; (b) A tighter bound of the elbow trajectory is obtained by using two SSCHs with 4 points.", "texts": [ " In the initial stage, we build the collision detection model of the arm using the sphere-swept line (SSL). For the rest of the robot and obstacles, sphere-swept convex hulls (SSCHs) are used to construct their collision detection models. The SSL and SSCHs are pre-computed by using an offline volume optimization procedure. They contain all the points from the CAD model in the lowest possible volume. At runtime, the approximate SVs formed by the self-motion under the joint limits are computed by calculating the SVs of the SSL, which encloses the links of the arm (see Fig. 5(a)). The non-convex SVs must be split into several convex subparts to perform the approximation because the collision detection algorithm proposed in this instance is applicable only to convex bodies. The motion trajectory of the elbow is contained in two SSCHs, as shown in Fig. 5(b). Each SSCH is composed of a radius and four vertices. Thus, the SVs generated by the SSL 1 0( ,{ } )i iV r p = of the lower arm are formulated as follows: 1 4 6 1 0 1 0 1 2 0 3( ,( ,{ } )) ( ,{ ,{ } }) ( ,{ ,{ } })L r i i i i i iSV r p V r p p V r p p= = = + (7) where 4 1 0 1( ,{ ,{ } })i iV r p p = is the SSCH with a radius r and a set of vertices 4 0 1{ ,{ } }i ip p = , and 6 2 0 3( ,{ ,{ } })i iV r p p = is the SSCH with a radius r and a set of vertices 6 0 3{ ,{ } }i ip p = . Then, the collision check pairs are constructed between the swept volumes of the self-motion manifolds and the other collision detection models, and the nearest distance and potential contact points of each collision check pair are computed by using the Gilbert-Johnson-Keerthi (GJK) algorithm" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000327_j.mechatronics.2015.06.014-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000327_j.mechatronics.2015.06.014-Figure6-1.png", "caption": "Fig. 6. Simulation model based rigid\u2013flexible mixed structure.", "texts": [ " (27) yields $fr \u00bc AW Fn0 AW Gn s s1 s2 s6\u00bd T \u00fe Fd0 : \u00f029\u00de Force balance equation of the moving platform can be restated by $F X6 i\u00bc1 $fi $fr \u00bc 0: \u00f030\u00de For active over-constraint parallel manipulator, $F can be achieved by $F \u00bc $g \u00fe $a $z \u00fe $W ; \u00f031\u00de where $g \u00bc md g; md r\u0302p g T represents force screw produced by gravity applied on platform, $a \u00bc md ad; md r\u0302p ad T\u00fe 0; Ided xd Idxd\u00bd T represents inertia force and moment screw caused by motion and $W represents all force screws applied on platform from environment. Substituting Eqs. (10) and (29) into Eq. (30) yields $F \u00bc F0 GF s s1 s2 s6\u00bd T ; \u00f032\u00de where F0 \u00bc Fs0 \u00fe AW Fn0 \u00fe Fd0 , and GF s \u00bc G f s h i \u00fe AW Gn s . When GF s 2 R6 6 is invertible, solution of Eq. (32) can be obtained by s1 s2 s6\u00bd T \u00bc GF s h i 1 F0 $F\u00f0 \u00de: \u00f033\u00de Then the desired driving forces applied on the sliders are obtained. To verify the dynamics with deformation compatibility proposed in the paper, simulation model based on rigid\u2013flexible mixed structure is built, as shown in Fig. 6. Rigid-body dynamics simulation model is built firstly, using the software of Adams. Then utilizing the software of ANSYS, all components with elastic deformations considered are transformed into finite element model. Related geometrical and inertial parameters about the model are shown in Tables 1 and 2. All inertia moments listed in Table 2 are about their own geometrical centers. The tension/compression stiffness of link rods of actuated limbs is 1.45539 108 N/m, and rotational stiffness of the constraint limb is 2", " Making stiffness of limbs 1 and 2 both reduce 100 times, distribution result is shown in Fig. 14. Table 3 Maximum differences of each limb between theoretical calculation and rigid\u2013flexible mixed simulation. Limbs #1 #2 #3 #4 #5 #6 Positive difference (N) 0.63 1.25 0.42 0.81 1.31 0.15 Negative difference (N) 1.04 0.62 1.40 1.26 0.32 1.47 Table 4 Maximum differences of each limb between deformation compatibility and traditional method. Limbs #1 #2 #3 #4 #5 #6 Positive difference (N) 3.22 2.52 2.56 3.17 3.70 2.28 Negative difference (N) 3.15 4.39 4.33 3.17 2.21 3.37 Fig. 6 represents the driving force distribution calculated by the method with deformation compatibility while stiffness of actuated limbs equal. Comparing Figs. 6\u20138 reveals that the method proposed in this article almost has the same result with traditional pseudo-inverse method when actuated limbs has the same stiffness. The differences between these two methods, from 4.69 N to +3.70 N, are mainly caused by the constraint stiffness of constraint limb which is neglected in the traditional method. Generally speaking, rotational stiffness of constraint limb is different from the axial tension/compression stiffness of actuated limb" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002426_j.engfailanal.2020.105211-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002426_j.engfailanal.2020.105211-Figure6-1.png", "caption": "Fig. 6. Radial displacement component at different Angle positions.", "texts": [ " The above formula can be used to calculate the internal load of the bearing. Fr, Fa are the radial and axial forces received by the hub unit. a is the distance between the load center and the center of bearing 1. When the TRB bears radial load Fr and axial load Fa, there will be a radial displacement \u03b4r and a axial displacement \u03b4a correspondingly. The bearing\u2019s stress state and deformation are shown in Fig. 5. Q is the contact load on the roller. \u03b1v is the contact angle between the roller and the outer raceway. As shown in Fig. 6, when the bearing has a radial displacement \u03b4r, for the roller at position Angle \u03d5i, the radial displacement \u03b4ri generated by the bearing outer ring relative to the inner ring is: \u03b4ri = \u03b4rcos\u03d5i (6) When the bearing bears axial load, the axial deformation component of all rollers is the same, that is, the axial displacement of rollers at different angle positions \u03b4ai is equal to the axial displacement of the bearing \u03b4a L.-H. Zhao et al. Engineering Failure Analysis 122 (2021) 105211 \u03b4ai = \u03b4a (7) Therefore, for the roller i, The total displacement upwards along the contact method of the outer raceway can be obtained by (the clearance of the bearing after preloading is 0) \u03b4ni = \u03b4ricos\u03b1v + \u03b4aisin\u03b1v (8) The contact load - displacement relation between roller and outer raceway can be obtained by [22]" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002212_tie.2020.3031535-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002212_tie.2020.3031535-Figure7-1.png", "caption": "Fig. 7. Flux line distributions when armature and PM work separately.", "texts": [ " The 12S22P ST2 delivers 66% and 32% higher torque compared with the 12S10P ST1 and 12S22P ST3, respectively. In addition, the 12S22P ST2 shows the lowest torque ripple when comparing with the 12S10P ST1 and 12S22P ST3. It is worth noting that all the three stators have the same slot/pole combinations. To explain this, the radial airgap flux density with slot-less rotors are compared, as shown in Fig. 6. It can be seen that the machines with different split tooth number have the same PM MMF harmonic order, albeit with different amplitude. Furthermore, Fig. 7 shows the flux line distributions of the three machines when the armature and PMs work separately. When only the PMs are employed, for ST1 machine the flux lines have no radial component in the tooth center adjacent to the airgap; for ST3 machine the flux lines have no radial component in the middle split tooth. The split tooth does not affect the harmonic order of the PM MMF, and thus the rotor pole selection. For ST1, the rotor pole number should be selected as 10 to produce the maximum torque, which is due to that the most dominant harmonic in the PM MMF is 6th and the armature pole pair number is 4", " However, the inductance tendency is not in accordance with the overload capability. This is because the armature MMF passes Authorized licensed use limited to: University of Gothenburg. Downloaded on November 17,2020 at 09:38:58 UTC from IEEE Xplore. Restrictions apply. 0278-0046 (c) 2020 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. through the middle split tooth in ST3 while the PM MMF does not, as shown in Fig. 7(c). Hence, ST3 have additional flux leakage path for the armature MMF, which deteriorates the overload capability. In addition, it can predict that further increasing the split tooth number is not beneficial to the output due to that the number of flux leakage path increases accordingly. B. Influence of Other Geometric Parameters The split-tooth stator slot PM machine with 2 split teeth (ST2) shows the best torque characteristics. Hence, the investigations of the effect of other geometric parameters are based on the 12S22P ST2" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001253_s40435-020-00629-8-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001253_s40435-020-00629-8-Figure1-1.png", "caption": "Fig. 1 Schematic diagram of the ball and plate system in x-direction", "texts": [ " This paper is organized as follows: The system structure and the conversion to the strict feedback form are presented in Sect. 2. Section 3 presents the ADRC based estimation of the uncertain model and derivatives, the cascaded interaction control design procedure and the theoretical results. Experimental setup and results are presented in Sect. 4. Section 5 contains the conclusions. The ball and plate system makes use of two same rotary servo base units made by Quanser. The plate is symmetrical. The dynamics along each axis is therefore assumed to be the same. As shown in Fig. 1, the motor drives the servo angle \u03b8 to further control the displacement x of the ball through the movement of the angle \u03b1 of the beam. The system is a typical cascaded, under-actuated system. From Newton\u2019s second law, we can state that the dynamic equations for the displacement of the ball and the output angle of theDCmotor can bewritten in the following general form: x\u0308 a11x + a12\u03b8 + a13 x\u0307 + a14\u03b8\u0307 + b1u + 1 \u03b8\u0308 a21x + a22\u03b8 + a23 x\u0307 + a24\u03b8\u0307 + b2u + 2 (1) where 1 and 2 denote the unknown nonlinear parts of the system and the disturbances, u is the control and represents the input voltage Vm of the servo motor" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0003137_1.1752269-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0003137_1.1752269-Figure5-1.png", "caption": "FIG. 5. Typical scale reading, amplitude number curves as the temperature is increased (a to d).", "texts": [ "124 On: Wed, 03 Dec 2014 21:07:45 The amplitudes are always observed at the same side and are consecutively numbered. n is the number of full periods. For practical reasons in a preliminary experiment the approximate values of the amplitudes are noted. As in the final experiments, counting the first observed amplitude as zero, two consecutive amplitudes are omitted and every third is noted. With a little experience tenths of scale units may be estimated. The values of the amplitudes in scale units are plotted as ordinates (see Fig. 5d) and the corresponding amplitude numbers as abscissae, the omitted amplitudes entering into the axis of abscissae with their respective num bers. In Fig. 5 the uppermost point corresponds to A o, the lowest to An; the number of observed amplitudes is 10, the number of the omitted, unobserved, is lS, hence the number of fuIl periods is 27, i.e., equal to the number of intervals between the points. If all the points, perhaps with exception of the first and largest amplitude (A 0), which may be slightly in error because of the primitive release, are lying on a straight line without any systematic deviation, we have pure nonviscous friction. Such straight lines are reproducible to within 3 percent for the same substance even after some wear of the pendulum and between 200 e and SO\u00b0c. Fig. 5 a to c show the case where the straight lines are chords of curves on which the observed points lie, i.e., the presence of viscous friction, notwithstanding the invisible thickness of the lubricating film. The more viscous the lubricant is, the easier this effect occurs. The electric heater minimizes these lubrication difficulties. In general it is possible with its help, without altering the lubrication, to suppress viscous and capillary effects to such a degree that on re peated oscillating the effects disappear and a straight line is obtained (Fig. 5a-d). The first straigh t line in such a series yields the correct coefficient of friction. This course of an experi ment is perhaps the safest way of obtaining the coefficient characteristic of the lubricant and not any other nonviscous coefficient produced by particles of wear, etc. In most cases such higher friction values, the extreme of which is the value for the unlubricated pendulum, may be obtained on prolonged operation of the pendulum. It appears, that the ease with which they are obtained depends somewhat on the oil. A sys tematic investigation might lead to a method of testing film toughness. For practical lubrication considerations it is worth mentioning that experimenting with the apparatus demonstrates that excluding viscous friction seems to be much more of a problem than realizing it. The example of Fig. 5d yields by means of Eq. (S') the value p=O.112S. For Ao the value from the straight line was taken. The main fields of application of the apparatus are lubrication research and practice and the physics of adsorption. With slight alterations the apparatus may be used for measuring the coeffi cients of friction of various materials. II In the first section it is pointed out that the formula used to find the coefficient of friction is an approximation. The simplicity of the result permits its use, but it is necessary to develop the accurate method and from that to show that the formula (S') may be safely used for relative measurements", "124 On: Wed, 03 Dec 2014 21:07:45 function of 'Y is a straight line and then find the variation of the value of p as measured at different values of a'Y. To give a numerical ex ample the assumption is made for the straight line that a'Y-a-y+l = const = 30'. According to Eq. (8') this corresponds to a value of 0.1132 for p. Fig. 7 shows the values of p calculated from Eq. (22) for different values of a'Y Extrapo lation of this curve to zero gives a value of p=0.1146. This value of p is within one percent of the value given by Eq. (8'). The method, therefore, becomes what of finding the equation of the straight line (Fig. 5d) and from this accurate equation calculating p from the lower part of the straight line. A further discussion of the curve in Fig. 7 is necessary, since thequestion is raised as to why the value of p found by extrapolating the curve to zero is any better than any other value of p given from the curve. This goes back to the fact that Eq. (22) does not give a straight line rela tion, which means that if the measurements are made from a straight line a viscous friction effect must also be present. This viscous effect may be shown to be negligible and to have negligible effect upon the value of p determined by ex trapolating the curve in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000718_978-1-4471-7278-9-Figure15.12-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000718_978-1-4471-7278-9-Figure15.12-1.png", "caption": "Fig. 15.12 Slice-based feature detection based on the polar coordinate profile", "texts": [ " Based on the simulated image samples of known materials, we can train an artificial neural network (e.g. Radial Basis Function) to recognize the signature intensity. To adapt to a broader range of data and noise levels, the data can be preprocessed with normalization algorithms and filters. Surface-based detection: In the continuous case, curvature is defined as the rate of change of slope. In our case, the discrete space, the curvature description must be slightly modified to overcome difficulties resulting from violations of curve smoothness (Fig. 15.12).15 We start by slicing the digital model horizontally and then averaging the points between the slices. The curvature scaler descriptor calculates the ratio between the total number of boundary pixels (length) and the number of boundary pixels where the boundary direction changes significantly. The smaller the number of direction changes, the straighter the boundary. In this case, we map the points on the slice 15Goldgof DB, Huang TS, Lee H (1989) A Curvature-Based Approach to Terrain Recognition, November 1989 (Vol" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000252_j.promfg.2020.01.002-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000252_j.promfg.2020.01.002-Figure6-1.png", "caption": "Fig. 6. Simulation of AM layer toolpath in STEP-NC Machine.", "texts": [ "com/StepNcLadprer/RepRap-machine-modelPrusa-Mendel-for-STEP-NC-Machine.git. An alternative way to this method may be to include the kinematic model directly in the STEP document by using the data structure of the AP-242. In A4, virtual simulation for toolpath verification is performed by opening the AM STEP-NC program into STEP-NC Machine. Simulation is appended by loading the AM machine\u2019s kinematic model from pull-down list \u201cMachine Tool\u201d and importing the shape geometry for the tool (hotend for 3D printing) from the tab \u201cTool\u201d as shown in Fig. 6. Information of part shape geometry can also be included from the tab \u201cPart Properties\u201d in STEPNC Machine. Thus, besides verifying manufacturing toolpath, the AP-238 STEP-NC program file can be enriched with more data of tool and part geometry. That is because the AP-238 supports integration with other STEP resources for representation of 3D-geometry and GD&T data. That gives possibilities to perform quality checks of the part considering the need for corrections or surface finishing. A G-code program, understood by the CNC controllers of legacy AM machines, is generated by using postprocessor of STEP-NC Machine to manufacture the test part (Activity A5 of the IDEF0)" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001312_s00170-020-05434-3-Figure18-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001312_s00170-020-05434-3-Figure18-1.png", "caption": "Fig. 18 The fabricated samples with different scanning paths", "texts": [ " However, for the narrow side of the cladding layer formed by changing scanning speed, the total height is low due to less accumulation of materials, which is similar to the geometric morphology of single track, so it is not easy to distort. For laser direct metal deposition, the scanning path is very important and directly affects the geometric accuracy and performance of the formed parts. In the forming process of thinwalled parts with variable width, different scanning paths also directly affect the geometry of the formed parts. The forming samples of parallel scanning and reverse scanning are shown in Fig. 18, respectively. It can be seen from the figure that the test piece formed by the parallel scanning has a high uniformity, but a tilt collapse phenomenon at the narrow end. For the sample formed by the reverse scanning, the sample is uniform in height. The formed part has no obvious defects and the forming result meets the requirements. The experienced phenomenon can be explained based on the consideration of cladding path. As shown in Fig. 19, a schematic of the laser direct metal deposition is shown" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002702_iciccs51141.2021.9432177-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002702_iciccs51141.2021.9432177-Figure1-1.png", "caption": "Fig. 1. A real photo of a bearing.", "texts": [ " The fault unless detected at an early level, it can lead to catastrophic results in terms of unfavorable maintenance time and cost, decrease in production level, and severe risk of loss of human lives [2]. A bearing's primary function in rotating machinery is to sustain the load and minimize friction. The outer race, inner race, cage, and balls are the main components of the bearing. The cage keeps the balls in place within the inner and outer races, allowing for free and smooth rotation [3]. A real image of a bearing is shown in Fig 1. Different parts of the bearing are shown in Fig 2. Concerning the possibility of induction machine defects, bearing faults are the most common fault type, accounting for more than 30 percent of all failures. In addition, this fault is the common reason for the motor shut down, causing in a major loss of property and safety [4, 5]. Pertaining to the previous reasons, bearing fault diagnosis is a crucial part of development as well as engineering research. Fault detection and condition monitoring of the machine provide significant information about its status during the operation at each moment, keeping the working condition" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002186_icra40945.2020.9196692-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002186_icra40945.2020.9196692-Figure1-1.png", "caption": "Fig. 1: (a) The pictorial representation of the robot design with major components. (b) The free-body-diagram of the mobile robot and the principle of actuation. (c) The depiction of the path-planning-based control problem with a representative medical application of a tumor region, which requires simultaneous reachability of a desired 5-DoF position and orientation.", "texts": [ " Therefore, pitch angle is not reachable to any desired configuration, which leads the experiments to be controlled in 4-DoF. In this study, we use an SSND-type robot with the following major properties: (1) The density of the robot is very close to the density of the fluidic medium; (2) Center of mass (CoM) and center of the volume (CoV) positions are very close to each other; (3) The spherical permanent magnet is located in a spherical cavity in a free-to-rotate form; (4) The permanent magnet is off-centered from CoM of the robot [29]. The robot design is depicted in Fig. 1a. Basic robot components, as well as the assembled final version of the robot, are provided in Fig. 2. The resultant robot parameters are given in Table I. In order to emulate the magnetic field environment of an MRI system, a custom-built pseudo-MRI system is used (Fig. 3). This system consists of a (1) commercial MRI gradient coil insert, (2) a water-cooled uniform one-axis field generator, and (3) a motion stage. The MRI gradient coil insert used in this study can create a homogeneous 200 mT/m maximum instantaneous magnetic field gradients and can provide 66 mT/m maximum magnetic field gradient continuously at around 6 cm away from the bore center", " The feedback rate of the system is fixed to 10 Hz in order to emulate the low-feedback rate condition of MRI-powered robotic actuation systems. The untethered magnetic robot presented in this study performs rigid body motion in 3D fluidic environments. The state of the robot can be defined as x = [ r q v \u03c9 ]T (1) where r = [ x y z ]T represents the position of the CoM in the inertial frame. The objective is to control the 5- DoF position and orientation (pitch and yaw) of the robot as shown in Fig. 1c. In order to avoid any singularity, the quaternion representation, q = [ qw qx qy qz ]T , is used, which defines the orientation of the robot in the inertial frame. v = [ u v w ]T and \u03c9 = [ p q r ]T are translational and rotational velocities described in robot\u2019s frame, respectively. The major forces acting on the robot are magnetic force, Fm, gravitational force, Fg, the buoyancy force, Fb, and fluidic forces and torques due to the relative velocity of the robot with respect to the stationary fluid, F f and T f ", " The robot is neutrally buoyant if these two forces are exactly equal in magnitude. Due to the manufacturing errors and the slight imbalanced design, CoM and CoV locations are non-coincident and the robots are not perfectly neutrally buoyant. Manual assembly, 3D printing accuracy, insufficient controllability on color coding and sealing with lacquer during the manufacturing of the robot are the main causes contributing to such manufacturing errors [29]. Due to the discrepancy between the CoM and the CoV shown in Fig. 1, a buoyancy torque might cause pitching and rolling. Additionally, if there is a relative motion between the robot and the fluid, then, the robot also experiences drag force and torque. These forces and torques are computed with COMSOL numerical simulations for the robot [29]. The continuous time optimal control problem in the absence of terminal state cost is formulated as J\u2217 =min u(t) \u222b T 0 L (x(t),u(t))dt s.t. x\u0307 = f (x(t),u(t)) x(0) = x0, (12) where J\u2217 is the optimal cost, T is the terminal time, L (\u00b7, \u00b7) is the Lagrangian, and u(t) is the control input" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001403_j.ymssp.2020.107075-Figure21-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001403_j.ymssp.2020.107075-Figure21-1.png", "caption": "Fig. 21. Axial and radial loading direction of the bearing.", "texts": [ " Therefore, such parameters as raceway curvature coefficient and misalignment values should be considered during the design stage of bearing. The geometric analysis above is simple and effective in the initial design stage of bearing parameters, but the operating condition and load state of the bearing are not considered, which may lead to an unexpected error. The dynamic modal, proposed in this paper, is applied to study how the radial and axial loads affect the contact state of the bearing and loading state of non-main loading inner-half ring through a comparative analysis of simulation and experiment. Fig. 21 shows the load direction of the bearing (Table 3) with shim width gi \u00bc 0:236mm, which is applied to analyze the contact force of non-main loading inner-half ring and thus clarify the effect of different loading conditions on the three-point abnormal contact. Fig. 22 illustrates the contact force variation of non-main loading inner-half ring when the rotate speed of the inner ring is 600r/min. From the figure, the contact force is nonzero and gradually increasing with the decrease of axial load and the increase of radial load, which means that the ball is in contact with the non-main loading inner-half ring, and thus the three-point abnormal contact comes into being" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000962_s11740-019-00916-0-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000962_s11740-019-00916-0-Figure4-1.png", "caption": "Fig. 4 The part geometry used in model validations, adapted from [10], a program file; b SLA printed NIST part from program file in Fig.\u00a06a", "texts": [ " Larger geometries will lead to longer print times and thus can increase energy consumption during a print. Additionally, different geometry sizes will alter material consumption and waste. Layer thickness is another critical parameter to be considered as it affects the total time it takes to print a part as well as the surface quality of the finished part. To validate the proposed UPLCI model, experiments are performed to investigate the energy flow and material flow in an SLA process when fabricating the NIST AM test artifact shown in Fig.\u00a04a, b. In this study, an SLA test bed is developed and used, which contains several components including a projector as the UV source, a motor that powers the building platform, a control board, control software, and a material tank. A Yokogawa CW10 power meter with 0.01\u00a0W resolution is used to measure the power consumption, a Honeywell MiniRAE 3000 VOC monitor is used to measure the real-time VOC emissions from the liquid resin, and a Taishi 200 digital scale is used to quantify the material consumption and waste generated" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001628_0954406220976154-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001628_0954406220976154-Figure5-1.png", "caption": "Figure 5. Test-bed of experimental setup.", "texts": [ "22 Gradientbased methods are widely used to fit both preceding and subsequent parameters in the ANFIS model.45 One of the problems with gradient methods is that the response is situated in the local optimality, and the rate of convergence is too slow. The algorithms of metaheuristic optimization such as PSO is used as an efficient solution into problems related to methods that are based on gradient. The procedure for ANFIS model training using PSO is given in Figure 4.46,47 In this study, vibration signals were collected from a 2:2MW high-speed wind turbine48 as shown in Figure 5. Measurements were taken over 50 consecutive days using MEMS-based accelerometers mounted radially on the bearing support ring. The input shaft bearing speed is 1800 rpm, while the sampling frequency is equal to 97,656Hz, with an acquired signal time of 6 s for a real-world highspeed shaft bearing from a Wind Turbine Gearbox (WTG) provided by Green Power Monitoring Systems in USA. The bearings type is SKF 32222 J2 Tapered Roller Bearings(TRB). The TRB has an outside diameter of 200mm, a bore of 110mm and an overall length of 56mm" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000268_j.measurement.2015.04.025-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000268_j.measurement.2015.04.025-Figure2-1.png", "caption": "Fig. 2. Grid points around the lathe machine.", "texts": [ " As the output of the accelerometer is of low level and contains some unwanted frequencies, some form of pre-processing is required before analyzing the data. A 4-input, 1-output vibration collector system [23] (Spider-81 Vibration Controller System) amplifies the output data of the accelerometer and converts the data into frequency domain using Electronic Data Management (EDM) software. The entire process cycle for vibration data acquisition is shown in Fig. 1. The acoustic signatures are obtained using a noise level meter [24]. Fourteen equidistant grid points are selected around the lathe machine as shown in Fig. 2 and the sound level is measured at each point for every lathe machine. The noise level meter is kept at a distance of 0.2 m from the base of lathe machine and the readings are taken in the absence of any external noise field. The Noise level meter consists of a transducer, preamplifier, amplifier and an analysis module. A condenser microphone is used as a transducer for measuring the sound pressure level which is directly obtained on a readout screen. An FFT module converts the time domain data into frequency domain" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000538_tcns.2016.2609638-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000538_tcns.2016.2609638-Figure1-1.png", "caption": "Fig. 1. Interpretation of multi-agent coordination for three agents", "texts": [ " K Sakurama is with Graduate School of Engineering, Tottori University, 4-101 Koyama-Minami, Tottori-shi, Tottori, 680-8552 JAPAN (e-mail: sakurama@mech.tottori-u.ac.jp). S. Azuma and T. Sugie are with the Graduate School of Informatics, Kyoto University, Yoshida-honmachi, Sakyo-ku, Kyoto 606-8501 JAPAN (e-mail: sazuma,sugie@i.kyoto-u.ac.jp). ample is the consensus problem, which can be described by the convergence of x(t) to T = {x \u2208 Rn : x1 = x2 = \u00b7 \u00b7 \u00b7 = xn}, namely a line. In three-agent case, this problem is interpreted as Fig. 1 (a), where each axis corresponds to the each agent\u2019s state xi (i \u2208 {1, 2, 3}). Thus, the consensus problem is characterized by the one-dimensional target subspace. The second example is equally-spaced coordination shown in Fig. 2, which is a task to align the agents with the same interval. This coordination is described by a plane of T as sketched in Fig. 1 (b), namely a two-dimensional target subspace. In this way, various coordination tasks can be described via a \u03c4 - dimensional target subspace for a natural number \u03c4 . Typical coordinations are shown in Table I. For networked multi-agent systems, it is important to know what kind of network topology enables us to achieve a given task via distributed and relative control. Actually, it strongly depends on the network topologies whether the coordination to \u03c4 -dimensional target subspaces is achievable or not. It is known that the connectedness is required for consensus [11], [12], namely a one-dimensional target subspace. However, TABLE I EXAMPLES OF COORDINATION TO \u03c4 -DIMENSIONAL TARGET SUBSPACES \u03c4 Shape of T Type of coordination 1 Line (Fig. 1 (a)) \u2022 Consensus ([12], Example 1) \u2022 Position-based formation [27] 2 Plane (Fig. 1 (b)) \u2022 Equally-spaced coordination (Fig. 2) \u2022 Scale-free formation (Example 2) 3 3-dimensional space Alternately-spaced coordination (Example 3) 2325-5870 (c) 2016 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. it is not enough for higher-dimensional target subspaces in general. For example, as shown in this paper, coordination to two-dimensional target subspaces including the equally-spaced coordination can be achieved over the graph in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000616_s11071-016-3242-y-Figure14-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000616_s11071-016-3242-y-Figure14-1.png", "caption": "Fig. 14 Two-degrees-of-freedom robot", "texts": [ " From obtained results, it can be concluded that proposed control scheme provides perfect model following with improved accuracy in estimation using smaller values of \u03c4f and overcomes the problem of coupling phenomenon completely. The estimation accuracy can also be improved by using higher-orderfilters. This shows the effectiveness of proposed control scheme. In this section, application of proposed control scheme for a two-link robotic manipulator has been performed and validated. The schematic of two-link manipulator is shown in Fig. 14. The dynamics of two-link manipulator are taken from [42] and given by [M(q) + M(q)] q\u0308 + [C(q, q\u0307)+ C(q, q\u0307)] q\u0307 + [G(q)+ G(q)]+[F(q, q\u0307)+ F(q, q\u0307)]=\u03c4 +d (47) M(q)11 = l22m2 + 2l1l2m2 cos(q2) + l21(m1 + m2) M(q)12 = l22m2 + l1l2m2 cos(q2) M(q)21 = M(q)12 M(q)22 = l22m2 C(q, q\u0307)q\u0307 = [\u2212m2l1l2 sin(q2)q\u030722 \u2212 2m2l1l2 sin(q2)q\u03071q\u03072 m2l1l2 sin(q2)q\u030722 ] G(q) = [ m2l2g cos(q1+q2)+(m1+m2)l1g cos(q1) m2l2g cos(q1+q2) ] F(q, q\u0307) = [ Fv1q\u03071 + Fc1sgn(q\u03071) Fv2q\u03072 + Fc2sgn(q\u03072) ] The l1 and l2 are the lengths of the link, m1 and m2 are the masses of the joints, g is the gravitational acceleration, Fv1, Fv2 are viscous friction coefficient and Fc1, Fc2 are Coulomb friction coefficient for link 1 and 2, respectively, and \u03c4 = [\u03c41, \u03c42]T is an input torque vector" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000745_iet-map.2018.6170-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000745_iet-map.2018.6170-Figure6-1.png", "caption": "Fig. 6 Distributions of (a) The E-field,", "texts": [ " By setting up another Itype metal strip along the y-direction, a dual-polarised bandabsorptive FSR is designed and presented in Fig. 5. Two cross Itype metal strips are printed on the same surface of the substrate. In order to prevent overlap of the lumped resistors, one resistor is welded on the bottom surface of the lossy layer and connects to the corresponding I-type metal strip by two metalised vias. The lossless layer is constructed by Jerusalem-cross metal strip. E-field and currents distributions of the dual-polarised bandabsorptive FSR are presented in Fig. 6. As shown in Fig. 6(a), a coupling capacitance is generated between the two cross I-type metal strips. Its currents distribution is shown in Fig. 6(b). According to the distributions of E-field and currents, ECM of the dual-polarised FSR is established and its detailed circuit parameters is shown in Fig. 7. Ra, La and Ca are similar to those of the single-polarised band-absorptive FSR and they are formed by the I-type metal strip along the E-field direction. Lb is an equivalent inductance of the I-type metal strip perpendicular to the E-field direction. The equivalent capacitance Cb is generated between the two cross I-type metal strips. (a) Perspective view, (b) Top view, (c) Bottom view A B C D = 1 0 1/Za 1 cos \u03b8p jZpsin \u03b8p jsin \u03b8p/Zp cos \u03b8p 1 0 1/Zb 1 = cos \u03b8p + j Zp Zb sin \u03b8p jZpsin \u03b8p Za + Zb ZaZb cos \u03b8p + j 1 Zp + Zp ZaZb sin \u03b8p cos \u03b8p + j Zp Za sin \u03b8p (1) 1778 IET Microw" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001738_j.mechmachtheory.2019.103629-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001738_j.mechmachtheory.2019.103629-Figure4-1.png", "caption": "Fig. 4. Inverse singular configuration in which Link-1,1 is perpendicular to x -axis.", "texts": [ " These kinds of singularities typically identify boundaries of the workspace [45] , unless the workspace is limited by other factors such as mechanical interference. Referring to Eq. (25) , inverse kinematic singularities arises when one of the diagonal elements of the inverse Jacobian matrix vanishes: \u02c6 i \u00b7 l i, j = 0 (40) The left-hand side of the latter equation represents the magnitude of l i , j along x -direction. Consequently, an inverse kinematic singularity occurs whenever any link of an arm is perpendicular to the x -axis. Fig. 4 illustrates such a singular configuration of the manipulator under study. In this type of singularities, by locking the end-effector, the i th actuator can have an infinitesimal motion and an infinitesimal movement of the end-effector along particular direction cannot be accomplished. Direct kinematic singularity arises when the direct Jacobian matrix ( J D ) is rank deficient. In this type of singularities, as opposed to the other one, the end-effector of the manipulator gains one or more uncontrollable degrees of freedom while the actuators are locked" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002086_tia.2020.3015693-Figure23-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002086_tia.2020.3015693-Figure23-1.png", "caption": "Fig. 23. The VF IPMSM with six probes along the length of PM.", "texts": [ " The 18th harmonic is generated due to the 17th and 19th harmonics which are the slot harmonics in the back EMF. The FEA results show that the harmonics are minimized with the 3-step and the 5-step skewed PM-pole. VI. DEMAGNETIZATION OF PMS A negative d-axis current pulse is applied to demagnetize the PM-pole permanently in the field weakening region above the base speed. Thus, the efficiency of a VFM can be enhanced when operating above the base speed compared to a PMSM. Six probes are placed along the length of each PM- pole to investigate the demagnetization as shown in Fig. 23. Demagnetization of each step of the skewed PM-pole has to be investigated because there is a difference in the angular position between the steps of a skewed PM-pole. A high value of angular deviation between each step causes non-uniform demagnetization of the step skewed PM-pole. So the minimum optimum skewing angle, 20/3\u2070 is chosen to minimize the cogging torque and the torque ripple. The magnetization level of the PM-pole is quantitatively evaluated by measuring the back EMF waveform after magnetization or demagnetization", " Demagnetization of the unskewed PM-pole is easier compared to the skewed PM-pole. The applied demagnetizing current pulse is shown in Fig. 24. The flux density of the unskewed PM-pole before and after applying the demagnetizing current pulse is shown in Fig. 25. The demagnetizing current pulse brings the magnetic flux density close to zero which ensures complete demagnetization of the unskewed PM-pole of the VF IPMSM. The FEA results in Fig. 26 show that both the PMs are uniformly demagnetized. Fig. 23. VF IPMSM with six probe along the length of PM. Authorized licensed use limited to: Cornell University Library. Downloaded on September 01,2020 at 03:48:51 UTC from IEEE Xplore. Restrictions apply. 0093-9994 (c) 2020 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. Each step of the 3-step skewed PM-pole in the VF IPMSM has to be demagnetized uniformly. Analytical model of the local demagnetization in VFM due to armature winding MMF was analyzed in [27]" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000442_tsmc.2016.2560528-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000442_tsmc.2016.2560528-Figure4-1.png", "caption": "Fig. 4. Preserving the structural controllability after adding a new node. Existence of one link (or more links) of links I to VI is necessary to preserve the structural controllability of the system after adding new distinct node 7.", "texts": [ " In contrast, in the structural controllability of network systems with dynamics and self-loop (see Fig. 3), controlling node 1 is sufficient for full control over the network system. Using Theorem 2, we can state Corollary 1 on the scalability of a controllable network system. Corollary 1: Consider a controllable network system. The network system can preserve the structural controllability after adding a new node, if the node is distinct and there is a path from any node of the network system to the new node. An example is presented in Fig. 4 to illustrate Corollary 1. It shows a controllable network which consists of six nodes. Node 7 is a distinct new node which is added to the system. Hence, to preserve the structural controllability, the existence of one or more link(s) of links I to VI is a requirement. To show the structural controllability of a network system with a globally reachable driver node in the proof of Theorem 2, it is assumed that the diagonal elements of network system matrix A, are distinct. Next, we develop the results to a more general case that the network system contains identical diagonal elements of network system matrix" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001717_jmes_jour_1959_001_016_02-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001717_jmes_jour_1959_001_016_02-Figure2-1.png", "caption": "Fig. 2. Geometry of &formation Original position. - - - Final position.", "texts": [ " Dahl (3), de Lieris (6), Turner and Ford (14), and Salzmann (11), have all examined, mainly by strain energy methods, certain aspects of various expansion bellows shapes, including toroidal elements, without studying the torus per se. Internal pressure loading of a complete torus has been examined by Dean (4). The membrane solution for internal pressure by Foppl (5) is of course well known and is discussed, for example, by Timoshenko (13). Nakamura (9) has recently studied internal pressure in vessels with torispherical heads. DERIVATION OF GOVERNING EQUATIONS The governing equations may be found in terms of various quantities. They are briefly derived here in terms of the rotation 4 of the shell meridian, Fig. 2, and the horizontal force F per unit length, Fig. 1, since these are convenient variables in some engineering applications. From the initial geometry, Fig. la, it is seen that, in the limit and d r = d S i n e . . . . (1) * A numerical list of references is given in Appendix 111. t The topic is erroneously stated to be corrugated pipes by Timo- shenko .( 13). J O U R N A L M E C H A N I C A L E N G I N E E R I N G S C I E N C E EQUILIBRIUM Considering only small elastic deformations which do not substantially change the six and shape of an element of shell wall, Fig", " (5) The horizontal and vertical components F and IV can be related to the normal and shear forces at any section by N,= wcose+Fsine ~ = w s i n e - F c o s e . . . (6) $ It must be borne in mind that X may well be a function of radius, for exampZe arising from cenm&gal force, and p may well be a function of depth (and hence of radius in general), for example, hydrostatic pressure. Vol I No 2 1959 at UNIV OF VIRGINIA on June 5, 2016jms.sagepub.comDownloaded from STUDY OF THE SYMMETRICAL ELASTIC LOADING OF SOME SHELLS OF REVOLUTION 115 GEOMETRY OF DEFORMATION Analysis of the geometry of deformation, Fig. 2 4 shows that the horizontal (radial) movement u is related to the rotations 4, and meridional strains C?, by du dr - = - 4 a t e + ~ ? , . . . (7) where the - notation refers to mid-wall values. The mid-wall circumferential strain is U ZS = ; . . . . (8) and the variation of these strains through the wall thickness, Fig. 2b, gives d4 e, = C,+Z a . . . . (9) \" e, = C*+Z- sine . . . (10) r Elastic laws For a Hookean material, neglecting for thin-walled shells the stresses through the wall thickness, then 1 1 e, = E ( P e - V P * ) eS = E(IJ,--vP,) . . . - (11) These equations may also be written with the - notation to refer to the mid-wall values. MANIPULATION OF EQUATIONS Using equations ( 1 x 1 1) just derived, the bending moment m, per unit length of wall is -h Similarly the circumferential bending moment mS per unit length is m4 = D s i n 8 (;+vZ) 4 d\" ", " It is also found that the sin 20 terms on the right-hand side of equations (16) and (17) are usually of order 72 smaller than corresponding sin 0 terms*. Performing the above simplifying steps and writingt kvV kpb2(2-v)773/2 2hX(3+v)b2$/2 B=-+ + the equations are finally obtained in the form . . (18) P Pr)'l2 2P d2P d82-ppco~e= Asin8 . . (19a) and in this form the equations have been used for the following calculations. The homogeneous part of these equations is of the same form as that used by Nakamura based on Reissner's solutions*. Deflections and stresses The vertical deflection of an element of the shell wall can be seen from Fig. 2a to be whence Thus for a torus, the deflection between points dw=+dzsine+codzcose dV/& = ++So cot 6 and O2 is V = - - ~ b +sine+e;,cosede . (20) s: The horizontal (radial) deflection u is, similarly, 02 80 u = +cos--e;,sinede . (21) The second term in these equations is usually much smaller than the first, for the type of problem considered here. Alternatively, u may be found from the strains and stresses by equation (8). The slope 4 is given by bP + = r ) ~ ( q + k cos e y 2 * t $ See further discussion, p " ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001440_physrevfluids.5.084004-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001440_physrevfluids.5.084004-Figure1-1.png", "caption": "FIG. 1. Schematics of (a) experimental and (b) computational setups depicting, respectively, a fully submerged hemispherical and a half-submerged spherical surfer at a water-air interface. The area of the hemisphere dip coated with a layer of Dawn soap is colored red in panel (a), similar to the active area of the sphere in panel (b). The color map and vector plots in panel (b) represent the concentration distribution and liquid velocity field in the vicinity of the surfer, respectively.", "texts": [ " In what follows, we first describe the experimental setup (Sec. II), and then explain the simulation details (Sec. III). The results are discussed next (Sec. IV), and concluding remarks are given in Sec. V. The experiments involve fabricating a cylindrical disk of radius R = 2.25 mm and thickness \u03bb = 1.5 mm, and a hemisphere of radius R = 2.375 mm, both from polydimethlysiloxane, which has a density of \u03c1 = 965 Kg/m3. The disk and hemisphere are dip coated into a solution of water and 50% soap (Dawn) to a depth of d = 1 mm, as shown in Fig. 1(a). This process results in an asymmetric coating of the particle with a layer of an active agent (i.e., Dawn soap), which is known to reduce the interfacial tension of water from 72 to \u224830 mN/m [52]. The disk or hemisphere is gently placed on the water-air interface and released using tweezers. The dissolution of the coating layer creates a surface tension imbalance that, in turn, leads to the propulsion (also referred to as surfing) of the particle. The experiments are conducted in a clear-walled rectangular container of length L = 560 mm and width W = 45 mm, that is filled with water [see Fig. 1(a)]. To investigate the 084004-2 effect of confinement on the motion of the particles, the depth of water is varied from H = 1.8 to 10 mm. We employ particle image velocimetry to analyze the flow field at the water-air interface and beneath the surfer. PIV measurements are performed using uniformly dispersed tracer particles of 40-\u03bcm diameter that are illuminated by a 300-mW argon-ion laser light sheet of thickness 1 mm. The laser is oriented either parallel to the interface just below the surface (in order to obtain the interfacial velocity profiles) or normal to the surface passing up through the fluid from the bottom of the container [see Fig. 1(a)]. Through this second configuration, the velocity field underneath the surfer is measured. A high speed camera at a maximum frame rate of 110 fps is utilized to capture the images, which are then processed via a commercial PIV software package developed by LaVision. Also, a digital camera is used to track the motion of the surfer. The videos are captured at 17 fps and fed into a particle tracking software (Tracker) to measure the speed and direction of the particle\u2019s propulsion. For the purposes of the simulations, we model the motion of a particle located at the interface between a semi-infinite layer of gas and a layer of liquid\u2014with density \u03c1, viscosity \u03bc, and surface tension \u03b3 \u2014that is bounded from below by an impermeable solid wall at a distance H from the interface [see Fig. 1(b)]. Consistent with the experiments, the movement of the particle is caused by an asymmetric release of a chemical species\u2014with diffusivity D\u2014from the \u201cactive\u201d region of the particle\u2019s surface, which we denote Sa p [see Fig. 1(b)]. The shape of the particle is considered to be an oblate spheroid of equatorial radius R and aspect ratio (ratio of polar to equatorial radius) \u03b5. We also consider, for the sake of comparison with the experiments, a cylindrical disk of radius R and half-thickness-to-radius ratio of 2/3. To reduce the complexity of numerical calculations, the following justifiable assumptions are made. (i) The liquid-gas interface is flat. (ii) The particle is half submerged forming a 90\u25e6 contact angle. (iii) The released chemical species is soluble into the bulk of the liquid layer and its concentration is constant at Sa p", " (iv) The liquid is Newtonian with constant density and viscosity that are unaffected by the presence of the solute. 084004-3 (v) The surface tension of the liquid varies linearly with the concentration of the active agent. (vi) The particle undergoes a pure translational motion along a straight line parallel to the interface. Let r = xex + yey + zez be the position vector in the Cartesian coordinate system (x, y, z), with the unit vectors ex, ey, and ez such that ex is parallel to the direction of the surfer\u2019s propulsion and ez is normal to the interface pointing away from the liquid [see Fig. 1(b)]. Also, let u(r, t ) = uxex + uyey + uzez, p(r, t ), and c(r, t ) represent, respectively, the velocity and pressure fields of the liquid and the concentration distribution of the chemical species, where t denotes the time variable. With these definitions and the above-mentioned conditions, the equations that govern the spatiotemporal evolution of u, p, and c are \u03c1 ( \u2202u \u2202t + u \u00b7 \u2207u ) = \u2212\u2207p + \u03bc\u22072u and \u2207 \u00b7 u = 0, with u(r, 0) = 0, u = Uex for r \u2208 Sp, u = 0 for r \u2208 Sw, uz = 0 and (I \u2212 nn) \u00b7 (n \u00b7 ) = \u03bc ( \u2202ux \u2202z ex + \u2202uy \u2202z ey ) = \u2212\u2207s\u03b3 = \u2212 ( \u2202\u03b3 \u2202x ex + \u2202\u03b3 \u2202y ey ) for r \u2208 Si, (1) \u2202c \u2202t + u \u00b7 \u2207c = D\u22072c, with c(r, 0) = 0, c = cs for r \u2208 Sa p, n \u00b7 \u2207c = 0 for r /\u2208 Sa p, (2) where = \u2212pI + \u03bc[\u2207u + (\u2207u)T ], \u03b3 = \u03b30 \u2212 \u03b1 c", " In particular, the PIMPLE algorithm is employed to treat the pressurevelocity coupling, the Laplacians are discretized via the second-order linear Gaussian integration, the corrected scheme (with the number of corrections set to 2) is used to calculate surface normal gradients, the time derivatives are approximated by the second-order backward differentiation formula, and the equation of motion for the surfer is integrated using the Newmark method with the relaxation parameter set to 0.5. Our computational domain is a rectangular box (of length L, width W , and height H) with the submerged volume of the surfer carved out of its top center [see Fig. 1(b)]. This domain is discretized using the SNAPPYHEXMESH utility in a multiblock fashion, where the mesh is densely distributed near the surfer. Grid-independence tests are performed by refining the mesh and 084004-4 repeating the simulations. In all cases considered, the computational grid is chosen such that the change in the results due to the refinement is marginal. When the simulations are intended for comparison with the experiments, the size of the domain and the depth of the particle\u2019s active region are matched with those reported in Sec", " 2, despite the uncertainty in the actual value of Re in the experiments. The data in Fig. 5 correspond to the steady-state motion of the surfers. In this condition, the net force acting on the particle in the direction of propulsion is zero, which means that the surface tension force Fst = ex \u00b7 \u222b p \u03b3 t d is balanced by the sum of pressure and viscous forces Fp + Fv = F = ex \u00b7 \u222b Sp n \u00b7 dS [see Eq. (3)]. Given the orientation of the (x, y, z) coordinate system with respect to the location of Sa p, Fst is always a positive quantity [see Fig. 1(b)]. Of fundamental interest here are the relative contributions of pressure and viscous forces. It is also of value to have an understanding about the distribution of the forces over Sp. For instance, it is informative to know that if the surfer is divided into two equal halves, say active (which encompasses the release site) and inactive, how Fp, Fv , and their sum F would split between these two regions. Table I provides this information for illustrative points in Fig. 5, where the forces and their subdivisions (denoted by the superscripts a and ia) are normalized by Fst " ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001312_s00170-020-05434-3-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001312_s00170-020-05434-3-Figure5-1.png", "caption": "Fig. 5 Schematic diagram of variable width structure", "texts": [ " 4, when the overlapping rate is large, the order of deposition from one side to the other side tends to cause the upper surface of the cladding layer to be inclined, which is disadvantageous for the deposition of the next layer. The order of deposition from the middle to the outside can obtain a relatively flat cladding layer for the next layer of deposition. Therefore, the use of a two-way deposition order from the middle to the sides facilitates the formation of the variable width thin-walled members. A schematic cross-sectional view of a variable width thinwalled part is shown in Fig. 5, and w is the width of variable width thin-walled structures. The number of required clad tracks can be calculated from the maximum width wmax. n \u00bc wmax d l m \u00f01\u00de where n is the number of tracks in the layer and n \u2265 2, d is the laser beam diameter. \u2308x\u2309 represents the smallest integer not less than x. The overlapping rate at different widths can be calculated by the following equation: \u03b7 \u00bc nd\u2212w n\u22121\u00f0 \u00ded \u00f02\u00de where \u03b7 is the overlapping rate. The scanning speed v at different overlapping rates is calculated according to the number of tracks and the overlapping rate" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001226_j.cja.2020.01.001-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001226_j.cja.2020.01.001-Figure1-1.png", "caption": "Fig. 1 Three-dimensional interception geometry.", "texts": [ " Thirdly, finite-time robust 3D composite guidance laws based on fixed-time convergence disturbance observers are presented, where the target maneuver can be fast estimated and compensated. Moreover, the solutions to deal with the peaking effect of the fixed-time convergent observers are proposed and tested. The interceptor missile and the target are regarded as point masses, and the velocities of the missile VM and the target VT are assumed to be constants. Then, the 3D interception geometry can be described as in Fig. 1. Oxyz is the inertial reference frame whose origin is moved to the missile gravity center; r is the relative range from missile to target; hL and wL are the elevation and azimuth angles of the LOS respectively; hM and wM are the flight-path angle and heading angle of the interceptor missile respectively; the same definition is for hT and wT of the target; all of the angles are Euler angles with respect to the inertial reference frame; er, eh, ew are the unit vectors along the corresponding rectangular LOS coordinate axes" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001403_j.ymssp.2020.107075-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001403_j.ymssp.2020.107075-Figure3-1.png", "caption": "Fig. 3. Interaction between the ring and the ball of rolling ball bearing with double half-inner ring.", "texts": [ " The transfer from the inertial frame to the ball azimuth frame can be expressed as rabc \u00bc Tiaribc \u00f02\u00de Where Tia is the transfer matrix from the inertial frame to the ball azimuth frame, which can be calculated as Tia \u00bc Tia g n k\u00f0 \u00de \u00bc cosncosk cosgsink\u00fe singsinncosk singsink cosgsinncosk cosncosk cosgcosk singsinnsink singcosk\u00fe cosgsinnsink sinn singcosn cosgcosn 2 666664 3 777775 \u00f03\u00de The load vector between the ball and the inner ring is collinear with rabc , so the contact angle between the inner ring and the ball is expressed as ai \u00bc arctan rabc1 rabc3 \u00f04\u00de where rabc1 and rabc3 are components of rabc on the xa and za, respectively. As for the rolling ball bearing with double half-inner rings, the relationship between the ball center and the curvature center of double half-inner rings should be considered detailedly because of the special structure. As shown in Fig. 3, the relationship between the raceway curvature center and the ball center can be given as rrbcr \u00bc rrbcl \u00fe gi;0;0\u00f0 \u00de \u00f05\u00de The contact load calculation between the ball and the inner-right raceway is taken as an example to discuss the modeling process. The vector from the ball center to the curvature center of the inner-right raceway in the contact frame can be expressed as rcbcr \u00bc rcrcr rcrb \u00f06\u00de where rcrb is the vector from the origin of the ring fixed frame to the ball center. rcrcr is the vector from the origin of the ring fixed frame to the curvature center of the inner-right raceway" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000286_b978-0-08-102884-1.00001-7-Figure1.10-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000286_b978-0-08-102884-1.00001-7-Figure1.10-1.png", "caption": "FIG. 1.10 The Whole Arm Manipulator\u2019s anthropomorphic hand. The manipulator was designed to access a glovebox through the same rubber glove as used by human operators.", "texts": [ " The development of dextrous hands or end effectors has been of considerable importance to the academic robotic research community for many years, and while the following examples are in no way exhaustive they do however present some of the thinking that has gone into dextrous robotic systems. A significant robotic end effector was the University of Southampton\u2019s Whole Arm Manipulator (Crowder, 1991). This manipulator was developed at for insertion into a human sized rubber glove, for use in a conventional glove box. Due to this design requirement, the manipulator has an anthropomorphic end effector with four adaptive fingers and a prehensile thumb, Fig. 1.10. Due to size constraints the degrees of freedom within the hand were limited to three. The Stanford/JPL hand (sometimes termed the Salisbury hand) was designed as a research tool in the control and design of articulated hands. In order to minimise the weight and volume of the hand the motors are located on the forearm of the serving manipulator and use Teflon-coated cables in flexible sleeves to transmit forces to the finger joints. To reduce coupling and to make the finger systems modular, the designers used four cables for each three degree of freedom finger making each finger identical and independently controllable (Salisbury, 1985)", " It is normal practice for the precision and power grasp not to occur at the same time. In the design of robotic dexterous end effectors, the main limitation is the actuation technology: it is recognised that an under-actuated approach may be required, where the number of actuators used is less than the actual number of degrees of freedom in the hand. Under-actuation is achieved by linking one or more finger segments or fingers together: this approach was used in Southampton\u2019s Whole Arm Manipulator, Fig. 1.10. As discussed by Birglen et al. (2008) approximately 50% of all robotic hands are of underactuated design, in an attempt to replicate the complexities of the human hand with its 20 degrees of freedom. The location and method of transmission of power is crucial to the successful operation of any end effector, in particular the end effector size should be compact and consistent with the size of the manipulator. Both fully and under-actuated dexterous artificial hands have been developed using electric, pneumatic or hydraulic actuators" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002507_j.measurement.2021.109189-Figure15-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002507_j.measurement.2021.109189-Figure15-1.png", "caption": "Fig. 15. The components of measurement system for calibration.", "texts": [ " A reflective standart material was placed under the prototype. Therefore the LED light of the camera was reflected through the prototype and robot arm connection hole whose axis pass through spherical center and polar of the spherical surface. Thus the location of the prototype was changed relative to the camera until the reflection was completed as shown in Fig. 14. The last calibration was performed by assembling the probe, which was printed in a 3D printer, into the tripod downward as shown in Fig. 15. After the probe was fixed by panning lock nut, it could be regarded as perpendicular to the floor. The probe had a hole for the camera to perform a shrink-fit assembly. Thereby the camera was restricted to rotate about X or Y axes to prevent high skew coefficient in processed image [67]. Finally, the floor and the probe of the set-up was controlled by using a standart watergage. The coordinates of the coupler point can be recorded by the help of image processing technology by using open access codes [68]" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002696_j.jmapro.2021.05.069-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002696_j.jmapro.2021.05.069-Figure3-1.png", "caption": "Fig. 3. Tensile specimens of the alloy deposits.", "texts": [ " The specimen was cut along the cross section using a wire-cut electrical discharge machine (Fig. 2e). The polished specimens were etched with Keller solution (2 mL HF, 3 mL HCl, 5 mL HNO3, and 190 mL H2O). The microstructure of the specimen was observed using an optical microscope (OM), scanning electron microscopy (SEM) in conjunction with an energy dispersive spectrometer (EDS) detector, and electron backscatter diffraction (EBSD). Transverse and longitudinal tensile tests were performed, as shown in Fig. 3. The hardness of the alloys was measured through micro Vickers, and the microhardness of the alloy deposits was measured under a 100 g load and 10 s load time. Fig. 4 shows the resultant forming quality and surface morphology when La2O3 powders of different particle sizes were added between the layers. Fig. 4a shows Specimen A, which represents the sample without La2O3 powder. The overall forming quality was good, and the roughness of the side where alloy deposition took place was less. Specimen A had the largest specimen height and the narrowest width (Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002700_s12541-021-00510-4-Figure20-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002700_s12541-021-00510-4-Figure20-1.png", "caption": "Fig. 20 Movement of the meshing point", "texts": [ " The final rotation angle was C , and whether the distance between G1 and Gg was greater than zero was judged at this moment. (22)G1 = M2 ( z1 1 z2 ) G. (23)Mp = min(dis tan ce((G1,M)) After completing the design of the gear pair, its basic theoretical characteristics, such as pump displacement, flow ripple coefficient, trapped oil volume, were determined. Unlike the straight meshing line of an involute internal gear pump, the meshing line of the conjugated straight-line internal gear pump was a curve (Fig.\u00a020); hence, the change between the meshing point angle and the gear rotation angle 1 was nonlinear. The relationship between and 1 could not be expressed by an explicit analytic equation, and the flow characteristics of the pump must be solved by a discrete numerical method. The swept area method refers to the same area swept by any two arcs rotating at the same angle on the same circle. In other words, in Fig.\u00a021, the arcs OA and OB become OA\u2019 and OB\u2019 after rotating the same angle; thereby, Now, applying this theory, the pressure oil volume change of the gear can be expressed as (24)SAOA = SBOB (25)dV1 = B r2 a1 \u2212 r2 n1 2 d 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001677_j.actamat.2020.116558-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001677_j.actamat.2020.116558-Figure1-1.png", "caption": "Fig. 1. Schematic illustrations of weld pool diagrams to show the relationship between associated angle and n on the (a) (001), (b) (011), and (c) (111) planes, and the angles \u03b1, \u03b2 , and \u03b3 between [hkl] preferred dendrite growth direction and the x-, y-, and z-axis.", "texts": [ " Therefore, the ffect of substrate orientations of different crystallographic planes an be determined by cos \u03c8 . R ( 111 ) = \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 \u221a 6 3 cos ( 60 \u25e6 \u2212 \u03be( 111 ) ) \u221a 6 3 cos ( \u2212 \u221a 6 3 sin ( 60 \u25e6 \u2212 \u03be( 111 ) ) \u221a 6 3 sin ( \u221a 3 3 hkl How to consider the effect of cos \u03c8 hkl on the different plane? ccording to the geometrical model developed by Rappaz et al. 21 , 22 ], cos \u03c8 hkl is determined by n \u00b7 uhkl . n is the normal vector o the solidification front in the melt pool and uhkl is the unit ector along the preferred dendrite growth direction [hkl]. For diferent planes, uhkl varies. Fig. 1 presents the schematic diagrams f a weld pool to show the situations for the three conventional lanes (001), (011), and (111). After this relationship has been contructed, a matrix for transforming uhkl along [hkl] crystallographic eference system into an x-y-z reference system can be determined ased on the preset initial orientation (001) = [ cos \u03be( 001 ) \u2212 sin \u03be( 001 ) 0 sin \u03be( 001 ) cos \u03be( 001 ) 0 0 0 1 ] (3) (011) = \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 cos \u03be( 011 ) \u2212 \u221a 2 2 sin \u03be( 011 ) \u221a 2 2 sin \u03be( 011 ) sin \u03be( 011 ) \u221a 2 2 cos \u03be( 011 ) \u2212 \u221a 2 2 cos \u03be( 011 ) 0 \u221a 2 2 \u221a 2 2 \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 (4) + \u03be( 111 ) ) \u2212 \u221a 6 3 cos \u03be( 111 ) + \u03be( 111 ) ) \u2212 \u221a 6 3 sin \u03be( 111 ) \u221a 3 3 \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 (5) ere \u03be (uvw) (uvw = 001, 011, 111) is the angle between the laser canning direction and the initial direction on the (uvw) crystalloraphic plane" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000798_j.neucom.2019.04.056-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000798_j.neucom.2019.04.056-Figure4-1.png", "caption": "Fig. 4. Transverse swing p ( z, t ) of the suspension cable without control.", "texts": [ " l The initial conditions are set as s (z, 0) = 0 , s \u2032 (0 , t) = 0 , \u02c6 D\u0304 (0) = 10 , and \u02c6 m (0) = 1 . 0 . All the design parameters are \u03b21 = 1 , \u03b22 = 1 , \u03bb1 = 1 . 0 \u00d7 10 \u22123 , \u03bb2 = 0 . 5 \u00d7 10 \u22123 , \u03b3 = 0 . 01 , k 1 = 1 . 5 \u00d7 10 6 , and k 2 = 1 . 0 \u00d7 10 3 . Figs. 4 \u20138 are the responses of the helicopter suspension cable system which is described by (10) under boundary conditions (16) and (17) with cable swing constraint and unknown dead-zone. The transverse oscillation amplitude without control and with the proposed control scheme are illustrated by Fig. 4 and Fig. 5 , respectively. The slop of the suspension cable with the proposed control scheme is illustrated by Fig. 6 . It shows the closed-loop system can achieve a good performance by using the proposed unilateral boundary adaptive control scheme even considering cable swing constraint and unknown dead-zone. Fig. 7 illustrates that cable swing constraint can not be violated under the designed unilat- 0 50 100 150 \u22120.03 \u22120.02 \u22120.01 0 0.01 Time (s) s (L ,t ) Fig. 7. The boundary slope s \u2032 ( L, t ) with proposed control scheme" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0003084_adem.202100611-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0003084_adem.202100611-Figure7-1.png", "caption": "Figure 7. Flexural damage results: a) the 0 sample, b) the 45 sample, c) the 90 sample, and d) fracture surfaces.", "texts": [ " The 0 sample has the lowest flexural strength, 948MPa. It is 56 and 78MPa lower than that of the 45 and the 90 samples, respectively. The 90 sample has the highest ultimate displacement, and it is about two times larger than that of the 0 sample. The bend angles of the 0 , the 45 , and the 90 samples are calculated according to the standard. The 90 sample has a 42 of \u03b1/2, and shows a great ductility than the other two build direction samples. Damage results of the three-point bending samples with three different build directions are shown in Figure 7. In Figure 7a, the 0 samples fractured incompletely with long cracks located Figure 6. The sketch of three-point bending test. Table 4. Flexural test parameters of the 0 , the 45 , and the 90 samples. Build direction L [mm] D [mm] d [mm] c [mm] f [mm] \u03b1 [ ] Ultimate flexural load [kN] Flexural strength [MPa] 0 100 30 15 32.5 16.7 42 12.6 948 90 32.9 82 13.7 1026 Figure 5. Load\u2013displacement curves of three-point bending tests. Adv. Eng. Mater. 2021, 2100611 2100611 (4 of 12) \u00a9 2021 Wiley-VCH GmbH in the middle position of specimens. For the 45 samples, some samples have the same fracture pattern as the 0 sample, and some others fractured completely, shown in Figure 7b. All the 90 samples fractured completely shown in Figure 7c. Cracks began at the center of the lower surface, where the shear lips can be observed on both sides. Samples contacted with the indenter is subjected to compressive load, which lead to the collapse of the upper surface. The trapezoidal fracture surfaces can be observed both on 45 samples and 90 samples shown in Figure 7d. Scanning electron microscope (SEM) was used to study fracture surfaces of three-point bending samples and the results are shown in Figure 8. Figure 8a1\u2013a3 shows the magnification of the crack initiation region, stable growth region and fracture region of the 45 sample in Figure 8a, respectively. And Figure 8b1\u2013b3 shows that of the 90 sample in Figure 8b. For the 45 sample, the fracture surface has a 45 angle with the deposited layers, which cause the hill feature and lager cracks shown in Figure 8a2" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000399_j.ejcon.2016.04.001-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000399_j.ejcon.2016.04.001-Figure1-1.png", "caption": "Fig. 1. Spacecraft in formation. r1, r2 and r3 are the position vectors to the centre of mass.", "texts": [ " Let the attitude of spacecraft i be given by the rotation matrix Ri representing the transformation from body fixed frame of spacecraft i to a common inertial frame. Attitude dynamics of the i-th spacecraft, for i\u00bc 1;2;3 are given by _Ri \u00bc Ri b\u03a9 i \u00f06\u00de Ji _\u03a9 i \u00bc Ji\u03a9i \u03a9i\u00fe\u03c4i \u00f07\u00de where JiAR3 3 is the moment of inertia, \u03a9iAR3 is the angular velocity, and \u03c4iAR3 is the control torque of i-th spacecraft. \u03a9i and \u03c4i are represented in the i-th spacecraft's body fixed frame. 2.3. Formation specifications Consider a spacecraft as illustrated in Fig. 1. \u00f0x; y; z\u00de is the Earth Centred Inertial reference frame. Axes \u00f0X0;Y 0; Z0\u00de, \u00f0X\u2033;Y \u2033; Z\u2033\u00de and \u00f0X000;Y 000; Z000\u00de are the body fixed frames of spacecraft one, two and three, respectively. Here spacecraft 1 is the leader while spacecraft 2 and 3 are the followers. It is required that the leader spacecraft should track a desired absolute attitude and position trajectory. Let rd1\u00f0t\u00de and vd1\u00f0t\u00de be the desired position and velocity of the leader spacecraft ecraft attitude and position tracking control, European Journal of satisfying, _rd1 \u00bc vd1: \u00f08\u00de Let the desired time varying attitude of the leader be given by Rd 1\u00f0t\u00de" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002414_012043-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002414_012043-Figure4-1.png", "caption": "Figure 4. Some natural frequencies and oscillations of the rotor system: first, second and third bending and higher vane shapes (4-6).", "texts": [ " For von Mises stresses distribution, a characteristic feature is that they reach maximums at the edge of the blade disc. Given that the mechanical properties of aluminum alloys are relatively low, this creates some problems with the impeller strength. HERVICON+PUMPS 2020 Journal of Physics: Conference Series 1741 (2021) 012043 IOP Publishing doi:10.1088/1742-6596/1741/1/012043 The next stage was the harmonic analysis of this rotor system. The previously obtained pre-stress data were used to account for the specific effect of stiffening. Figure 4 shows the selected eigenfrequencies and eigenmodes of the rotor system. It can be seen that the first two self-oscillations mainly correspond to the behavior of the hard disk and the hard shaft on the elastic supports. However, there is some contribution from the deformation of the disk and the shaft, which must be taken into account. It is especially indicated when varying the parameters of the rotor system. If we consider the higher oscillations of the rotor system, they correspond to the excitation of the blades" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002026_j.isatra.2020.06.015-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002026_j.isatra.2020.06.015-Figure1-1.png", "caption": "Fig. 1. The assembly diagram of the multiple manipulator.", "texts": [ " Introduction With the continuous expansion of modern manufacturing scales, the multiple manipulator system plays an important role in many high-precision and complex tasks. The multiple manipulator system is a complicated system with multiple inputs and multiple outputs, which has strong coupling, nonlinear characteristics. Compared with a single manipulator, the multiple manipulator system provides greater load capacity and wider operating workspace. The control algorithm of multiple manipulators is the key to accomplishing such objectives. The assembly diagram of multiple manipulator is shown in Fig. 1. Recently, the synchronized control problem of the multiple manipulator system has received significant attentions. The graph theory is used to describe the communication network topology between manipulators, and the synchronous control of the multiple manipulator system is enabled by the leader\u2013follower control mode. According to the graph theory, a new synchronization error is defined to design the synchronized controller. In [1], the distributed output feedback dynamic controller is designed according to the concept of nonlinear H\u221e control, which solves the problem of synchronous control of multiple manipulators under the condition of model uncertainty" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002316_rcar49640.2020.9303297-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002316_rcar49640.2020.9303297-Figure1-1.png", "caption": "Fig. 1. The concept of shell gripper in the circular (a) and the parallel (b) configuration, respectively.", "texts": [ " This method simplifies the fabrication process and allows us to reduce the finger thickness. The gripper concept and working principle are introduced in Section II, followed by the explanation of the fabrication method in Section III. Experimental tests on packaging cucumbers were presented in IV and Section V concludes the paper and suggests some future work. To grasp an object firmly and handle it at a high speed, we proposed the concept of shell gripper, which can be arranged in a circular or a parallel configuration, as shown in Fig. 1, respectively. As the name suggested, shell gripper is constructed by rigid shells and soft chambers. The soft 978-1-7281-7293-4/20/$31.00 \u00a9 2020 IEEE 188 Authorized licensed use limited to: Central Michigan University. Downloaded on May 14,2021 at 03:22:08 UTC from IEEE Xplore. Restrictions apply. chamber was bonded to the rigid shell. Once the chamber was inflated, the gripper can grasp an object. Due to the rigid shell, the gripper can generate larger force than soft gripper made of entire soft material" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0003110_520237-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0003110_520237-Figure2-1.png", "caption": "Fig. 2 - Point of load application under assumed worst conditions of Fig. 23 - Method of determining stress distribution along root of tooth loading near heel and toe ends", "texts": [ " (b) The teeth, which theoretically would have line contact if there were no elastic deformation and no mismatch, will have an area of contact in the form of an elongated ellipse. 3. The maximum stress is considered to occur when the tooth is carrying its maximum load, and in a position when that load is at its highest point on the tooth. 4. The stresses will not exceed the elastic limit of the material. The purpose of the above assumptions is to make the theory and equations simple enough for practical use. However, in several cases certain modifying factors must be applied, as discussed later. 315 Fig. 2 illustrates, by the broken lines, the zone of action of a bevel-gear pair. The rectangle superposed on this diagram shows the zone of action for a pair of cylindrical gears with the same face width and the same mean line of action. In the following discussion of bevel gears, it will be assumed that this rectangle may be substituted for the annular area. Usually, the theoretical contact between a pair of gear teeth is a line. However, when the tooth surfaces are mismatched, the line becomes a point" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002178_icra40945.2020.9196731-Figure8-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002178_icra40945.2020.9196731-Figure8-1.png", "caption": "Fig. 8: CAD of SoRX and materials.", "texts": [ " Results indicate that the actuator can apply 10.67 N at 20 kPa. As such, our hexapedal robot can lift a maximum weight of 3.26 kg when it follows an alternating tripod gait.4 Note that the bending part was not actuated in this test. However, as pressure increases over a critical point, the leg will passively bend; this effect can lead to the sharp increase observed in Fig. 7 at approximately 13 kPa. The new soft actuators are used to create the pneumatically-actuated soft robotic hexapod SoRX (Fig. 1, and Fig. 8). SoRX measures 230 mm L \u00d7 140 mm W \u00d7 1https://www.sofa-framework.org/ 2https://project.inria.fr/softrobot/ 3Material properties may vary due to fabrication, e.g., it is very difficult to remove all air bubbles during casting despite using a degassing chamber. 4That is, three legs are touching the ground at all times. 422 Authorized licensed use limited to: UNIVERSITY OF ROCHESTER. Downloaded on September 20,2020 at 15:48:49 UTC from IEEE Xplore. Restrictions apply. 100 mm H and weighs 650 g. The frame of SoRX was manufactured by combining laser-cut wood and acrylic sheets (Universal Laser Systems VLS 3" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000173_01691864.2019.1680316-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000173_01691864.2019.1680316-Figure3-1.png", "caption": "Figure 3. Model of the base of the robot.", "texts": [ " There are two laptops in the operator station; one is for operating the robot, and the other is for visualizing the sensor information. The gamepad, which sends commands to the robot, is connected to the laptop for operation. The two laptops are connected to a router by an ethernet cable. The robot operating system (ROS) is used as the robot\u2019s middleware, and the laptops and robot communicate on the intranet through the router. T2 Snake-4 is composed of a base and a folding arm. The base has active joints and active wheels. Figure 3 shows a model of the base. We define a module to be a group of two links and a yaw joint between the two links. The modules are serially connected by the pitch joint, and a pair of active wheels are coaxially placed with respect to the pitch joint. Figure 4 shows the details of the base. We use the Dynamixel XM540-W270R and XM430-W350-R (Robotis Co., Ltd.) as actuators for the joint and wheel, respectively. The composition of joints and wheels is almost the same as those of the ACM-R4 series [20\u201322] and T2 Snake-3 [16,19]", " If the robot uses a backward traveling wave without considering the relative relationship between itself and the stairs, it is possible for the robot to become stuck or fall down. Thus, we designed a novel stair climbing motion for the articulated mobile robot so that it can avoid becoming stuck or falling. Figure 8 shows the stairs assumed in this section. The riser heights and tread depths are all the same. We define that depth d satisfies the condition r \u2264 d < L + r, (1) where L is the module length of the robot and r is the wheel radius, as shown in Figure 3 and Table 2. The robot cannot climb stairs whose parameters satisfy condition (1) using the method of [16]. The arm is treated as a virtual module similar to the case of basic steering control considering the folding arm. We assume that the robot approaches the stairs parallel to the x axis, as shown in Figure 8. The posture shown in Figure 8 is used as the basic posture of the robot when climbing stairs. One pair of wheels contacts one tread. The part of the body connecting two treads is called the connecting part" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000471_j.ijsolstr.2016.07.004-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000471_j.ijsolstr.2016.07.004-Figure3-1.png", "caption": "Fig. 3. Wrinkled and Lengthened Surfaces ( Lu et al . , 2001 ).", "texts": [ " o ( Z s m E t 3 b t t A e a v s f t o t s o A o B i v c fi E E d m e F w s m t a E Therefore, it may be reasonable to assume that the strain f a wrinkled membrane will be comprised of two components Jarasjarungkiat et al., 2008; Raible et al., 2005; Akita et al., 2007; iegler et al., 2003 ), i.e. the elastic strain E e and the wrinkling train E w corresponding to the plastic part of a perfectly plastic aterial E p : = E e + E w (7) The physical meaning of wrinkling strain will be discussed in he following subsection for a better understanding. .3. Physical meaning of wrinkling strain Fig. 3 shows the deformation of a uniformly wrinkled memrane ABCD . The unit vector t is along the tensile direction while he unit vector w is perpendicular to the vector t . The configuraion AB\u2019C\u2019D is obtained from the stretch of the wrinkled membrane BCD to its full length without changing its stress state. It can be asily understood in the figure that the length \u03b2 l ( \u03b2 = B B \u2032 AB > 0 ) is ctually the amount of contraction of the membrane ABCD along ector w owing to the wrinkling deformation. If the wrinkling train is defined as lim A B \u2032 \u2192 0 B B \u2032 A B \u2032 = \u03b2 1+ \u03b2 for the case of the uniaxial de- ormation, it is clear that the wrinkling strain reflects the proporion of contraction of an infinitesimal membrane compared to its riginal length ( A B \u2032 )", " When a membrane is wrinkled or slack, its stress state can be etermined by: \u00af \u00b7 \u02dc \u03c3 \u00b7 t\u0304 = t \u00b7 F\u0304 \u22121 \u00b7 \u02dc \u03c3 \u00b7 F\u0304 \u2212T \u00b7 t = J \u22121 t \u00b7 \u02dc S \u00b7 t \u22121 > 0 \u21d2 t \u00b7 \u02dc S \u00b7 t or [ U 1 ] T [ \u0303 S ] = { \u02dc \u02c6 S 11 > 0 if wrinkled \u02dc \u02c6 S 11 = 0 if slack (26a) \u00af \u00b7 \u02dc \u03c3 \u00b7 w\u0304 = w \u00b7 F\u0304 \u22121 \u00b7 \u02dc \u03c3 \u00b7 F\u0304 \u2212T \u00b7 w = J \u22121 w \u00b7 \u02dc S \u00b7 w = 0 w \u00b7 \u02dc S \u00b7 w or [ U 2 ] T [ \u0303 S ] = \u02dc \u02c6 S 22 = 0 (26b) \u00af \u00b7 \u02dc \u03c3 \u00b7 t\u0304 = t\u0304 \u00b7 \u02dc \u03c3 \u00b7 w\u0304 = t \u00b7 F\u0304 \u22121 \u00b7 \u02dc \u03c3 \u00b7 F\u0304 \u2212T \u00b7 w = J \u22121 t \u00b7 \u02dc S \u00b7 w = 0 t \u00b7 \u02dc S \u00b7 w or [ U 3 ] T [ \u0303 S ] = \u02dc \u02c6 S 12 = 0 (26c) here ( \u0304t , w\u0304 ) , \u02dc \u03c3 and \u02dc S are the wrinkling coordinate system, the auchy stress tensor and the PK2 stress tensor in the fictitious onfiguration (corresponding to AB\u2019C\u2019D in Fig. 3 , in which nly the elastic deformation remains), respectively; [ \u2022 ] denotes a atrix; a variable with a caret \u02c6 \u2022 denotes one in the wrinkling oordinate system; S ij is the components of PK2 stress tensor; F\u0304 s the deformation gradient tensor used to transform the current onfiguration ( , ABCD in Fig. 3 ) into the fictitious configuration nd it can be formulated for a wrinkled membrane as: \u0304= I + b w w (27) nd \u0304= I + a t t + b w w (28) or a slack membrane; a and b are the same scalars as hose in Eq. (16) ; [ U 1 ] = [ m 2 1 m 2 2 2 m 1 m 2 ] T , [ U 2 ] = n 2 1 n 2 2 2 n 1 n 2 ] T , [ U 3 ] = [ m 1 n 1 m 2 n 2 m 1 n 2 + m 2 n 1 ] T , \u03b1 = t \u00b7 g \u03b1 = g \u03b1 \u00b7 t , n \u03b1 = w \u00b7 g \u03b1 = g \u03b1 \u00b7 w, \u03b1 = 1 , 2 . With the fact hat only the elastic deformation exists in , the stress vector [ \u0303 S ] an be obtained from the elastic constitutive relationship as: \u0303 S ] = [ C][ \u0303 E ] = [ C][ E e ] (29) n which [ \u0303 E ] is the Green\u2019s strain vector in and it equals [ E e ] as erived from: E e ] = [ E] + \u03bc1 [ U 1 ] + \u03bc2 [ U 2 ] (30) here \u03bc1 = 1 2 ( a 2 + 2 a ) , \u03bc2 = 1 2 ( b 2 + 2 b ) " ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000557_aim.2016.7576989-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000557_aim.2016.7576989-Figure2-1.png", "caption": "Figure 2. Overview of a developing equipment", "texts": [ " Based on these results, we experimentally investigated the conveyance of the developer in the conditions of the highest conveyance. Fig. 1 shows the electrophotographic imaging system structures used for image forming of the copier or printer. The photoreceptor drum formed an electrostatic latent image by charging the charging unit. After the toner on the photoreceptor drum is electrostatically transferred to the paper, image is fixed by the heat and pressure in the fusing section is output. 978-1-5090-2065-2/16/$31.00 \u00a92016 IEEE 1539 Fig. 2 shows an overview of the developing equipment comprising a screw and a developing roller. The developing device mixes the developer comprising the toner ( 5 \u03bcm) and the carrier ( 35 \u03bcm) and charges the toner. In the commonly used developing equipment, to provide the circulation, the conveyor is arranged to face the two screws. To replenish the toner that has been consumed by the development, the toner supplied from the screw end is mixed by the screw during the conveyance of the developer. The mixed developer is pumped up by the magnetic force to the developing roller acting on the carrier comprising magnetic material The printing speed of digital printing machines has been increasing continuously" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000187_s10846-019-01113-y-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000187_s10846-019-01113-y-Figure2-1.png", "caption": "Fig. 2 Demonstration of crawling function", "texts": [ " So, in order to achieve broader workspace, two innovative functions for space robot system have been proposed and concluded by Liu et al. [5]. Noticing the symmetric structure, space-manipulator\u2019s base and end-effector can switch with each other to transport itself from one platform to another, performing the crawling function. The other one is cascade connection. In practice, two manipulators which are the Core Module Manipulator (CMM) with larger size and the Experimental Module Manipulator (EMM) with smaller size, can end-to-end connect with each other, thus presenting 14DOFs. Figure 2 and Fig. 3 demonstrate these innovative functions. In this paper, mathematic solutions are provided. The working environment of space manipulators is quite complex, involving four space stations and other surrounding obstacles, as is illustrated in Fig. 4. In most cases, manual operation of manipulators is high-demanding and inaccurate. Hence, automatic trajectory planning with collision-avoidance for space manipulators is of crucial importance. Potential collision should include the collision against the environment and against the manipulator itself [6\u20139]", " sin\u03b86 = \u2212a7 \u00b7 a2 (6) tan\u03b87 = (n7 \u00b7 a2) / (o7 \u00b7 a2) (7) Furthermore, the position of Coordinate V can be calculated as p5 = l7cos\u03b87 (o7 \u2212 n7) + l8a7 + p7 (8) So, \u03b84 is given by 2l3sin(\u03b84/2) = \u2016p5 \u2212 (l2 + l4 + l6)a2 \u2212 p2\u20162 (9) Position of Coordinate III is deduced as p3 = p5 \u2212 (l2 + l4 + l6)a2 + p2 2 + v \u2016v\u20162 l3cos(\u03b84/2) (10) with v being v = a2 \u00d7 (p5 \u2212 p2) (11) So, \u03b83 can be figured out from ( \u2212l3sin\u03b83 l3cos\u03b83 l2 1 )T = T\u22121 2 ( pT 3 1 )T (12) Finally, \u03b85 is given by A5 = T\u22121 4 T7(A6 5A 7 6) \u22121 (13) Obviously, different value of \u03b81 will cause different configurations. In this paper, selection of the most appropriate configuration is achieved during the trajectory planning process, discussed in Section 5.3.2. Generally speaking, a manipulator with fixed base cannot cover a satisfying workspace. Taking full advantage of the symmetric structure, crawling function is proposed. The manipulator can switch its base and end-effector and achieve large range transfer with the platforms outside space stations as is shown in Fig. 2. Both CMM and EMM can implement this function. With appropriate D-H coordinates established in Section 2, the same kinematic model can apply to both situations by adjusting joint angles. \u03b81,7 new = 180 \u00b7 \u2212\u03b81,7 old (14) \u03b82,3,4,5,6 new = \u03b86,5,4,3,2 old (15) Herein, \u03b8i new stands for new joint angles after the switch, and \u03b8i old for the original angles. It is worth noting that \u03b8i is just a virtual parameter used in kinematics. The change of \u03b8i does not necessarily mean that the robot actually moves" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002298_tec.2020.3045063-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002298_tec.2020.3045063-Figure5-1.png", "caption": "Fig. 5. The geometry of the mathematical model for the conjugate heat transfer problems.", "texts": [ " The research considers the following configurations of the sub-slot duct profile: constant, stepped (variable), smoothly tapering to the rotor centre up to 6% of the inlet, smoothly tapering to the rotor centre up to 17% of the inlet. The geometry of the variable profile corresponds to that of the generator in operation and has three sections along the axial rotor length. The geometric characteristics of the radial and sub-slot ducts of the turbogenerator under review are as follows: \u00bd of the rotor length 2450 mm, two-row radial duct arrangement, the number of radial ducts \u2013 76, the ratio of the sub-slot duct section to the total area of radial ducts \u2013 0.26; rotation speed \u2013 3000 rpm. The numerical model shown in Figure 5 includes the air of the sub-slot duct and radial ducts; the rotor winding and ground insulation, the rotor core. The stationary problem is considered. The main solver settings, turbulence model, and detailing mesh techniques were taken from the aerodynamic model for the full-scale experiment. On the surface of the inlet to the computational domain, the axial component of the velocity was calculated through flow rate, and the temperature was set at 40\u00b0C. The flow rate was calculated separately at the stage of the computational air dynamic simulation" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001075_ecce.2019.8912286-Figure9-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001075_ecce.2019.8912286-Figure9-1.png", "caption": "Fig. 9. (a) Model view of the magnetically active portions of the magnetic gear. Views of the Rotor 3 plastic core with PMs (b) with a penny for a size reference and (c) zoomed in on the PMs and walls. (d) The initial and (e) revised Rotor 1 plastic cores and end caps with PMs and a size reference penny.", "texts": [ " 8 illustrates the impact of axially extending the Rotor 1 and Rotor 3 PMs beyond the modulators with the modulators stack length fixed at 37.8 mm. Fig. 8 shows that extending the Rotor 1 PMs increases the torque more than extending the Rotor 3 PMs the same amount. However, the Rotor 3 PMs can be extended several millimeters axially beyond the Rotor 1 PMs without increasing the gear\u2019s overall volume because the Rotor 3 end caps must extend axially beyond the Rotor 1 end caps. For this prototype, the axial lengths of the Rotor 1 and Rotor 3 PMs were selected as 47.8 mm and 51.8 mm, respectively. Fig. 9(a) shows part of a crosssection of the magnetically active portions of the design, and Figs. 9 (b)-(d) show the initial versions of the Rotor 3 and Rotor 1 plastic cores with the PMs inserted. However, when the PMs were inserted into the Rotor 1 plastic core, the outward forces on the radially magnetized PMs caused it to bulge outward by as much as 0.5 mm in some places. Therefore, a new Rotor 1 plastic core was designed with the radially magnetized Rotor 1 PMs moved inward by 0.5 mm, which reduced the simulated Rotor 2 slip torque from 31.1 N\u2219m to 30.4 N\u2219m. Additionally, the Rotor 1 plastic core was axially shortened and the Rotor 1 end cap was redesigned to interlock with the magnets to provide additional support. The redesigned Rotor 1 is shown in Fig. 9(e). Additionally, a magnetically ideal version of the design was simulated with no holes in the modulators, no modulator bridge, 1 mm inner and outer air gaps, and 100% PM fill factors. 3D FEA predicted that this ideal design would have a Rotor 2 slip torque of 69.1 N\u2219m. Thus, the very conservative choices made to facilitate the fabrication and assembly of this prototype reduced the slip torque by about 56%. IV. IMPACT OF BACK IRONS The prototype base design was also simulated with 5 mm thick back irons on both Rotors 1 and 3", " 11 illustrates the tooling for inserting the Rotor 1 PMs. This enabled the PMs to be aligned axially beyond most of the magnetic fields and then pushed into place. Similar tooling was created for inserting the Rotor 3 PMs. Also, the radially magnetized PMs were inserted before the tangentially magnetized PMs to avoid strong forces pushing the PMs being inserted away from the plastic cores. With this strategy, all of the PMs were successfully inserted, and each of the authors still has 10 unmaimed fingers. Fig. 9 shows the resulting assemblies. The finished prototype is illustrated in Fig. 2(b). End supports were included to mount the prototype to the testbed. The overall mass of the prototype was 4.4 kg with the end supports and 3.6 kg without the end supports. Table II provides a breakdown of the measured masses of some of the different prototype components. The measured Rotor 2 slip torque was 31.2 N\u2219m, which is 2.6% higher than the simulated slip torque of 30.4 N\u2219m. This yields GTD values of 25.2 N\u2219m/kg considering only the active material, 8" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001024_s00170-019-04168-1-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001024_s00170-019-04168-1-Figure5-1.png", "caption": "Fig. 5 (a) Experimental array of lattice strut elements according to specified independent (Table 2). (b) Tessellated experimental array of lattice strut elements. (c) Schematic of circular cross section versus other polygon order sections", "texts": [ " Where the radius of a circle with a perimeter equal to the closed curve\u2019s (rp,) is: rp \u00bc p 2\u03c0 \u00f02\u00de And the radius of a circle with an equivalent area to the curve (rc) is: rc \u00bc ffiffiffiffiffiffi Ac \u03c0 r \u00f03\u00de The isoperimetric quotient can then be written in terms of the closed curve\u2019s area (Ac) to its perimeter (p) as: Q \u00bc Ac Ap \u00bc r2c r2p \u00bc Ac \u03c0 p 2\u03c0 2 \u00bc 4\u03c0Ac p2 \u00f04\u00de The effective diameter (Deff) is defined as the diameter of a circle that has equal area to that of the closed curve in question and is given as Deff \u00bc 2 ffiffiffiffiffiffi Ac \u03c0 r \u00f05\u00de The polygon order (P) describes the number of equally spaced linear segments, used to generate the cross-sectional geometry to be manufactured. Three polygon orders were analysed (P = 3,4,8) to compare with the reference circular geometry (P \u2192 \u221e), and are displayed in Fig. 3. Manipulation of the polygon arc length (s) allows the area of the polygon to be maintained to that of a reference circle with a specified nominal diameter (Dnom). As P increases, the isoperimetric quotient tends towards that of a perfect circle (i.e. as P\u2192\u221e, Q\u21921.0) as shown in Fig. 4. A physical lattice element structure was designed (Fig. 5) to enable a relevant experimental DOE to be established for independent variables: nominal strut diameter (Dnom); strut centreline inclination angle (\u03b1) and polygon order (P) (Table 2). For this initial study, inclination angles useful for the design of lattice structures were selected, specifically, \u03b1 = 90\u00b0, 45\u00b0, and 35\u00b0 are specified for vertical support, face centred and body centred structures, respectively [21, 22]. Lattice strut elements were oriented such that a symmetry plane, defined by the strut axis and side face normal, was oriented to align with the powder bed. For the case of triangular cross section (P = 3), the two orientations which meet this requirement, identified as downward (P = 3dn) and upward (P = 3up), were included (Fig. 5). A total of 120 permutations of DOE variables were fabricated with SLM using the process parameters of Table 1. Computed microtomography (\u03bcCT) was used to acquire geometrical data of as-manufactured specimens. The \u03bcCT was performed to a voxel size of 8 \u03bcm, enabling qualitative and quantitative inspection of the effect of DOE variables on lattice strut geometry. Custom scripts were developed to process \u03bcCT data into individual slices, perpendicular to the strut inclination axis, and to extract the associated isoperimetric Table 2 Design parameters" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001256_j.measurement.2020.107831-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001256_j.measurement.2020.107831-Figure1-1.png", "caption": "Fig. 1. (a) Front view of the fixed canards, (b) diagram of projectile.", "texts": [ " There are several models for trajectory simulation, including 6DOF (six-degree-of-freedom) trajectory model [20,21] with highest accuracy, point-mass trajectory model [22] with least computational cost and the modified point-mass trajectory model [23,24] with appropriate trajectory prediction accuracy, and low computational cost which is suitable for real-time application on the onboard computer. Therefore, the modified point-mass trajectory model is used to establish of the extended Kalman filter without velocity input for trajectory estimation. The model of a spinning projectile with fixed canards is established as shown in Fig. 1 [25]. As shown in Fig. 1a, there are two pairs of canards with fixed deflection angle de of the fixed canards. One pair of them is used to produce rotational moment to despin the projectile, while another pair of them is used to provide the steering forces to control the attitude of the projectile. In Fig. 1b, it shows the reference frame Ongf of the projectile, and the body frame O1nccgccfcc of the fixed canards. The reference frame Ongf of the projectile is defined to be rigidly attached to the spin axis of the projectile, not the rest of the body, where On points to the head of the projectile; Og is orthogonal to On and points upward; Of is orthogonal to On and Og with the usual right-hand rule. The definition of O1nccgccfcc is shown in Fig. 1a and Fig. 1b. The included angle between O1nccgccfcc and Ongf is the roll angle of fixed canards cc . The roll angle of the fixed canard is defined as 0 when the canards provide upward steering forces as shown in Fig. 1a and the roll angle increases as canards rotate counterclockwise. To study the motion of high-spinning projectile with fixed canards, the angle of attack is the focus that we need to pay attention to. Based on the angular motion equations of the projectile [20], the angle of attack of high-spinning projectile with fixed canards consists of three parts, the angle of attack of circular motion D1, the angle of attack produced by gravity D2, and the angle of attack produced by fixed canards D3. The gross angle of attack of the projectile Dwith fixed canards is given by D \u00bc D2 \u00fe D3 \u00bc \u00f0D2p \u00fe D3p\u00de \u00fe i\u00f0D2d \u00fe D3d\u00de \u00bc Dp \u00fe iDd \u00f01\u00de For D1 is a variation with mean value of 0, it is omitted in (1)" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001062_s00202-019-00874-x-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001062_s00202-019-00874-x-Figure3-1.png", "caption": "Fig. 3 Stator slot details for 11\u00a0kW, 400\u00a0V, eight-pole, 50\u00a0Hz motor with S = 48 stator slots: Din= 204.4, Dout= 320, l = 200.6, g = 0.4, bts= 8.5, bos= 2.5, hos= 1, hw= 2, bs1 = 5.3, bs2 = 9.0, hs= 28.5, hcs= 26.3 (all dimensions in millimetres)", "texts": [], "surrounding_texts": [ "After the initial design process for an 11\u00a0kW IM with S = 48 slots and an arbitrarily chosen R = 30 rotor bars, the model parameters were obtained, as shown in Table\u00a02. While the initial design was for 30 rotor bars, by employing the PWF model, the steady-state electromagnetic torque is acquired for any desired number of rotor bars and hence, the torque ripple factor can be calculated using (4)\u2013(7). IM operation with three different arbitrarily chosen number of rotor bars 26, 33 and 40, respectively, is examined initially with the PWF model and then with FE models to cross-verify the results. A wider study of possible rotor bar numbers was subsequently undertaken in the PWF model only due to the prohibitive execution time constraints associated with the FE model analysis. An example of the PWF and FE model outputs for electromagnetic torque for the three investigated cases is shown in Figs.\u00a05, 6 and 7. The presented transients simulate the process of IM direct start-up for no-load conditions, and once steady-state is reached at t = 0.6\u00a0s, the rated load is applied. A comparison between the corresponding PWF and FE model results reveals a good level of agreement, both quantitatively and qualitatively. This is particularly apparent for the R = 33 case, shown in Fig.\u00a06. The notable high spikes in the PWF torque during startup, for the case of non-skewed rotor bars, arise due to the stepwise shape of the mutual inductance derivative curve between stator phases and rotor loops, given later in the text (Fig.\u00a016). When the rotor bars are skewed, the inductance derivative function is smooth resulting in no spikes in the PWF model predicted torque (Fig.\u00a016) [18]. The natural consequence of rotor bar skewing is a somewhat reduced value of electromagnetic torque and hence, a longer time period needed for machine acceleration. A comparison of the PWF and FE model steady-state electromagnetic torque predictions for the three examined rotor bar number cases is shown in Figs.\u00a08, 9 and 10. Again, the obtained results are seen to be in good agreement, further confirming the PWF model predictions validity. It is interesting to note that the introduction of rotor bar skew has a significant mitigating effect on the electromagnetic torque ripple for R = 26 and R = 40. However, for the case of R = 33, rotor bar skew does not significantly modify the existing ripple level and is thus, an almost unnecessary measure in this respect. Furthermore, it can be shown that this observation is valid for almost all odd rotor bar numbers. To enable an evaluation of the rotor bar number choice for the examined IM from the perspective of torque ripple minimization, a wider study was performed with the PWF model for rotor bar numbers ranging from R = 20 up to R = 73 (approximately in the range 0.5S \u2264 R \u2264 1.5S). The obtained results for the torque ripple factor are shown in Table\u00a03 for each considered number of rotor bars. Furthermore, both skewed and non-skewed rotor designs were assessed for each evaluated rotor bar number. 1 3 Table\u00a03 also provides torque ripple factors obtained from the FE model for R = 26, R = 33 and R = 40. The corresponding FE model and PWF model data are seen to be in good agreement. Furthermore, identical trends are obvious from both models. For example, both report the smallest torque ripple factor value for the case R = 33, and the largest value 1 3 1 3 1 3 for the case R = 40. For ease of interpretation, the torque ripple factors from Table\u00a03 are also graphically presented in Figs.\u00a011 and 12." ] }, { "image_filename": "designv11_5_0000908_tpel.2019.2934984-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000908_tpel.2019.2934984-Figure1-1.png", "caption": "Fig. 1. LVPM motor. (a) Cross-section. (b) Prototype.", "texts": [ " In Section II, the mathematical model of the LVPM motor is established. In Section III, an improved mover flux observer is proposed and analyzed. In Section IV and Section V, the main simulation and experimental results are given to verify the feasibility of the proposed method. Finally, conclusions are summarized in Section VI. II. MATHEMATICAL MODEL In this work, a LVPM motor is chosen as the control object, which consists of the short mover and the long stator. The cross-section and prototype of the LVPM motor is shown in Fig. 1. For long stroke applications, both PMs and windings are installed in the short mover and only cores and salient poles are located in the long stator. This design contributes to reduce the cost, and according to the flux modulation principle [1], the motor can provide high thrust density. The motor model in the stationary reference frame is s s s s s s d u R i L dt d u R i L dt (1) Then, the mover flux can be drive as f s s f s s L i L i (2) The relation between the mover flux and electrical angular position of PMs inducing in the windings can be expressed as cos sin f PM e f PM e (3) arctan f e f (4) The electrical angular position also has the relationship with the primary position 2 e x (5) So, the electrical angular frequency can be expressed as 2 ew v (6) A main characteristic of the LVPM motor is that the motor can provide high thrust density" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002356_978-3-030-50460-1-Figure15.17-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002356_978-3-030-50460-1-Figure15.17-1.png", "caption": "Fig. 15.17 Equilibrium at unit cell and internal stress resultants", "texts": [ " The relation between stress resultants and deformations is described in the local system shown in Fig. 15.2, where the coordinate mapping between the moving trihedral x \u03be2 \u03be3 and the reference system xyz is given with\u23a7\u23a8\u23a9 x \u03be2 \u03be3 \u23ab\u23ac\u23ad = \u23a1\u23a3 1 0 0 0 cos \u03b8 sin \u03b8 0 \u2212 sin \u03b8 cos \u03b8 \u23a4\u23a6\u23a7\u23a8\u23a9 x y z \u23ab\u23ac\u23ad . (15.21) For writing convenience of subscripts, we prefer to replace the symbol \u03be2 with s. Fig. 15.2 indicates points in the reference configuration, y(0)(\u03be2) = Rsin\u0394\u03c8 ; z(0)(\u03be2) = R [1\u2212 cos\u0394\u03c8]\u2212 c , (15.22) where \u0394\u03c8 = \u03c8 \u2212 \u03c80 = \u03be2 R \u2212 \u03c80 = \u03ba0\u03be2 \u2212 \u03c80 , (15.23) Point B in the reference configuration shown in Fig. 15.17 is given with: y (0) B = P 4 ; z (0) B = \u2212c . (15.24) Cylindrical bending implies that curvature deformations vanish, \u03b51x = \u03b51xs = 0. We also assume that the midplane strain transverse to the corrugations and the shearing 15 Manufacturing and Morphing Behavior of High-Amplitude Corrugated Laminates 253 strain vanish, \u03b50x = \u03b50xs = 0. As the laminate is free of couplings between in-plane shear and normal line forces as well as between twist and bending moments, \u03b50xs = 0. The remaining relations between line loads and laminate deformations are: Nx = A12\u03b5 0 s Mx = D12\u03b5 1 s Ns = A22\u03b5 0 s + B22\u03b5 1 s Ms = B22\u03b5 0 s + D22\u03b5 1 s (15.25) Fig. 15.17 redraws the mechanical unit cell identified in Fig. 15.16, where \u03b8 coincides with \u0394\u03d5 if no deformation is applied, and indicates stress resultants reacting to the essential boundary conditions. A specified line force NB = Ny = N causes the reaction line force NA = \u2212N at point A. The internal moment at B is obtained from the cross product of the reaction NA and distance between the line along which NA acts and point B: \u23a7\u23a8\u23a9 0 0 zB \u23ab\u23ac\u23ad\u00d7 \u23a7\u23a8\u23a9 0 \u2212N 0 \u23ab\u23ac\u23ad = \u23a7\u23a8\u23a9 zBN 0 0 \u23ab\u23ac\u23ad = \u23a7\u23a8\u23a9 MB 0 0 \u23ab\u23ac\u23ad (15.26) The local mid-surface in-plane line force N2 (\u03be2) along \u03be2 is a component of N obtained from the transformation rule (15", "34) Note that the stretch of the line elements becomes significant at the end of morphing action when the corrugated laminate has become almost flat. Positions of points in the deformed configuration are found by integrating the slope (Kress and Winkler, 2009) and thereby considering the stretch (Ren and Zhu, 2016) of line elements: duy = du0 y + du1 y ; duz = du0 z + du1 z . (15.35) At position \u03be2 the stretching of the line element gives the incremental displacements: y(t) (\u03be2) = \u222b \u03be2 0 cos (\u03b8)\u03bbd\u03c3 ; z(t) (\u03be2) = \u2212 \u222b \u03be2 0 sin (\u03b8)\u03bbd\u03c3 . (15.36) At the reference configuration the point B indicated in Fig. 15.17 has the position y (0) B = P/4, z(0)B = \u2212c. Deformation changes that position to: y (t) B = \u222b Ls 0 cos (\u03b8)\u03bbd\u03c3 ; z (t) B = \u2212 \u222b Ls 0 sin (\u03b8)\u03bbd\u03c3 . (15.37) Displacements follow from the difference between current and reference configurations: uy = y(t) \u2212 y(0) ; uz = z(t) \u2212 z(0) , (15.38) which yields: uy = \u222b \u03be2 0 cos (\u03b8)\u03bbd\u03c3 \u2212 y(0) ; uz = \u2212 \u222b \u03be2 0 sin (\u03b8)\u03bbd\u03c3 \u2212 z(0) . (15.39) Deformed configurations of the problem at hand depend most directly on an applied line force N . On the other hand, it is desired to specify morphing stretch \u03bb\u0302morph" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001800_978-3-030-40410-9-Figure1.3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001800_978-3-030-40410-9-Figure1.3-1.png", "caption": "Fig. 1.3 a 5R manipulator, b RRdR manipulator (SCARA robot)", "texts": [ " For instance, there are assembly robots (Kramberger et al. 2017; Makris et al. 2017) which carry heavy parts or PCB manipulators where dynamic loads donot have to be considered\u2014of course, both of the previous examples require precise motion. There are many types of industrial robots, and each is used according to the purpose and desired duty (Spong et al. 2005). Themost common type of robots are serial robot manipulators, which are a series of rigid bodies, called links, joined together by means of joints (Ghafil et al. 2015), see Fig. 1.3. In manufacturing lines of the vehicle industry, it is economically undesirable to design all the robot manipulators according to the same criterion because the robot\u2019s joints and links are subject to different loads in different lines of production. A robot manipulator in \u00a9 Springer Nature Switzerland AG 2020 H. N. Ghafil and K. J\u00e1rmai (ed.), Optimization for Robot Modelling with MATLAB, https://doi.org/10.1007/978-3-030-40410-9_1 1 2 1 Introduction an assembly line will suffer from more stresses than one in the painting or welding lines" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000958_13621718.2019.1666222-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000958_13621718.2019.1666222-Figure2-1.png", "caption": "Figure 2. The location of the welding electrodes.", "texts": [ " For the hybrid tandemMAGwelding experiment, this auto carriage machine is re-designed. An additional electrode holder is added to control feeding angle and position of the hot-wire. Each holder has a linear and a rotational stage to adjust the contact tip to work distance (CTWD) of each filler wire. As a result, a leading electrode, a hot-wire and a trailing electrode are adjusted by the carriage. The base and web plates used in our experiments are sliced to the size of 250mm \u00d7 80mm \u00d7 25mm and 250mm \u00d7 80 mm \u00d7 15mm, respectively, as shown in Figure 2. The flux-cored wire (E71T-1C) which has a diameter of 1.4mm is used for both a leading and a trailing electrode. Both wires are set to the DC-EP. In addition, the solid wire (ER70S-3), with a diameter of 1.2mm is selected as a hot-wire material and is heated with the DC-EN, which is placed between the leading and trailing electrodes. The chemical compositions of the AH36 plate and filler wires are listed in Table 1. During the welding experiments, the CTWD is maintained at 18mm for the leading electrode and 21mm for the trailing electrode, as shown in Figure 2. The travel speed of the carriage is maintained at 1.5mmin\u22121, and the travel angle of both electrodes is assigned to \u00b110\u25e6 around the hot-wire that has a 0\u25e6 travel angle. The work angles of the three electrodes are all set to 42\u25e6 and, 100% CO2 shielding gas is used with a flow rate of 18 Lmin\u22121. The inter wire distance, which is the distance between the leading and trailing electrodes, was fixed to 25mm. In preliminary experiments, when the inter wire distance was smaller than 25mm, the weld beads were irregularly formed due to overflowing themelt pool" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000465_j.jmapro.2016.06.021-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000465_j.jmapro.2016.06.021-Figure3-1.png", "caption": "Fig. 3. Mechanical design of the laser system: (a) design dr", "texts": [ " It is difficult to use mechanical adjustment pproaches to achieve the purpose. In our study, we adopted an ptical adjustment approach by using a two-axis gyro-mirror sysem. As shown in Fig. 2(a), we can easily fine-tune the position of he scanned laser line by changing the positions of the two ends f the laser line (i.e. End1 and End2). Using the control software of he gyro-mirror system, the position of the laser scan line can be ccurately adjusted to align with the tooltip opening. .3. Immersed accumulation tool A compact structure as shown in Fig. 3(a) is designed to encapulate all the laser optics in order to move the laser scanning system n the fabrication process. The accumulation tool has opaque side- alls to prevent the laser beam from exposing to liquid resin that s not immediately below the bottom opening. In addition, the tool ip needs to be sealed by some transparent film such that no resin ill be able to get inside the tool when the accumulation tool is mmersed in liquid resin. The following two main functions need to be considered in the ool tip design: (1) providing a linear light source to the liquid resin, nd (2) reducing the separation force between the tool tip and Please cite this article in press as: Mao H, et al. LISA: Linear http://dx.doi.org/10.1016/j.jmapro.2016.06.021 he photo-cured line segments. The first main function requires transparent tooltip opening to enable laser beam to pass through he tool tip. One of such designs is shown in Fig. 3(a). The second ain function requires the tool tip to have a small separation force source for LISA. between the newly cured line segments and the film coated on the tool tip. As shown by Chen et al. [7], the coating material on the tool tip plays a significant role on reducing the separation force. It is desired to use Teflon film or some other coating materials with similar properties to ensure the newly cured line segments will detach from the tool tip and attach to the previously cured layers. Hence, a 3D object can be continuously fabricated line by line. Fig. 3(b) shows a tool tip design based on Teflon film. The tool tip design is critical in the LISA process. The following design parameters need to be considered. \u2022 Proper opening gap size d: This opening gap is designed to pass the laser beam. It can also physically regulate the laser beam size, i.e. the gap can play the role of an aperture. In our experiment, we have tried the gap size from 0.5 mm to 7 mm, and found that the 0.5 mm can largely reduce the diffraction of laser and generate a better surface finish" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001158_s00170-020-04957-z-Figure8-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001158_s00170-020-04957-z-Figure8-1.png", "caption": "Fig. 8 The layout of temperature measuring points", "texts": [], "surrounding_texts": [ "Based on the works above, the IGWO and GRNN are combined to establish the thermal error model. The specific procedures of IGWO-GRNN are given as follows. (1) Determine the input and output variables of the model. The inputs of the model are the temperature sensitive variables obtained in Section 2.2. The thermal drifts (on X-direction and Y-direction) of the hobbing machine workpiece shaft are the outputs. (2) Initialize the IGWO and the structure of GRNN. Determine the structure of GRNN (the number of neurons in every layer) according to the sample size and the dimensions of input and output. Set the population and the maximum number of iterations of the IGWO in advance. Initialize the location parameters of gray wolf individuals randomly. The positional parameters of the gray wolves represent the values of smooth factor \u03c3. The fitness function is defined as the root mean square error (RMSE) of the model outputs, which is expressed by Eq. (24). The smaller the error of the Temperature sensors Mounting positions Displacement sensors Mounting positions predicted results, the better is the fitness of the gray wolf individual. F \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 m \u2211 m i\u00bc1 yi\u2212yi0 2s \u00f024\u00de where m is the dimension of the output vector, yi is the predicted value of the ith dimension of the output, and yi0 is the corresponding true value. The fitness values of \u03b1, \u03b2, and \u03b4 wolves are set to infinity. (3) Train the model. Update the smoothing factor \u03c3 of GRNN and calculate the fitness of every gray wolf. Compare the fitness of all gray wolves to update the best three individuals. The positional parameters of other wolves are updated under the guidance of the best three individuals. Then a new population of gray wolves is generated. Record the parameters of the best three individuals and start to do the next iteration. Repeat these processes until the terminal conditions are met. If any of the following conditions are satisfied, the iteration of model will stop. (1) The number of iterations reaches the given maximum. (2) The fitness of gray wolf individual is lower than a given value. The setting of the terminal conditions of the algorithm guarantees the generalization performance of the algorithm. (4) Prediction of thermal error for gear hobbing machine using the IGWO-GRNN. Obtain the optimal prediction model (optimal smoothing factor) to predict the thermal drift of workpiece axis. The execution flow of the IGWO-GRNN is shown in Fig. 6 and described as follows." ] }, { "image_filename": "designv11_5_0000200_s00202-019-00882-x-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000200_s00202-019-00882-x-Figure5-1.png", "caption": "Fig. 5 Regions of a coaxial magnetic gear", "texts": [ " The material used for the steel segments and rotor yokes is a low-carbon steel AISI 1008 [21]. The stack length of the constructions is 300\u00a0mm [18]. In order to analyze magnetic field distributions, behavior and characteristics of magnetic gear, we developed an approach for magnetic field modeling at different modes and conditions [17\u201319]. The approach employs the transient finite element method (FEM) and ANSYS Maxwell software to analyze the steady-state and transient performances of the CMG [22\u201326]. The regions of interest in the magnetic gear construction are presented in Fig.\u00a05. They are as follows: permanent magnet regions of the two rotors, ferromagnetic parts of magnetic circuit, modulating steel segments and air gaps. Modeling domains \u03a91, \u03a92, \u03a93, \u03a94, \u03a95, \u03a96 and \u03a97 correspond to: the inner rotor yoke, the inner rotor permanent magnets, the air gap between the inner rotor and the segments, the modulating segments ring, the air gap between the steel segments and the outer rotor, the outer rotor permanent magnets and the inner rotor yoke, respectively. Time-dependent magnetic vector potential\u2013scalar electric potential formulation is used [25]" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002702_iciccs51141.2021.9432177-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002702_iciccs51141.2021.9432177-Figure2-1.png", "caption": "Fig. 2. Parts of a bearing.", "texts": [ " The fault unless detected at an early level, it can lead to catastrophic results in terms of unfavorable maintenance time and cost, decrease in production level, and severe risk of loss of human lives [2]. A bearing's primary function in rotating machinery is to sustain the load and minimize friction. The outer race, inner race, cage, and balls are the main components of the bearing. The cage keeps the balls in place within the inner and outer races, allowing for free and smooth rotation [3]. A real image of a bearing is shown in Fig 1. Different parts of the bearing are shown in Fig 2. Concerning the possibility of induction machine defects, bearing faults are the most common fault type, accounting for more than 30 percent of all failures. In addition, this fault is the common reason for the motor shut down, causing in a major loss of property and safety [4, 5]. Pertaining to the previous reasons, bearing fault diagnosis is a crucial part of development as well as engineering research. Fault detection and condition monitoring of the machine provide significant information about its status during the operation at each moment, keeping the working condition" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002037_s00158-020-02686-1-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002037_s00158-020-02686-1-Figure3-1.png", "caption": "Fig. 3 Experimental rig (left), experimental rig schema (top right), and a pair of etalons (bottom right)", "texts": [ " 2017, 2018), the authors learned that the computational time is similar to the time required to input the requirements needed for the calculation of boundary condition values. Thus, it was concluded that NSGA-II is suitable for this kind of problem and number of variables; it is wellestablished by scholars, is reliable, and has a wide range of applicability. The experiment was necessary to validate the analytical results. The open-circuit experimental rig was designed in order to measure the frictional power losses in gear pairs (see Fig. 3). The rig consists of two AC motors, first having the nominal power of 4 kW while operating at 2800 min\u22121 and second of 12 kWwhile operating at 3000min\u22121. The former is providing the operational torque, working as driving machine, while the latter is providing the reactive torque (driven machine). Two torque transducers were mounted between the shafts and motors to determine the gear pair power losses. The 20-Nm transducer was mounted onto the input shaft, while the larger, 50-Nm transducer is mounted onto the output shaft", " The overall power loss between the transducers is calculated as: Ploss\u03a3 \u00bc 2\u03c0 nt1 T tt1\u2212nt2 T tt2\u00f0 \u00de \u00f012\u00de where Ttt1 [Nm] is torque measured by 20-Nm transducer (input), Ttt2 [Nm] torque measured by 50-Nm transducer (output), nt1 [s \u22121] the number of rotations per secondmeasured at the input shaft, and nt2 [s \u22121] number of rotations per second at the output shaft. The rig was designed to examine gear pairs with various centre distances since it was necessary to ensure that pairs found in the optimisation process could be examined. Variations in centre distance were achieved by moving the bearing housings. Since centre distance tolerances in gear reducers are narrow, thus requiring the high precision machining, etalons for their adjustment (see Fig. 3) were designed and manufactured to ensure the necessary shaft positions. These etalons ensured the correct relative positions of axes, providing both the dimensional and geometric accuracy. The manufacturing tolerance of etalon bore spacing was + 0.02 mm, with bore tolerances of H6. Torques required to rotate the bearings must be determined before comparing the measured and calculated power loss values. The manufacturer provided a model for calculation of torques (Linke et al. 2016): PlossB \u00bc 2\u03c0n1 60 M rr\u2212shaft1 \u00feM sl\u2212shaft1\u00f0 \u00de \u00fe 2\u03c0n2 60 M rr\u2212shaft2 \u00feM sl\u2212shaft2\u00f0 \u00de \u00f013\u00de Two pairs of self-aligning ball bearings 1205-K were used (bearing series 12)" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001746_j.mechmachtheory.2019.103779-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001746_j.mechmachtheory.2019.103779-Figure6-1.png", "caption": "Fig. 6. Schematic of the bevel gear system equipped with FLSFD.", "texts": [ " 2 , k and k e are connected in series, and k is one order of b b b magnitude larger than k e , which implies that the equivalent support stiffness is mainly dominated by k e . Therefore, for the sake of simplification, the radial support stiffness k b is ignored in this paper. Besides, the oil-film is represented by the fluid reaction forces in the x- and y-directions ( F lx and F ly ), where l = p, g refers to the pinion and gear, respectively. c a denotes the damping coefficient of gears in the lateral direction. The simplified dynamic model of BGS with FLSFDs is shown in Fig. 6 . Only three translational displacements and one rotational displacement around the Z-axis are incorporated, the general- ized coordinates of the system are, { q } = { x p , y p , z p , \u03b8p , x g , y g , z g , \u03b8g } (16) According to D\u2019Alembert principle, the motion equations of the gear-SFD system can be written as: m p \u0308x p + c e \u0307 xp + k e x p + n px F m \u2212 F px \u2212 m p e p \u03c9 2 p cos ( \u03c9 p t ) = 0 (17) m p \u0308y p + c e \u0307 yp + k e y p + n py F m \u2212 F py \u2212 m p g \u2212 m p e p \u03c9 2 p sin ( \u03c9 p t ) = 0 (18) m p \u0308z p + c a \u0307 zp + k a z p + n pz F m = 0 (19) m g \u0308x g + c e \u0307 xg + k e x g + n gx F m \u2212 F gx \u2212 m g e g \u03c9 2 g cos ( \u03c9 g t ) = 0 (20) m g \u0308y g + c e \u0307 yg + k e y g + n gy F m \u2212 F gy + m g g \u2212 m g e g \u03c9 2 g sin ( \u03c9 g t ) = 0 (21) m g \u0308z g + c a \u0307 zg + k a z g + n gz F m = 0 (22) I p \u0308\u03b8p + \u03bbpz F m \u2212 T E = 0 (23) I g \u0308\u03b8g + \u03bbgz F m \u2212 T L = 0 (24) where m l is the mass of the gear l ( l = p, g ) , and \u03c9 l denotes rotation speed without considering the torsional vibration" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001650_j.measurement.2020.108950-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001650_j.measurement.2020.108950-Figure1-1.png", "caption": "Fig. 1. A CAD model of the test rig: (a) isometric view, (b) fixture apparatus.", "texts": [ " A special test rig was designed as a part of the study, and the effect of tooth profile on impact strength was investigated. The purpose of this study is to investigate the effect of drive side pressure angle on the impact strength of the involute spur gears. Therefore, a special test setup was designed, and a novel experimental procedure was proposed to investigate the effect of the tooth profile on impact strength. The 3D CAD model of the test rig was designed using SolidWorks\u00ae (Waltham, MA, USA) software. A detailed illustration of the test setup is presented in Fig. 1a. A novel test setup was designed, which measured the impact strength via releasing a load from a given height, allowing it to free fall. The basic principles underpinning how the drop-weight test system works are inspired by designs in previous studies. The test rig consists of a rigid body that is mounted to a base using steel bolts. The other components of the gear impact test rig are: a pneumatic load releasing system, a loading mechanism, a load cell attached force impactor, and the fixture apparatus. This study designed and manufactured a unique fixture apparatus compatible with the asymmetric tooth profile. Bolt connections added to the fixture apparatus allow the movement of the test gear within the apparatus boundaries, as presented in Fig. 1b. Thus, the impact force can be applied to the desired diameter of the test tooth. During the tests, the fixture apparatus\u2019s stability was guaranteed with a fixed rack gear which is in contact with the gear test sample. Cast nylon sheets were placed on both sides of the gear sample for fixation. Besides, it was ensured that the rotation of the test sample does not occur because a full meshing was provided (Fig. 1b). The pressure angle in the direction where torque transfer is provided for involute gears is called the drive side pressure angle (\u03b1d), while the other side is called as coast side (\u03b1c) pressure angle. Fig. 2 shows the identifying geometric characteristics of asymmetric involute spur gears. In a comparative experimental study, different drive side pressure angles can be used to examine the effect of asymmetric tooth profile on gear impact strength. However, it was observed that among them, the drive side pressure angles varying between 25\u25e6 to 35\u25e6 are frequently preferred in the literature [4\u20138,14,21,34]" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002870_s00500-021-06159-5-Figure12-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002870_s00500-021-06159-5-Figure12-1.png", "caption": "Fig. 12 Section view and design variables of rubber bushing", "texts": [ " P h x\u00f0 \u00de; g x\u00f0 \u00de; r\u00f0 \u00de \u00bc r Xp i\u00bc1 hi x\u00f0 \u00de\u00bd 2\u00fe Xm i\u00bc1 g\u00fei x\u00f0 \u00de 2( ) ; g\u00fei x\u00f0 \u00de \u00bc max 0; gi x\u00f0 \u00de\u00f0 \u00de \u00f09\u00de In addition, the weight factors have been added to constraint functions hi x\u00f0 \u00de by using the scalar weight function method to balance between cost and constraint functions. h x\u00f0 \u00de \u00bc Xm i\u00bc1 wihi x\u00f0 \u00de; Xm i\u00bc1 wi \u00bc 1 \u00f010\u00de where wi is the weight factor of ith constraint function (hi). Finally, the unconstrained target function is as follows. U \u00bc kTr kir kTr \u00fe 0:3 0:33 kTa kia kTa 2 \u00fe0:33 kTt kit kTt 2 \u00fe0:34 kTc kic kTc 2 ! \u00f011\u00de where r is the vector of penalty parameters has been taken as 0.3 concerning the experience. The section of rubber busing is shown in Fig. 12. The reference dimensions on the section view have been defined as design variables, as shown in Table 13. Shore hardness and swaging are not dimensions. Shore hardness is the durometer level of the rubber. Swaging is a process parameter, and it reduces the outer diameter of the bushing after the vulcanization process. The geometry in Fig. 12 is created parametrically in the Abaqus CAE software. 3D first-order eight-node hexagonal brick element (C3D8) and six-node five-sided triangular prism elements (C3D6) have been used in the finite element model to be used in the solution. The hyperelastic material model is used to simulate rubber in finite element analysis. In addition, hybrid elements have been selected since it is mandatory to use hybrid elements with the hyperelastic material model in Abaqus. Hyperelastic materials in finite element analysis can be modeled using different material models such as Mooney\u2013 Rivlin, Ogden, Yeoh and neo-Hookean (Gu\u0308ven et al" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001788_j.ijmecsci.2020.105516-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001788_j.ijmecsci.2020.105516-Figure3-1.png", "caption": "Fig. 3. (colour online) Investigated bearing configurations. Diameter of a feed hole depicted in part (a) is valid for all configurations. Section A-A shows dimensions of the oil supply bore and section B-B depicts dimensions of the cirfumferential groove.", "texts": [ " Thus, the elastic deformations of the shaft are considered in the odel of the bearing gap. The floating rings are considered rigid and hey have only three degrees of freedom \u2014 two lateral motions and otation. It is a reasonable approximation because the angular misalignent of the floating ring is roughly 10 \u22123 rad at maximum. .1. Geometry of the full-floating ring bearings Basic specifications for FFRBs and the floating rings are given in able 1 c. Six layouts of the axial and circumferential grooves are inroduced in Fig. 3 and they are described below: a) There are no grooves in the bearing. b) There is a circumferential groove in the outer surface of the ring. c) There is a circumferential groove in the inner surface of the bearing housing. d) There is a partial circumferential groove in the inner surface of the bearing housing. e) There is a partial circumferential groove in the inner surface of the bearing housing. Furthermore, there are four equally spaced axial t p c i i 3 s grooves in the inner surface of the ring, which are located between the feeding holes", " Thermal effects including temperature-dependent earing clearances, temperature-dependent viscosity and heat transfer n both turbine and compressor stages are considered. Figs. A.13 c and Fig. A1. Results of steady-state simulations when thermal effects are considered. (a, b) Contour plots show a simulated vertical vibrations of a journal supported in the bearing with no grooves and in the bearing with the proposed groove layout, respectively. Summary of the results shows (c) maximum vibrations in the bearings and (d) averaged total power loss due to hydrodynamic friction. Bearing configurations are defined in Fig. 3 . (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.) d e R [ [ [ [ [ [ [ [ [ [ [ [ [ [ [ [ [ suggest that the conclusions presented above hold even if the thermal ffects are included in a computational model. eferences [1] Chen WJ. Rotordynamics and bearing design of turbochargers. Mech Syst Signal Process 2012;29(7):77\u201389. doi: 10.1016/j.ymssp.2011.07.025 . [2] Brouwer MD, Sadeghi F, Lancaster C, Archer J, Donaldson J" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001800_978-3-030-40410-9-Figure8.6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001800_978-3-030-40410-9-Figure8.6-1.png", "caption": "Fig. 8.6 a Translational compliant displacements \u03b4 and b rotational compliant displacements \u03c6", "texts": [ "3) and the Jacobian of the elastic deformation J is obtained by J = \u2202 f \u2202\u03b4 (8.12) where \u03b4 = ( \u03b41 \u03b42 \u00b7 \u00b7 \u00b7 \u03b4n )T is the translational displacements of the origins of the attached frames to the manipulator. Thus, the total compliance matrix is CT = C + C jt (8.13) The global stiffness matrix can be obtained using Eq. (8.10). Example 8.1 Consider the 2DOF planar manipulator in Fig. 8.5, where we need to find the stiffness matrix using the Komatsu formulation for the given external wrenches Fx and Fy. Figure 8.6a, b shows the linear and angular displacements to the joint 1 (point A) and end-effector (point B), respectively. M2 = 0 (8.21) From the theory of elasticity, we can formulate the translational displacements of point A and point B, which are \u03b41 and \u03b42, respectively: \u03b41 = 31 3E1 I1 F1 + 21 3E1 I1 M1 (8.22) \u03b42 = 32 3E2 I2 F2 + 22 3E2 I2 M2 (8.23) Also, from the theory of elasticity we can formulate the rotational displacement of point A and point B, which are \u03c61 and \u03c62, respectively: \u03c61 = 21 3E1 I1 F1 + 1 E1 I1 M1 (8" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002175_icra40945.2020.9197401-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002175_icra40945.2020.9197401-Figure3-1.png", "caption": "Figure 3. (a) Components of the hybrid soft-rigid joint concept (left), assembled joint (right), (b) A-A cross section of assembled joint with unpressurized (left) and pressurized (right) pressure chambers in the rigid compartment (by pressurization, increasing stress at the contact surface is represented by dashed black line), (c) assembled SCS. lSCS: 100 mm, wSCS: 10 mm, hSCS: 8 mm.", "texts": [ " Meanwhile, the extension causes curvature because of the stiffness difference [25] between the materials of the FRA (Shore Hardness: 10A) and the encapsulation layer (Shore Hardness: 30A). In addition to its strain limiting purpose, it integrates the SCS with the FRAs. B. Stiffness Control Structure The proposed stiffness concept is composed of modular rigid links and soft pressure chambers which are made of UV curable plastic and DragonSkin 10, respectively. By the integration of two rigid links using alignment pins, a rigid compartment in a revolute joint form appears, then a pressure chamber is positioned in it (Fig. 3a). Lateral and axial expansions of the pressurized chamber are constrained by the rigid compartment. Upon pressurization, the stress at the contact surface between the pressure chamber and the rigid compartment increases. This causes increasing friction at the surface and resultant stiffness of the joint (Fig. 3b). As a result, the stiffness can be easily controlled by the input pressure. The SCS includes three joints (i.e. four rigid links and three pressure chambers), all of them controlled independently (Fig. 3c). The entire SCS was aimed to be rotated between -90\u00b0 - +90\u00b0 in the x-y plane, thus the rotation of each joint is mechanically limited between -30\u00b0 - +30\u00b0 around their axial direction to add up 90\u00b0. Thanks to the modular design of the rigid links, the total rotation angle can be increased or decreased by changing the number of rigid links. All the rigid parts of the SCS and molds for silicone casting were designed using SolidWorks (Dassault Syst\u00e8mes, France) and fabricated by ProJet 3600 HD Max (3D Systems, USA) 3D printer using UV curable plastic (VisiJet M3 Crystal, 3D Systems, USA)" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000857_s0263574719000857-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000857_s0263574719000857-Figure1-1.png", "caption": "Fig. 1. Planar flexible space manipulator with an elastic base.", "texts": [ " In Section 5, experimental datum from a computer simulation by Matlab is presented. Finally, the conclusions are given in Section 6. In this section, first the kinematics analysis of a flexible space manipulator with an elastic base was made. Then the total energy associated with the manipulator system was computed using the kinematics formulations. Finally, the dynamic equations were derived by the Lagrange method. A space manipulator with flexible links Bi(i = 1, 2) and an elastic base B0 is shown in Fig. 1. OC is the mass center of the system, OC0 the mass center of the base and Oi the rotational center of the revolute joint between Bi\u22121 and Bi(i = 1, 2). \u03b80 is the attitude angle of the base and \u03b8i is the ith link angular displacement (i = 1, 2). Oi \u2212 xiyi is the local right handed coordinate frame of Bi(i = 0, 1, 2) and O \u2212 xy is the inertial right handed coordinate frame. The elastic base was dynamically simplified as a rigid body with a constant stiffness k linear spring. The spring\u2019s elastic displacement is l0", " (38) can asymptotically track the terminal trajectory qt with the link flexible vibrations and the elastic base deformation are well damped out. The operational control strategy of the proposed TSMC based on hybrid trajectory is shown in Fig. 3. https://www.cambridge.org/core/terms. https://doi.org/10.1017/S0263574719000857 Downloaded from https://www.cambridge.org/core. Carleton University Library, on 21 Aug 2019 at 12:28:19, subject to the Cambridge Core terms of use, available at To show the performance of the proposed controller, simulations are carried out on a planar space manipulator system shown in Fig. 1. The actual parameters of the system are as follows: l0 = l2 = 1.0 m, l1 = 2.0 m, m0 = 200 kg, \u03c11 = 4 kg/m, \u03c12 = 1.6 kg/m, ma1 = ma2 = 1 kg, mP = 14.4 kg, I0 = 100 kg \u00b7 m2, IP = 41.7 kg \u00b7 m2, (EI)1 = 500 N.m2, (EI)2 = 300 N.m2, Ia1 = Ia2 = 0.03kg m2, k = 5000 N/m, \u03c4 d = 10 \u00d7 [sin(0.25\u03c0 t) cos(0.25\u03c0 t) sin(0.25\u03c0 t) \u2212 cos(0.25\u03c0 t)]T. The terminal angle vector is defined as qt = [000.5\u03c0 ]T. The initial state of the space manipulator is l0 = \u03b411(0) = \u03b412(0) = \u03b421(0) = \u03b422(0) = 0 m, \u03b80(0) = 0.1 rad, \u03b81(0) = 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000110_tie.2019.2941156-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000110_tie.2019.2941156-Figure6-1.png", "caption": "Fig. 6. Section planning approach using tangent lines in the TPS algorithm.", "texts": [ " We plan the trajectory using three sections to satisfy the boundary conditions (initial and final states) using the follow- 0278-0046 (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. ing variables: two tangent lines and three turning angles. Note that \u03b8f \u2212 \u03b8s = \u03c6I + \u03c6C + \u03c6F . We suggest an initial tangent line from ITS to CS and a final tangent line from FTS to CS. Note that the asymptote of ITS is the initial tangent line while the asymptote of CS is the final tangent line, as shown in Fig. 6. First, we plan the trajectory by varying final rotation angle \u03c6F using Newton\u2019s method for optimization (the solution to f \u2032 (|\u03c6F |) = 0) in Line 1. Because the arrival time according to \u03c6F has a convex form, Newton\u2019s method for optimization can be applied to the derivative of the function [37]. Inside the outer loop, we plan ITS, CS, and FTS by varying the initial rotation angle \u03c6I as well as T2 and T5 using the 3D-Newton\u2019s method in Line 2 to satisfy the configuration conditions (dc = 0, dI = 0, and dF = 0)" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001712_hlca.19560390332-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001712_hlca.19560390332-Figure2-1.png", "caption": "Fig. 2.", "texts": [ " 38, 508 (1955); 15'. Simon, Doris Meuche & E. Heilbronner, Helv. 39. 290 (1956). Volumen XXXIX, Fasciculus 111 (1956) - Xo. 106. 585 a p p a r a t u r . In Fig. 1 ist die Titrationsanlage schematisch dargestellt. Das im vertikal verschiebbaren Reservoir 1 enthaltene Quecksilber shomt durch eine Polarograplienkapillare (Fig. 3) in die thermostatierte Burette 7. Eine genaue Einregulierung des Quecksilberdruckes kann mit Hilfe des in ern und mm eingeteilten Niveaurohres 2 erreicht werden. DerHahn 5 (vgl. 1 in Fig. 2), der durch einen Servomotor (vierUmdrehungen/Min.) hedient wird, ermoglicht ein Starten und Stoppen des Quecksilberflusses durch die Kapillare und somit auch des Ausfliessens der Titrationsfliissigkeit aus der Burette. R 11 r e t t e. -i dnsicht von oben. B Ansicht von vorne. C Ansicht von der Seite. a Xormalschliff Nr. 7 (Verbindung der Polarographenkapillare mit dem Quecksilber- reservoir 1 [siehe Fig. 11). h Normalschliff Xr. 7 (Einfugung der Polarographenkapillare in den Glasmantel der Burette)", " i Platinkontakt. k Dreiweghahn: Fiillung und Reinigung der Burette. 1 Hahn: Starten und Stoppen der Titration. Polarographenkapi l la re . a Normalschliff Nr. 7 (Verbindung der Kapillare mit dem Quecksilberreservoir 1 (siehe b Normalschliff Nr. 7 (Einfugnng der Iiapillare in den Glasmantel der Burette). Fig. 1). E i c h u ng de r Au s f l u s sge s c h w in d i g ke i t. B Batterie 1,5 Volt. R 1 Widerstand 15 K f?. B 2 Widerstand 100 R. R 3 Widerstand 40 R. 1 h, i Platinkontakte (vgl. h und i in Fig. 2). (R 1, R 2 und R 3 w-urden fur die Verwendung eines 10-mV-Schreibers gewahlt. Bei der Verwendung anderer Schreiber mussen diese Widerstande entsprechend geandert werden. ) Der Ablauf des Titrationsvorganges wird iiber ein Relais-Aggregat (Fig. 5 ) durch die Mikroschalter MS 1, MS 2 und einen dritten Mikroschalter MS 3, der sich im Schreiber befindet, gesteuert. Die Mikroschalter MS 1 und MS 2 werden durch Nocken auf der Antriebswelle zum Hahn 5 betatigt. Durch den Dreiweghahn 6 kann die Burette in kurzer Zeit aus dem Reservoir 4 mit Quecksilber gefiillt werden", " Hahn 5 offen: MS 1 offen, MS 2 geschlossen. Hahn 5 geschlossen: MS 1 geschlossen, MS 2 offen. Mikroschalter, der durch den Federschlitten des Schreibers betiitigt wird. Ruhekontakt. 220 Volt Wechselstromrelais : RL 1 : Ein Arbeitskontakt, ein Umschaltkontakt. RL 2: Zwei Arbeitskontakte. RL 3: Ein Arbeitskontakt. (RL 2 kann mit einem zusatzlichen Arbeitskontakt ausgeriistet werden, so dass die Moglichkeit besteht, mit Hilfe dieses Kontaktes eine Markierung auf dem Papier des Schreibers bei Titrationsstart auszulosen.) Biiretteka) : Die Fig. 2 zeigt die Einzelheiten der Burette. Die Polarographenkapillare die die Stromungsgeschwindigkeit des Quecksilbers bestimmt, hat einen inneren Durchmesser von ca. 50 p und ist mittels des Glasschliffes b (siehe Fig. 2 und 3) in das restliche System eingef ugt5). Mit einer einzigen Kapillare kann die Titrationsgeschwindigkeit ohne 4 ) Zur Reinigung der Burette wird diese iiber den Dreiweghahn 6 mit dem Wasserstrahlvakuum verbunden, dann werden der Reihe nach destilliertes Wasser, Methanol, Ather und filtrierte, trockene Luft hindurchgesogen. 48) Die Biirett,e wurde von der Firma Gebr. MoZZer, Glasblaserei, Zurich, hergest.ellt . 5) Um eine Verunreinigung der Polarographenkapillare zu vermeiden, ist es not- wendig, stets frisch destilliertes Quecksilber zu verwenden. Bei der Installation der Burette muss der Raum zwischen der Polarographenkapillare und dem Mantel der Burette unbedingt vollstandig mit Quecksilber gefullt werden. Dies wird dadurch erreicht, dass die Burettenspitze mit Schlauch und Quetschhahn abgeschlossen wird und bei geoffnetem Hahn 1 (Fig. 2) das Burettensystem durch Hahn k (Fig. 2) abwechslungsweise evakuiert und aus dem Reservoir 4 (Fig. 1) mit Quecksilber gefullt wird. 888 HELVETICA CHIMICA ACTA. weitmes von 70 bis 400 mm3 Titrationsfliissigkeit pro Std. durch Anderung des Quecksilberdruckes variiert werden. Der Einsatz verschiedener Kapillaren gibt ausserdem die Moglichkeit, ein noch grosseres Gebiet von Titrationsgeschwindigkeiten zii erfassen. Die Reproduzierbarkeit der Titrationsgeschwindigkeiten ist von der Grossenordnung 0,001 und besser. Zur Konstanthaltung der Viskositat des Quecksilbers und zur Elimination von Thermometereffekten muss das ganze Burettensystem auf & 0,1 O tliermostatiert werden. Das Volumen der Kapillare e (Fig. 2A), die die Titrationsflussigkeit enthalt, betragt ca. 100 mm3 (gemessen zwischen den Platinkontakten h und i). Eine Bestimmung dieses Volumens wird am besten mit Hilfe einer Wagemethodec) vorgenommen. Die Ausflussgeschwindigkeit der Titrationsfliissigkeit in mm3/min kann d a m durch eine Messung derjenigen Zeit erhalten werden, die das Quecksilber bei konstantem Druck benotigt, urn von Kontakt i zu Kontakt h zu fliessen (Fig. 2B)7)R). Meistens ist es praktischer, die Ausflussgeschwindigkeit mit der Papiervorschubgeschwindigkeit des verwendeten Schreibers zu verknupfen. Unter Verwendung des einfachen Schemas der Fig. 4 ist es moglich, den Durchgang des Quecksilbers bei den Kontakten i und h auf dem Papier des Schreibers zu markieren. Die Ausflussgeschwindigkeit der Burette kann folglich direkt in mrn3/cm Papier oder bei Kenntnis des Titers der Titrationsfliissigkeit in Aquivalenten Titrationsflussigkeit pro cm Papier angegeben werden", " Details for the construction of an automatic microtitrator are given. Organisch-chemisches La,boratorium der Eidg. Technischen Hochschule, Zurich. 6 , P. L. R i r k , Quantitative Ultramicro Analysis, p. 36, New York 1950. \u2019) Im Teil I wurde gezeigt, dass eine strenge Linearitat zwischen Ausflusszeit und Ausflussmenge besteht. 8, Bei der Bestimmung der Ausfliisszeit muss unbedingt beachtet werden, dass die Messbedingungen moglichst analog denjenigen sind, die wiihrend der Titration vorliegen. Die Kapillare c (Fig. 2A) sollte also vor Mersbeginn mit der Titrationsflussigkeit, mindestens jedoch mit Wasser gefullt werden. Die Kapillare c muss stets horizontal liegen." ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001738_j.mechmachtheory.2019.103629-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001738_j.mechmachtheory.2019.103629-Figure1-1.png", "caption": "Fig. 1. Schematic of the mechanism.", "texts": [ " Then, in Section 3 its inverse and forward kinematics are provided. In Section 4 , first the velocity and acceleration analyses of the manipulator under study are presented and then its workspace and singularities are discussed. Thereafter, in Section 5 , the principle of virtual work is employed to derive the dynamic model. In Section 6 , the derived formulations are validated by comparing the computed forces with those obtained using MATLAB Simscape Multibody. Finally, Section 7 provides a conclusion and some directions for future work. As illustrated in Fig. 1 , the proposed mechanism is an over-constrained parallel manipulator with four degrees of freedom, including three translations and infinite tool rotation. This mechanism includes four parallel linear guideways on which four actuated carts can move independently. Each cart is connected via a revolute joint to a link, which in turn is connected by a revolute joint to a rhombus linkage. The lower two links connected to the manipulated platform control the position of the end effector interface in the direction parallel to the axes of the revolute joints, while the upper two links control the position of the end effector interface and its rotation around the same axis" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001659_j.foostr.2020.100170-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001659_j.foostr.2020.100170-Figure1-1.png", "caption": "Fig. 1. Patterns of constraint solution applied on xerogel surface for programming (a) flower and (b) cone transformations. Yellow colour represents xerogel and black colour represents constraint pattern.", "texts": [ " Chemical changes were observed using fourier transform infrared spectroscopy (Spectrum two FT-IR, Perkin Elmer) equipped with Miracle ATR single reflection ZnSe crystal. Wheat flour and xerogel sides (top and bottom) were analysed over 4000 \u2013 650 cm\u2212 1 at 4 cm\u2212 1 resolution. The xerogel was programmed with ethylcellulose to constraint it\u2019s swelling in water to desired shape transformation. Ethylcellulose (10 % (w/v)) was completely dissolved in the ethyl alcohol (99 %) using vortex mixer. The solution was applied to the top surface of the different shaped xerogels in specific patterns, as shown in Fig. 1. with micropipette manually. Then the xerogel was dried under the ambient condition to remove the ethanol in the constraint solution and form a solid film deposition. The programmed xerogel was immersed in a beaker containing distilled water heated till boiling before immersion. The shape transformation from the near-flat xerogel to programmed 3D shape over time was recorded using a GigE vision area scan camera (C1600, Genie color series, DALSA) with a magnifying lens (Zoom 7000, Navitar). The gelatinized wheat flour components were in two phases majorly comprised of amylose and amylopectin polymers" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001666_iros45743.2020.9341051-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001666_iros45743.2020.9341051-Figure5-1.png", "caption": "Fig. 5: GNC-system including the localization module mounted on the \u00b5AUV platform HippoCampus.", "texts": [ " Let the vector eR describe the robot\u2019s orientation error with regard to the desired orientation Rdes = R (\u0398des) eR = eRBx eRBy eRBz = 1 2 ( R>desR (\u0398)\u2212R (\u0398) > Rdes )\u2228 , (13) where \u2228 is the vee map operator which is used for the transformation from SO (3) to R3. Note that for the case of large orientation errors, we temporary reduce u1 to prioritize the minimization of the orientation error. Moreover, we define the error of the corresponding angular velocities as e\u03c9 = \u03c9 \u2212 \u03c9des. (14) Thus, the resulting control outputs yield[ u2 u3 u4 ]> = \u2212KReR \u2212K\u03c9e\u03c9, (15) where KR and K\u03c9 are diagonal matrices which allow convenient gain tuning of the individual components. The \u00b5AUV localization module mainly consists of three components, depicted in Fig. 5: a wide-angle camera, an SBC and a FCU. The camera is a low-cost wide-angle RaspberryPi camera with an opening angle of 140\u25e6. Note that the effective opening angle is approximately 120\u25e6 due to rectification. We reduce the resolution to 640 \u00d7 480 px in order to trade-off accuracy against processing time which allows the visionsystem to run at 10 Hz. Note that other camera mounting orientations are possible to adapt the system to the individual robot design. A Raspberry Pi 4 with 4 Gb RAM is chosen as an on-board SBC running the vision processing and the EKF. The PixRacer-platform is used as an FCU running the PX4-firmware [33] including a dedicated \u00b5AUV attitude controller. This modular design allows to physically split the low-frequent components e. g. the vision processing from the high-frequent components such as the attitude controller. Figure 5 shows the complete module in a 3D-printed rack which fits the dimensions of the HippoCampus \u00b5AUV, see Fig 4. However, all system components fit into a total volume of approximately 90\u00d7 50\u00d7 30 mm. The system is powered with 5 V and its power consumption is 9 W at full load which is comparatively low in comparison to the high power consumption of the four thrusters. For more details on the HippoCampus platform we refer the reader to [2], [35]. We evaluate the performance of our system in two experimental settings" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001197_s42405-020-00259-6-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001197_s42405-020-00259-6-Figure6-1.png", "caption": "Fig. 6 Boundary conditions and typical loadcase of a conventional honeycomb panel", "texts": [ " It must be emphasized that the mesh quality must be guaranteed in these calculations, otherwise abnormal data will appear and destroy the above range. This requirement can be realized using the quality check function of the software. By optimizing the meshing of elements, the quality indexes such as aspect ratio, warpage, skew, and Jacobian value of all shell elements must comply with the element check criterion of RADIOSS/OptiStruct, which is built into the software. A square conventional honeycomb panel with thickness of 20 mm and side length of 200 mm is shown in Fig. 6, of which the parameters of sandwich structure are consistent with the example in the former section. There are four inserts on each corner and one insert in the center of the panel. The equivalent modeling method proposed in this paper was used to model the panel, and the bonding position between inserts and honeycomb sandwich structure was simulated by node fitting. Assume a simplified typical loadcase: the honeycomb panel is placed vertically, the inserts at four corners are fixed, and a weight of 10 kg (converted into a force along Y axis direction, marked as F in Fig. 6) is mounted on the central insert.Under the above conditions, the stress anddeformation of the entire honeycomb sandwich structure were analyzed, as shown in Fig. 7. Table 7 Equivalent elastic parameters of interlayer in the payload mounting panel E1 (MPa) E2 (MPa) E3 (MPa) G12 (MPa) G23 (MPa) G31 (MPa) \u03bc12 \u03bc23 \u03bc31 \u03c1c (g cm\u22123) 0.035 0.035 646.63 0.021 135.10 90.07 0.99986 0.00002 0.33 0.0259 To distinguish the inserts from the honeycomb sandwich structure, the inserts are not color-mapped in the quarter model of Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001793_s12652-020-01781-x-Figure9-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001793_s12652-020-01781-x-Figure9-1.png", "caption": "Fig. 9 Robot testing environment (Li 2011)", "texts": [], "surrounding_texts": [ "Based on the improved algorithm of robot kinematics error identification in the third section above, the simulation and comparison experiments are carried out with the least square method. Referring to the simulated experimental equipment diagram and experimental environment diagram shown in Figs.\u00a08 and 9, the main experimental robots are composed of genetic algorithm controller, sixdegree-of-freedom parallel robot model and corresponding control modules (corresponding experimental devices and experimental robots cited in reference (Li 2011). In order to control the variables of the simulation experiment, three groups of experiments were carried out based on the least square method and the algorithm proposed in this paper. The corresponding experimental conditions are: 1, 4 test points and 1 repetition. 2. The number of test points is 4 1 3 and the number of repetitions is 10. 3. The number of test points is 16 and the number of repetitions is 10. The corresponding experimental results are as follows: 1. The number of test points is 4, and the number of repetitions is 1. The parameter identification error of the connecting rod based on the least squares method is shown in Table\u00a02. The identification error of the connecting rod parameters based on the algorithm is shown in Table\u00a03. The corresponding tables a, b, and c are the angles on the three-coordinate vector error. It can be seen from Tables\u00a02 and 3 that the algorithm proposed in this paper has obvious advantages. Figure\u00a010a, b correspondingly, the angular trajectories and errors of parallel six-degree-of-freedom robots around Z-axis are obtained by using the proposed algorithm. 2. The number of test points is 4, and the number of repetitions is 10. The parameter identification error of the connecting rod based on the least squares method is shown in Table\u00a04. The identification error of the connecting rod parameters based on the algorithm is shown in Table\u00a05. The corresponding tables a, b, and c are the angles on the three-coordinate vector error. As can be seen from Tables\u00a04 and 5, the algorithm proposed in this paper has obvious advantages. Figure\u00a011a, b correspondingly, the angular trajectories and errors of parallel six-degree-of-freedom robots around Z-axis are obtained by using the proposed algorithm. 3. The number of test points is 16, and the number of repetitions is 10. The parameter identification error of the connecting rod based on the least squares method is shown in Table\u00a06. The identification error of the connecting rod parameters based on the algorithm is shown in Table\u00a07. The corresponding tables a, b, and c are the angles on the three-coordinate vector error. As can be seen from Tables\u00a06 and 7, the algorithm proposed in this paper has obvious advantages. Figure\u00a012a, b correspondingly, the angular trajectories and errors of parallel six-degree-of-freedom robots around Z-axis are obtained by using the proposed algorithm." ] }, { "image_filename": "designv11_5_0001452_j.mechmachtheory.2020.104061-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001452_j.mechmachtheory.2020.104061-Figure4-1.png", "caption": "Fig. 4. Ring gear under moving meshing loads of the planets.", "texts": [ " \u03b81 and \u03b82 are the angular coordinations of the two nodes of the element, and D, B , and H are matrices given in Appendix B . 2.3.2. Moving meshing force f rp The ring gear of a PGT is subjected to vibration under multiple moving meshing forces, f rp,n (n = 1 , 2 , . . . , N) , and due to the elastic behavior of the ring gear applying these moving loads on the model in a proper way is significant. The ring gear is supported by six equally-spaced elastic supports while moving loads, caused by the planet-ring meshing pairs, rotate throughout the inside of the ring gear (see Fig. 4 ). Each moving load is assumed to have a constant angular velocity, \u03c9 c , and is at a fixed angular distance 2 \u03c0 / N form its neighbors. The acting point of moving loads are: \u03b8r,n = \u03c9 c t + 2 \u03c0( n \u2212 1 ) N n = 1 , . . . , N (14) where \u03b8 r,n is the angular coordination of the n th moving load inside the ring gear. In the following, the process leads to the ring gear dynamic equation under moving meshes were presented. To develop the ring gear dynamic equation under the moving load, the meshing force in the rp mesh should be defined" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002614_j.compositesa.2021.106449-Figure19-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002614_j.compositesa.2021.106449-Figure19-1.png", "caption": "Fig. 19. The illustration of the hole cut of FE mesh (a) 900, (b) -450.", "texts": [ " Four plane cut are performed in sequence to obtain perfect straight edges, which allows the specific boundary conditions to be applied in FE analysis. The hole cut, however, requires a mesh boolean process. This can be done in the FE package or 3D computer graphics software, such as Blender or OpenSCAD, which is used for backends mesh boolean computation. The mesh file of a cylinder is used as a hole cutter to trim the geometries. The open-hole geometry of 900 and -450 are used for demonstration purpose as shown in Fig. 19. It can be noted the mesh near the cutting boundary can be very distorted. This issue can be tackled by refining the boundaries or combining the distorted elements into adjacent elements such as the built-in function in Abaqus, \u201ccollapse edge (tri/quad)\u201d. The main advantage of the mesh trimming approach is to preserve the structured tow mesh within the part (at the expense of distorted mesh at the boundaries). An alternative to FE mesh trimming way is to cut the geometry prior to using an unstructured meshing algorithm on the irregular trimmed tow shapes" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001620_tbme.2020.3043388-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001620_tbme.2020.3043388-Figure4-1.png", "caption": "Fig. 4. Workspace analysis of REP. X-Y-Z is the world coordinate system, X0-Y0-Z0 is the REP coordinate system, P0 is the origin of REP coordinate system and the measuring point of the magnetic sensor, P1 is the section center of the biopsy forceps at the front end of the REP, P2 is the end of biopsy forceps, L is the extended length of biopsy forceps.", "texts": [ " In colonoscopy intervention, it is important to know the real-time end position of the biopsy forceps, such that it can successfully extend to the polyp. The collection of all possible end positions of the biopsy forceps represents the reachable workspace of this system and depends on both the pose of the REP and the extension length of the biopsy forceps. Therefore, the modeling of the workspace for predicting the end position of the biopsy forceps is shown here. To describe the end position of biopsy forceps, a world coordinate system is defined as the center of the magnetic source of the magnetic tracking system. As shown in Fig. 4, X-Y-Z defines the world coordinate system, and X0-Y0-Z0 is used to describe the local REP coordinate system. P0 is the measuring point of the magnetic sensor which is used to measure the pose of the REP, and is also the origin of the REP coordinate system. P1 is the section center of the biopsy forceps at the front end of the REP, and P2 is the end of biopsy forceps. L is the extended length of biopsy forceps from the front of the REP and along the direction of Y0 in the REP coordinate system. For simplification, the initial REP coordinate system is parallel to the world coordinate system, in which X, Y, Z axes have the same directions as that of X0, Y0, Z0 axis", " Pc is the origin of the camera coordinate system, the physical distance between Pc and the image plane is the focal length f. P is a random point in 3D space, (xp, yp ) is the projection of P on the image plane that expressed as pixel coordinates, and (x0, y0 )is the pixel coordinates of the camera center in the image plane. The coordinates of P in the camera coordinate system can be calculated by the following equation. \u23a1 \u23a3 Xc p Y cp Zcp \u23a4 \u23a6 = Zcp \u23a1 \u23a3 (xp \u2212 x0) /f (yp \u2212 y0) /f 1 \u23a4 \u23a6 (5) Considering the coordinate system transformation from the camera coordinate system to the world coordinate system shown in Fig. 4, the coordinates of point P in world coordinate system can be expressed as (6), in which [Xc p Z c p \u2212Y cp ] T is the rotated coordinates of point P from the camera coordinate system to REP coordinate system. \u23a1 \u23a3 Xp Yp Zp \u23a4 \u23a6 = R\u03c8,Z R\u03b8,YR\u03d5,X \u239b \u239d \u23a1 \u23a3 Xc p Zcp \u2212Y cp \u23a4 \u23a6+ P 0 c \u239e \u23a0+ P0 (6) Here, P 0 c is the position vector of Pc relative to P0 in the initial REP pose in the world coordinate system. As there is only a single camera on the REP, it cannot obtain the absolute depth of the target point P, which refers toZcp in (6)" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000557_aim.2016.7576989-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000557_aim.2016.7576989-Figure5-1.png", "caption": "Figure 5. Cross-sectional diagram of unit", "texts": [ " Using this continuous process, the bowel can transport boluses by application of a small force. We mimic the function of one circular muscle for one unit. The circular muscle is mimicked by a cylindrical tube. The straight-fiber-type artificial muscle mimics the longitudinal muscle. Each unit has circular contraction and relaxation and an axial contraction movement in order to imitate bowel peristalsis. The cross-sectional diagram of the peristaltic conveyor unit, the dimensions of the unit, and the unit specifications are shown in Fig. 5, Fig. 6, and Table I, respectively. We use a straight-fiber-type artificial muscle, a cylindrical tube, and flanges. The cylindrical tube is arranged inside the artificial muscle. The flange side has a vent for the supply and exhaust of the air with the vent connected to the chamber between the artificial muscle and the cylindrical tube. Upon application of air pressure, the artificial muscle expands on the outside, and the cylindrical tube expands on the inside. Fig. 7 shows the unit in the normal and pressurized states" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001158_s00170-020-04957-z-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001158_s00170-020-04957-z-Figure2-1.png", "caption": "Fig. 2 The position of the cutting point changes after the hob moves in the X-direction", "texts": [ " Through radiation, convection, and conduction, some of the heat diffuses into the air, while the rest is absorbed by the major components such as bed, coolant liquid, and cutter. The temperature gradient could lead to the tilting of the upright and the elongation of the bed. So, the relative displacement between hob and the center of workpiece axis is changed, as shown in Fig. 1. The change of the distance between the hob and the workpiece spindle is usually defined as the thermal error in X-direction, which is used to reflect the error in the center distance. As shown in Fig. 2, when the thermal error \u03b4x occurs, the cutting point moves from K toK\u2032. The lineKK\u2019 is perpendicular to the gear reaming, and it is defined as the total error of tooth profile. The calculation formula of KK\u2019 can be expressed as F\u03b1 \u00bc KK 0 \u00bc \u03b4xsin\u03b1 \u00f01\u00de where F\u03b1 represents the total error of tooth profile; \u03b1 is the angle between KK' and the horizontal direction. The error F\u03b1 will lead to the changes of the addendum, dedendum, tooth thickness, and the diameters of apex and root circles. This kind of error is relatively easy to eliminate, and it can be compensated by adjusting the feed speed in X-direction" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001934_j.mechmachtheory.2020.103895-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001934_j.mechmachtheory.2020.103895-Figure2-1.png", "caption": "Fig. 2. Three stiffness models of the Delta robot. From left to right: the fully flexible model 1, the flexible model 2 with rigid short sides of the parallelogram, and the model 3 with flexible arms P and forearms D.", "texts": [ " [1] : n O = n B \u2212 n J \u2212 n R \u2212 n F (1) The previous expression can be easily verified remembering that joints and rigid connections couple two boundary nodes and express the nodal displacements of one boundary node in terms of those of the other: it means that only one of the two boundary nodes can be selected as an independent node. Finally, each fixed node can be directly removed from the computation of Eq. (1) . To better explain this result, the Delta robot [55,56] , whose one of the three legs is shown in Fig. 1 , will be used as a demonstrative example considering three models with decreasing complexity, evaluated in terms of the number of dofs . The first model 1 shown in Fig. 2 is a fully flexible model in which all links are deformable. In model 2 the short sides of the parallelogram are rigid while in model 3 only the forearms D and arms P are flexible while the remaining parts are rigid. All boundary nodes have been marked with blank circles. The end-effector node E has been added to the set of boundary nodes. Starting from the fully flexible model 1, considering that the Delta robot has three legs connecting the base platform B to the moving platform M , by applying Eq", " The first rule comes from the fact, already mentioned, that only one boundary node of the two nodes involved in the displacement constraints created by a joint or a rigid connection can be independent. The same rule imposes that if several boundary nodes belong to the same rigid body at most only one of these is independent. The second rule states that all those nodes with imposed displacements must be considered independent. Fixed nodes are a special case of imposed displacements and can be removed from the final model because their displacements are all zero. The rules must be applied to each subsystem obtained cutting the flexible parts of a mechanism. As an illustrative example, in Fig. 2 the subsystems of the Delta robot are displayed with gray areas for the three stiffness models. As it can be observed by applying the rules to each subsystem of a model displayed in Fig. 2 , the choice of the independent boundary nodes is not unique. Certainly, some choices are to be preferred as these reduce the complexity of the final elastodynamic model. For example, considering the third model of Fig. 2 the subsystem of M includes the lower sides of the three legs (in the same figure only one leg is represented). By applying the two rules only one independent boundary node can be selected. Among the many choices, the boundary node on the end-effector point E should be preferred for the reason of symmetry. In the previous subsection, the independent boundary nodes have been determined. These nodes allow for expressing all the other boundary nodes by employing the displacement constraints generated by the joints" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001817_s40964-020-00123-9-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001817_s40964-020-00123-9-Figure3-1.png", "caption": "Fig. 3 EBSD analyses of the three simple geometries: a thin strut, b cube and c wall. The inverse pole figure (IPF) colouring orientation maps, and the {100} and {110} pole figures are included", "texts": [ "5\u00a0mm \u00d7 3\u00a0mm \u00d7 2\u00a0mm were cut from the LPBF-fabricated walls using electrical discharge machining (EDM), and were tested by a Satec 60UD tensile machine equipped with an Instron AVE 2 video extensometer. The microhardness testing was conducted using a CM-100AT Clark microhardness indenter with a load of 100\u00a0g and a holding time of 15\u00a0s. More than twelve indentations were taken for each plane. The simple geometries in this paper refer to the thin struts, cubes and walls whose laser scanning area remained unchanged throughout the LBBF process. Figure\u00a03 reveals the crystallographic texture of the three simple geometries manufactured in the present study. As shown in Fig.\u00a03a, the strut exhibited a complete single-crystalline core with its \u2329110\u232a crystallographic direction aligned with the BD. No stray grains were observed inside the strut core. The {100} and {110} pole figures in Fig.\u00a03a also confirm the single-crystalline nature of the strut core by showing a single orientation. This feature could also be observed in other thin struts produced using the same LPBF processing parameters. The crystallographic texture of the cube is shown in Fig.\u00a03b. It can be clearly observed that the \u2329110\u232a \u2225 BD crystallograhic texture was maintained. Furthermore, 1 3 the \u2329100\u232a crystallographic direction of the cube was observe to be aligned with the X axis. Therefore, the cube featured a {110}\u2329001\u232a Goss texture on the XY plane. Noted that the texture intensities in the pole figures were not as concentrated as those in the thin strut due to the presence of some stray grains. Therefore, the cube sample will be hereafter described as \u201csingle-crystalline-like\u201d" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002502_j.tws.2021.107512-Figure22-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002502_j.tws.2021.107512-Figure22-1.png", "caption": "Fig. 22. Transverse displacements and stresses in the radial direction near the ifurcation point.", "texts": [ ", digital image correlation (DIC) measurement systems. .3. Stress distribution of R3C5 The instability pattern \ud835\udc453\ud835\udc365 is simulated by the reduced model to study the stress distributions. As shown in Fig. 21, near the bifurcation point, both stresses exhibit a periodic pattern in the circumferential direction and a nearly periodic one in the radial direction. The region of maximum circumferential stress appears in the vicinity of the center, W while there does not seem to be any unique region of maximum radial stress. As shown in Fig. 22 for the instability pattern \ud835\udc453\ud835\udc365, both the radial nd circumferential stresses present nearly periodic pattern in the radial irection. The stresses on the top surface tend to be higher at the peak, hile the ones on the bottom surface tend to be higher at the trough. herefore, the direction of the transverse displacement determines the urface to be stretched, i.e., the top surface of the crest and the bottom urface of the trough tend to be stretched in the post-buckling stage. he amplitude of the radial stress remains at the same level, while hat of the circumferential stress decreases in the radial direction" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002073_lra.2020.3013862-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002073_lra.2020.3013862-Figure4-1.png", "caption": "Fig. 4. The d{1,2},G, the overall CoG position of the platform w.r.t. the body coordinate system of the TP, can be changed by manipulating the relative attitude \u03b1{r,p}.", "texts": [ " However, unlike conventional quadrotors, the T 3-Multirotor has additional actuators other than thrusters, which are two servomotors that control relative roll and pitch attitude \u03b1{r,p} between TP and FP. Therefore, if we can take advantage of two servomotors to generate some of the components of u independently of c0, then we can use this feature to obtain additional redundancy in control and utilize in fail-safe flight. 1) Obtaining Additional Control Redundancy: When controlling the attitude of the FP via the servomechanism, the center of gravity (CoG) position with respect to the body coordinate of the TP changes with \u03b1{r,p}. Fig. 4 shows an example of the y-directional CoG position being changed in accordance of the relative roll attitude \u03b1r. In the same manner, we can also change the CoG position in the x-direction with relative pitch attitude \u03b1p. The position of the altered CoG with respect to the TP frame is represented by dG = [d1,G d2,G d3,G] \u1d40 \u2208 R3\u00d71, and the relationship between the relative attitudes and the CoG position is as follows. dG (\u03b1r, \u03b1p) = 1 M (mT03\u00d71 +mF (dT +Rr(\u2212\u03b1r)Rp(\u03b1p)(\u2212dF ))) (5) Authorized licensed use limited to: Cornell University Library" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001242_s11432-019-2679-5-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001242_s11432-019-2679-5-Figure1-1.png", "caption": "Figure 1 (Color online) Reference frames.", "texts": [ " It is given by \u033a(\u03b6) , (\u03b6T\u03b6 \u2212 \u03b62o )/(2\u03b5\u03b6\u03b6 2 o + \u03b52\u03b6) with \u03b6 \u2208 R n, where \u03b5\u03b6 \u2208 R + denotes a projection tolerance bound, and \u03b6o \u2208 R denotes a projection norm bound. The projection operator Proj : Rn \u00d7 R n \u2192 R n is defined as [38] Proj(\u03b6, \u03be) , \u03be, if \u033a(\u03b6) < 0, \u03be, if \u033a(\u03b6) > 0 \u2227 \u25bd\u033a(\u03b6)\u03be 6 0, \u03be \u2212 (\u25bd\u033a(\u03b6))T\u25bd\u033a(\u03b6) \u2016\u25bd\u033a(\u03b6)\u20162 \u03be\u033a(\u03b6), if \u033a(\u03b6) > 0 \u2227 \u25bd\u033a(\u03b6)\u03be > 0, (2) where \u03be \u2208 R n and \u25bd\u033a(\u03b6) = [\u2202\u033a(\u03b6)/\u03b61, . . . , \u2202\u033a(\u03b6)/\u03b6N ]T. According to the definition of the projection operator, it renders (\u03b6 \u2212 \u03b6\u2217)T(Proj(\u03b6, \u03be)\u2212 \u03be) 6 0, (3) where \u03b6\u2217 \u2208 R n denotes the true value of the parameter \u03b6. According to [1], the kinematics of ASVs as shown in Figure 1 can be expressed by \u03b7\u0307 = J(\u03c8)\u03bd, (4) where \u03b7 = [x, y, \u03c8]T \u2208 R 3 denotes the earth-fixed position (x, y) and heading \u03c8; J(\u03c8) \u2208 R 3\u00d73 is the transformation matrix between the body-fixed and earth-fixed reference frame; \u03bd = [u, v, r]T \u2208 R 3 denotes the surge, sway and angular velocity in body-fixed frame. The kinetics of ASVs can be represented by [1] M\u03bd\u0307 = \u03c4s \u2212 C(\u03bd)\u03bd \u2212D(\u03bd)\u03bd + g(\u03bd, \u03b7) + \u03c4w(t), (5) with J(\u03c8) = j11 j12 0 j21 j22 0 0 0 j33 , M = m11 0 0 0 m22 m23 0 m32 m33 , C(\u03bd) = 0 0 c13 0 0 c23 c31 c32 0 , D(\u03bd) = d11 0 0 0 d22 d23 0 d32 d33 , j11 = cos(\u03c8), j22 = cos(\u03c8), j33 = 1, j12 = \u2212 sin(\u03c8), j21 = sin(\u03c8), c13 = \u2212m22\u03bd \u2212 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000647_s0081543816080186-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000647_s0081543816080186-Figure3-1.png", "caption": "Fig. 3. Body with two internal masses and an internal rotor.", "texts": [ " To this end, we analyze the free motion of an arbitrary body at nonzero circulation of the fluid around it. We show that in this case one can perform control by using the rotor alone. We also show that the presence of circulation leads to the necessity of stabilizing the final position of the body. We obtain conditions under which such a stabilization is possible. Consider the two-dimensional problem of the motion of a hydrodynamically asymmetric body with mass M and central moment of inertia I in an infinite volume of an ideal incompressible fluid (Fig. 3). We will assume that two material points with masses m1 and m2 and a rotor with mass mr and central moment of inertia Ir are contained inside the body. The motion of the material points is restricted by the shell of the body, inside which they can move along arbitrary smooth trajectories \u03c11 = (\u03be1(t), \u03b71(t)) and \u03c12 = (\u03be2(t), \u03b72(t)). The rotor has the shape of a circular cylinder, is homogeneous, and rotates with angular velocity \u03a9(t); the rotation axis of the rotor is perpendicular to the plane of motion of the body and passes through the center of mass of the rotor", " Such a set of control elements allows one to vary the following four parameters of the body in a consistent way: the coordinates of the center of mass, the moment of inertia, and the intrinsic angular momentum. Note that using various combinations of control elements, one can obtain different control systems, including those with deficiency of controls. To describe the motion of the body, we define two Cartesian coordinate systems: a fixed one, Oxy, and a moving one, O1\u03be\u03b7, attached to the body (see Fig. 3). The point O1 coincides with the position of the center of mass of the \u201cbody + rotor\u201d system. The position of the body in absolute space is characterized by the radius vector r = (x, y) and the angle of rotation \u03b1 of the moving PROCEEDINGS OF THE STEKLOV INSTITUTE OF MATHEMATICS Vol. 295 2016 coordinate system with respect to the fixed system. Thus, the configuration space of the system is H = R 2 \u00d7 S 1, and the pair (r, \u03b1) completely characterizes the position and orientation of the body. Let us denote by v = (v1, v2) the absolute velocity of the point O1 of the body referred to the axes of the moving coordinate system, and let \u03c9 be the angular velocity of the body", " However, below we will show that since the velocities depend in a nonlinear way on controls (because x\u0307 and y\u0307 depend on \u03b2i and \u03b2\u0307i), not every trajectory can be followed by a controlled motion. Let us introduce a natural parameterization of the trajectory. Then the velocity of the body can be represented as r\u0307(t) = v\u0302(s)\u03c4 (s), (3.18) where s is the length of the trajectory passed in time t and v\u0302(s) and \u03c4 (s) = (cos \u03b3(s), sin \u03b3(s)) are the absolute value of the velocity and the unit vector defining its direction (tangent to the trajectory) at the point s (see Fig. 3). It follows from the condition of absence of cusps that v\u0302(s) does not vanish on the trajectory. Substituting (3.18) into the first three equations of system (3.7) and making a change of variables and time, \u03b81 = \u03b21 + \u03b22 2 , \u03b82 = \u03b22 \u2212 \u03b21 2 , \u03b4 = \u03b3 \u2212 \u03b1, v\u0302(t) dt = ds, (3.19) we obtain a system of differential equations A(\u03b8)\u03b8\u2032 = \u03c4 rel, \u03b4\u2032 = \u03b3\u2032 \u2212 (b(\u03b8),\u03b8\u2032), (3.20) where A11 = a2 sin \u03b81 cos \u03b82K, A21 = \u2212a1 cos \u03b81 cos \u03b82K, A12 = a2 cos \u03b81 ( sin \u03b82K + 4m2r2(r sin \u03b82 +R sin \u03b81) ) \u2212 8m3r3 cos \u03b81 sin \u03b82 cos 2 \u03b82, A22 = a1 sin \u03b81 ( sin \u03b82K + 4m2r2(r sin \u03b82 +R sin \u03b81) ) \u2212 4m2r2 ( 2mr sin \u03b81 sin \u03b82 + a1R ) cos2 \u03b82, b1 = a1a2K 2mr \u2212D, b2 = \u22122mr cos \u03b81 cos \u03b82 ( Ra1a2 \u2212 2mr sin \u03b81 sin \u03b82(a1 \u2212 a2) ) , K = 2mr ( b+ 2mR2 + 2mRr sin \u03b81 sin \u03b82 ) , D = a1a2 ( b+ 2mR2 + 2mr2 + 4mRr sin \u03b81 sin \u03b82 ) \u2212 4m2r2 cos2 \u03b82 ( a1 cos 2 \u03b81 + a2 sin 2 \u03b81 ) , the prime denotes differentiation with respect to the variable s, \u03c4 rel = (cos \u03b4, sin \u03b4) is the unit vector tangent to the trajectory and referred to the axes attached to the body, and \u03b3\u2032 has the meaning of the curvature of the trajectory", "26) contains three equations and four unknown functions \u03c1, \u03c8, \u03d5, and \u03a9 and cannot be uniquely solved without invoking additional considerations. Below, we examine several partial statements of the problem in which we can give an answer to the question posed. 4.3.2. Complete stabilization of the body. In the case of complete stabilization of the body, we assume that the conditions x\u0307 = y\u0307 = 0 and \u03b1\u0307 = 0 are satisfied for t > T . In this case the following proposition holds. Proposition 4.3. The complete stabilization of the body during an infinite time interval is possible if and only if the center of the body (point O1 in Fig. 3) lies on the circle x = \u03b6 \u03bb + \u221a \u03c72 + \u03b62 \u03bb cos \u03b8, y = \u2212\u03c7 \u03bb + \u221a \u03c72 + \u03b62 \u03bb sin \u03b8, \u03b8 \u2208 [0, 2\u03c0), (4.27) and its orientation is defined by the formula \u03b1 = \u03b8 \u2212 arctan \u03b6 \u03c7 \u2212 \u03c0 2 . (4.28) In this case, the corresponding control actions are zero; i.e., the moving mass and the rotor are at rest. Proof. Necessity. Setting \u03b1\u0307 = 0 in equations (4.25), we obtain the following equations for the controls: m1\u03be\u03071 = x cos\u03b1+ y sin\u03b1+ \u03c7, m1\u03b7\u03071 = \u2212x sin\u03b1+ y cos\u03b1+ \u03b6, Ir\u03a9+m1(\u03be1\u03b7\u03071 \u2212 \u03be\u03071\u03b71) = \u03c72 + \u03b62 2\u03bb \u2212 x2 + y2 2\u03bb . (4" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002189_j.microc.2020.105584-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002189_j.microc.2020.105584-Figure1-1.png", "caption": "Fig. 1. Housing where LEDs and photoresistors are inserted (A), together with the scheme of the employed circuit (B).", "texts": [ "283 mM) were prepared by adding weighed amounts to controlled volumes of high purity deionized water. For the sake of comparison, all tests were also performed with a commercial benchtop spectrophotometer (Varian Cary 50 bio, Victoria, AUS) by using conventional polymethylmethacrylate (PMMA) cuvettes (optical path = 1 cm). The portable spectrophotometer consisted of five photoresistors and five LEDs placed in a suitable plastic housing in which cuvettes containing samples, used for absorbance measurements, could be inserted. This housing, shown in Fig. 1A, was made in polylactic acid (PLA) at LAMA FVG (Advanced Mechatronics Laboratory of the University of Udine), exploiting the 3D Ultimaker 2 printer and the CURA software. Its dimensions were 1.8 cm width, 3.7 cm thickness and 8.9 cm height and it could hold up to a maximum of 10 pairs LDR-LED. The distance from LED to cuvette and from LED to the photoresistor were 0.05 cm and 1.3 cm, respectively. During the design and printing phases, particular attention was paid to the precise alignment of light sources with the corresponding detection area (photoresistor). Light sources consisted of five LEDs characterized by different emission spectra. In particular, LEDs with a maximum emission wavelength of 454, 479, 513, 580 and 661 nm were chosen in order to cover the entire visible region. Five photoresistors LDR of the same type (CdS, 1\u201350 K\u2126) were used for signal detection [42]. Each photoresistor was connected to the Arduino microcontroller (Arduino Nano) through a voltage divider (Fig. 1B). The circuit consisted of two resistances in series that divided the voltage provided by a potential source into two aliquots. The output voltage (Vout) between two generic heads of the circuit was calculated by multiplying the voltage applied to the whole circuit (V) for the ratio of the considered resistance (R1) to the the sum of the two resistances placed in series (R1 and R2), according to Eq. (1). Vout = [R1/(R1 + R2)]V (1) In the circuit used for assembling the device, a variable photoresistance (1\u201350 K\u2126) was used as R2, while a resistance of 68 K\u2126 was used as R1" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002166_s10846-020-01259-0-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002166_s10846-020-01259-0-Figure5-1.png", "caption": "Fig. 5 Schematic diagram of magnetometer calibration", "texts": [ " However, it is difficult to accurately load a rotation vector onto a certain axis of the gyroscope in practice due to the error in the process of mounting the chip on the PCB and mounting the PCB on the measuring block. Therefore, the gyroscope is also calibrated by multiple loading methods. The tools used to calibrate the gyroscope is the same as that used to calibrate the accelerometer (Fig.2). During the experiment, the reference block is placed on the rotating platform with its rotation vector perpendicular to the plane of the turntable, as in Fig. 5. Since the Z axis of the reference block is perpendicular to the bottom surface, the rotation vector coincides with the Z axis of the reference block. That is, the coordinates of the input rotation vector in the reference block coordinate system are expressed as[ 0 0 \u03c9L ]T , where, \u03c9L is the value of the input rotation vector. The measuring block can constrain all its three rotational DOFs (Degrees of freedom) through its fitting with Vgroove. That is, the attitude of the measurement block in the reference block coordinate system is determined and known" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001075_ecce.2019.8912286-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001075_ecce.2019.8912286-Figure1-1.png", "caption": "Fig. 1. Magnetically active portions of (a) a conventional radial flux coaxial magnetic gear with surface PMs and (b) a radial flux coaxial magnetic gear with Halbach arrays and air cores.", "texts": [ " INTRODUCTION Magnetic gears perform the same function as mechanical gears, transforming power between low-speed, high-torque rotation and high-speed, low-torque rotation. However, unlike mechanical gears, magnetic gears rely on modulated magnetic fields instead of physically interlocking teeth [1]-[3]. This noncontact operation results in a plethora of potential benefits, including improved reliability, reduced maintenance, and reduced acoustic noise. Thus, magnetic gears have attracted significant interest for a wide variety of applications, from wind [4] and wave [5], [6] energy harvesting to electric vehicles [7] and aerospace platforms [8], [9]. Fig. 1(a) shows the magnetically active portions of a conventional radial flux coaxial magnetic gear with surface permanent magnets (SPMs). The gear consists of three rotors: an inner low pole count rotor (Rotor 1), an intermediate rotor comprised of ferromagnetic pieces (known as \u201cmodulators\u201d) separated by nonmagnetic slots (Rotor 2), and an outer high pole count rotor (Rotor 3). Usually, Rotor 1 is the high speed rotor and Rotor 2 or Rotor 3 is the low speed rotor, with the other rotor being fixed in place", " Halbach arrays have previously been applied to magnetic gears to increase torque density, improve efficiency, and reduce torque ripple [10]-[13]. Additionally, because Halbach arrays concentrate flux on one side of the array and reduce it on the other side, gears with Halbach arrays may achieve high torque densities without back irons [11]. Replacing the back irons with a lighter nonmagnetic material, such as plastic, can significantly reduce a design\u2019s mass. Because the nonmagnetic material has similar magnetic properties to air, such a design is described as having an \u201cair core.\u201d Fig. 1(b) illustrates the magnetically active portions of an example radial flux magnetic gear with Halbach arrays and air cores on both rotors. As minimizing mass is critical for aerospace applications, NASA has developed two prototypes with Halbach arrays [8]. Magnetic gears present several fabrication challenges. First, the modulators must be supported between the two sets of permanent magnets (PMs). These PMs create strong magnetic forces, which can cause the modulators to bend into the air gaps [14]" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002613_j.procir.2021.03.034-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002613_j.procir.2021.03.034-Figure1-1.png", "caption": "Fig. 1. Kinematics of the gear skiving process", "texts": [ " Skiving is suitable for the manufacturing of internal and external gears as well as worm gears and is mainly used when gear hobbing is not possible [4]. This is the case when manufacturing internal gears and gear shafts with interference [2]. Due to ongoing developments in the field of CNC control of machine tools, series production is possible using this process [5]. With hard skiving, a process variant is also available that enables hard machining of gears [1]. The kinematic of the skiving process is shown in Fig. 1 and can be described as follows: The tool is tilted relative to the workpiece by the axis cross angle \u03a3. The axis cross angle \u03a3 is the sum of the helix angles of the tool and the workpiece [6]. In the case of internal gears, the value of the tool helix angle \u03b20 is negated. If the axis cross angle \u03a3 is increased, the generating motion becomes a screw rolling motion [2]. Thus, the skiving kinematics is similar to the kinematics of two helical gears m shing [6]. Available online at www.sciencedirect", " Skiving is suitable for the manufacturing of internal and external gears as well as worm gears and is mainly used when gear hobbing is not possible [4]. This is the case when manufacturing internal gears and gear shafts with interference [2]. Due to ongoing developments in the field of CNC control of machine tools, series production is possible using this process [5]. With hard skiving, a process variant is also available that enables hard machining of gears [1]. The kinematic of the skiving process is shown in Fig. 1 and can be described as follows: The tool is tilted relative to the workpiece by the axis cross angle \u03a3. The axis cross angle \u03a3 is the sum of the helix angles of the tool and the workpiece [6]. In the case of internal gears, the value of the tool helix angle \u03b20 is negated. If the axis cross angle \u03a3 is increased, the generating motion becomes a screw rolling motion [2]. Thus, the skiving kinematics is similar to the kinematics of two helical gears meshing [6]. The screw movement results in a cutting velocity vc, which leads to chip removal if the axis cross angle \u03a3 is large enough" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000338_j.1538-7305.1960.tb01602.x-Figure20-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000338_j.1538-7305.1960.tb01602.x-Figure20-1.png", "caption": "Fig. 20 - A reversed segment in a magnetic tape.", "texts": [ " As an example, the successive outputs obtained as the bit pattern was propagated through the successive odd-numbered rungs, 7,9,11, 13, 15,17 and 19 respectively, are shown in Fig. 19(d). The second method is best introduced by considering the example of a magnetic tape, which has an easy direction of magnetization parallel to its long axis. Let it be fully magnetized in this direction, say from right to left. Now let a field be applied in a direction to reverse the magnetiza tion in a small segment, ab, as illustrated in Fig. 20; for example, by means of a small solenoid. In this case, flux closures must be completed by air-flux (not shown). Under these conditions, the segment ab is subjected to demagnetizing fields, which increase as the length ab is decreased, and, in fact, if cb is too small the reversed setting will be unstable once the applied field is removed. In consequence, it will be possible to produce a stable zone of reversed magnetization only if ab exceeds a critical length, which is dependent upon the thickness, width and magnetic properties of the tape" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002234_coginfocom50765.2020.9237841-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002234_coginfocom50765.2020.9237841-Figure3-1.png", "caption": "Fig. 3. Using Xbee chips for drone hacking. [39]", "texts": [], "surrounding_texts": [ "\u2022 Jammer \u2212 Its role is to eliminate the communication between the remote control and the drone, can selectively obstruct the communication channels, which covers the control (pilots) signal. It can also disrupt navigation, the drone can lose the GPS signal, get lost from the planned route, or be forced to land. More advanced drones use GPS receivers within the navigation and control loop which allows for some additional features: position hold, return to home, and autonomous flight.\n\u2022 Anti-drone gun \u2212 It's a gun-like device that works by jamming the signal between drone and pilot, in such a way grounding unwanted UAVs. Also, it controls the drone to land safely on the spot or returns to its point of origin. This device can disable drones operating on 5 frequencies and has a range of up to 2500 meters. RF output power is up to 60 W.\n\u2022 Spoofer \u2212 The GPS spoofing attack is an act of fooling the control system of the drone by sending it fake geographical coordinates. In this way, it is possible to hijack the vehicle taking it away from the authentic path. A similar attack is possible due to the leakage of GPS encryption rules, which is a common occurrence for civilian equipment.\n\u2022 Hacking \u2212 Certain software exploit toolkits are focused on hijacking drones targeting weak authentication and encryption methods and allowing hackers to send commands to the victim\u2019s drone.\n\u2022 Snaggers (catchers) \u2212 a net placed under a drone, shot from an air cannon or net-launching gun.\n\u2022 Drone catcher \u2212 There are situations on which destroying drones might not be suitable. For example, when a drone is being used in restricted areas and the task is to land it without losing potentially accusing evidence. In this case, a special drone-catcher device becomes useful. This device is basically an anti-drone drone, capable of catching an enemy UAV in a net from up to 20 meters away. Characteristics of this device are: controlled drone interception, equipped with gimbaled netgun, automatic drone detection, the target is selected by the operator, automatic tracking without operator input, the unwanted drone is carried to a safe place and brought down by a parachute.\nDestroyers are a special form of defenders which in the physical sense eliminate drones:\n\u2022 Laser \u2212 An optical device that emits a strong beam of light or radio energy at the drone. Relating to the power, the laser can burn out or blind the hardware or camera. For example, ATHENA (Advanced Test High Energy Asset) uses a specially developed tower with 30 kW output power [35].\n\u2022 Directed electromagnetic pulse \u2212 A generator emits a pulse of energy that, if strong enough, can damage weakly shielded electronics.\n\u2022 High-energy microwave \u2212 The antenna generates microwaves and if a high-energy microwave hits a drone, it will be destroyed.\nVI. JAMMERS\nJammers or commonly called signal blockers are devices used to block and interfere with authorized radio communication signals. Jamming equipment interferes with the remote control circuits and navigation signals of drones and forces them to land or return. The drone signal jammer blocks the communications between the enemy drone\u2019s video, telemetry, command, control and navigation systems and the operator [36]. Jammers are designed to disable a drone by blasting electromagnetic noise at active radio frequencies, and at a power high enough to eliminate any communication between the drone and its operator [37].\nThe civilian drone jammers include two bands: Band 1 - 2.4 GHz and Band 2 - 433 MHz. The RF circuit in them consists of an electronic oscillator whose frequency is controlled by a voltage (VCO- voltage controlled oscillator) and of a linear power amplifier whose function is to amplify the output power of the VCO. The tuning circuit consists mainly of a triangular wave generator to generate the tuning signal along with a noise generator to tune the VCO to the required RF jamming signal. The power supply unit is designed to generate appropriate DC voltages for the RF and the tuning circuits.\nFrequency jamming is a method that is even more subtle than the physical interception, and the anti-UAV defense system (AUDS) is one such solution. This system scans the space on drones and jams their control signals using their own high-powered radio signal.\nAn anti-drone rifle that uses targeted radio signals to disrupt drone controls (similarly to AUDS) ensures a more portable option. It currently has a range up to 1500 meters but may be able to reach even farther in future versions. The standard output power of these devices with VCO technology is up to 75 W and with high gain 12 dBi directional antenna.\nVII. DRONE HACKING\nUsing certain communication weaknesses, various forms of drone hacking are possible:\n1. A security researcher named Sasi infected his drone with Maldrone malware. This malware creates a persistent backdoor, enabling an attacker to remotely control the infected drone, cause it to drop from the sky, or hijack it to conduct surveillance [38]. Main steps that have been taken: normal flight controlled using API phone/computer, enabling autopilot, sending code and waiting for a reverse TCP connection from the drone, overriding the drone\u2019s autopilot.\n2. Some gadgets exploit DMSx, often used radio protocol by most remote-controlled commercial drones, allowing the hijacker to take control (for instance, the device called Icarus). The exchanged public key (not encrypted) can be easily reconstructed after completing the binding process by observing the protocol and using specific brute-force methods. Furthermore, there is a timing attack vulnerability wherein the\n000022\nAuthorized licensed use limited to: Raytheon Technologies Corporation. Downloaded on December 19,2020 at 07:38:28 UTC from IEEE Xplore. Restrictions apply.", "synchronization to the target transmission is done and transmission of offensive control packets in front of the target, the receiver accepts information and rejects the targets.\n3. There are devices featuring Xbee chips through which the drone receives the commands. They specifically implement a telemetry module featuring an Xbee radio chip. This module modifies the Wi-Fi commands that a flight planning software sent into low-frequency radio signals, to be further transmitted to a different Xbee chip located on the drone itself. The operator will therefore be able to control the drone even when considerably removed, further than otherwise possible.\nVIII. SPOOFING ATTACK\nWhen spoofing a certain receiver, signals from numerous satellites must be reproduced by the potential attacker, followed by the transmission of the spoofing signal with the aim of capturing the target GPS receiver. If the original satellite signals and spoofed signals cannot be distinguished by the targeted receiver, the spoofing has succeeded, the receiver is fooled into seemingly appear in another position or another point of time [40]. The final result of this attack is changed (spoofed) path of the drone relative to the desired plan. The approximate range (radius) of a spoofing device is 2500 m.\nIn order to prevent potential misuses, the regular GPS signal may include an encrypted binary P-code, known as Ycode, which is transmitted on L1 and L2 frequencies. This encrypted code has a frequency of 10.23 MHz, and it does not repeat over the course of an entire week. The P (precise) code is encrypted into the Y (or P(Y)) code when anti-spoofing mode is on. Given that without being in possession of the encryption key values, actually the generation of the Y-code is not possible, therefore, neither is it possible to spoof a receiver device into tracking Y-code.\nApart from other features, the GPS signals contain a C/A (coarse/acquisition) code, whose was supposed to help with acquiring the Y-code, though it is currently applied with all civilian GPS receivers. The C/A code is not encrypted, the binary symbols (logical 1s and 0s) have a frequency of 1.023 MHz, and the code itself (a unique sequence) repeats each millisecond. Since the C/A code structure is openly published in a signal-in-space (SIS) interface specification, a half-way capable attacker could reproduce it and then generate a spoofed version of the GPS signal.\nSpoofing attacks countermeasures \u2013 A reliable way of warding off spoofing is by means of directly tracking the encrypted Y-code. However, the precondition of such type of protection is that a GPS receiver be equipped with a Selective Availability Anti-Spoofing Module (SAASM). It is impossible for the SAASM receivers to track Y-code unless they are loaded with the functioning decryption key of that moment and the modules have protection so as to avoid attackers\u2019 reverse engineering. Only government authorized customers have access to such receivers, the purchases and distribution of such devises are closely monitored.\nAnti-spoofing countermeasures implemented in software on receivers can be grouped into two categories: amplitude and time-of-arrival discrimination.\nSuperior defense techniques include the following:\n\u2022 Consistency of navigation - inertial measurement unit (IMU-outputs gyroscopes, accelerometers and magnetometers raw data) cross-check\n\u2022 Polarization and angle-of-arrival discrimination\n\u2022 Cryptographic authentication\nThe basic prevention against GPS attacks is the use of cryptographic techniques, as well as the use of authentication processes (WAAS (Wide Area Augmentation System) message authentication) between transmitters and receivers avoiding interferences of external sources. These techniques are implemented only in the military sector and aviation.\nAs an efficient protection against spoofers, civilian users can rely on numerous receivers with the ability to track several GNSS (for instance GPS, GLONASS, Galileo, and BeiDou simultaneously), given that it would be necessary for the attacker to generate and transmit all possible GNSS signals simultaneously to spoof the target.\nIX. GPS SIGNAL JAMMING\nSignals stemming from the satellites are significantly weakened so that the elimination of GPS coverage for a particular territory is possible by a 10 W jammer (or less) [33]. A weak received signal power means that CDMA (Code Division Multiple Access) signals (not only GNSS), have become vulnerable to either unintended or deliberate interference (jamming). In case the level of interference surpasses a specific threshold, the GNSS signal is lost within the interfering signal.\nApplying jamming methods enables the interruption of GPS signal reception. In such a case, the drone may lose the ability to follow the route and calculate where it is, how high it is and where it is flying to. Because the communication itself is also jammed, the operators lose their control over the drone, allowing the UAV to fly uncontrolled up to the point when it smashes into the ground, makes a landing based on an emergency procedure, or simply runs out of fuel. The estimated range (radius) of such jamming equipment is 1500 m.\nA solution to these attacks is the production of equipment that has spoofing or jamming detectors and uses encrypted GPS signals (like the military does).\nAnother solution against jamming is using backup chips that contain gyroscopes, accelerometers, and a master clock - Time and Inertial Measurement Unit (TIMU). It is able to produce a record of vehicle movements, maintaining\n000023\nAuthorized licensed use limited to: Raytheon Technologies Corporation. Downloaded on December 19,2020 at 07:38:28 UTC from IEEE Xplore. Restrictions apply.", "coordinates accurate until the GPS connection is established again.\nProtection against jamming - There are several main strategies to help overcoming interference:\n1. Filtering the incoming traffic\n2. Receiver equipped with IMU\n3. Elimination of interfering signals using an adaptive antenna array.\nX. COGNITIVE APPROACH IN DRONE DETECTION TECHNOLOGY\nThe primary role in the successful defense of drones is their detection. In this sense, the adequate radar system in combination with an optical camera is effective and often used in practice. Cognitive technologies have a significant place in this area. The special form of radars, the so-called cognitive radars, run on the basis of the perception-action cycle of cognition with the ability of sensing the environment, learning from vital information about both target and background, while also adjusting the radar sensor to best meet the given mission\u2019s requirements dictated by the goal to be accomplished [41]. The notion of cognitive radar was initially exclusively implemented for active radar.\nCognitive radar (CR) \u201ccontinuously learns about the environment through experience gained from interactions with the environment, the transmitter adjusts its illumination of the environment in an intelligent manner, the whole radar system constitutes a dynamic closed feedback loop encompassing the transmitter, environment, and receiver\u201d [42].\nThe main characteristic of cognitive radar that makes it stand out from the standard radar is the use of active feedback between receiver and transmitter, which is presented in Fig. 4. The classical principle of adaptivity is, in this way, extended to the transmitter.\nThe key features of cognitive radar include the following [43]:\n1. it makes informed decisions based on theoretic approach;\n2. it has passive environmental and radar sensors; 3. is is equipped with learning algorithms to make the\nperformance better and adjust to novel environmental situations;\n4. it includes a knowledge database full of environmental, targets and other a priori information;\n5. the waveform solution space, for known targets; 6. receiver-to-transmitter feedback to mitigate\nclutter/interference and maximize target information. Cognitive radars are generally divided into active and passive:\n\u2022 Active radars adjust to the environment - work in spectrally dense conditions and quickly modify the waveform of the transmitted signals to avoid interference with the original user of the channel.\n\u2022 Active radars adjust transmit signal parameters to achieve a defined level of target tracking performance.\n\u2022 Passive radars are incapable of directly altering the transmitted waveforms, however, they can pinpoint the best option to make both the detection and tracking better\nThere are different advantages important for drone detection that could be attributed to CRs when compared to traditional active radar (TAR) with similar antenna performance [45]:\n1. Increased detection range. 2. Reduced time for parameter acquisition in target\ntracking and faster information delivery 3. Enhanced information accuracy of the target position. 4. Reduced risk in selecting a decision-making type in\nthe transmitter in conditions of real-time environmental uncertainties and disturbances.\n5. Better detection of smaller targets (such as drones and stealth flying objects).\n6. Increased capabilities of passive radar and multistatic radar systems in detection of stealth aircrafts better than conventional monostatic radars. Older stealth technologies reflected energy away from the line of sight, increasing radar cross-sections (RCS), which multistatic passive radars can detect.\n7. Use of waveforms difficult to be detected, including wideband signals.\n8. Lower probability of being tricked by fake target energy (including decoys).\n9. Use of built-in protection against misdirecting Direction-of-Arrival (DOA) measurements.\n10. More precise geolocalization due to networked CRs. 11. Enabling communication when other devices are\njammed.\n000024\nAuthorized licensed use limited to: Raytheon Technologies Corporation. Downloaded on December 19,2020 at 07:38:28 UTC from IEEE Xplore. Restrictions apply." ] }, { "image_filename": "designv11_5_0000557_aim.2016.7576989-Figure21-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000557_aim.2016.7576989-Figure21-1.png", "caption": "Figure 21. Conduit immediately after the experiment", "texts": [ " This is due to the smaller expansion change of the cylindrical tube as can be confirmed from Fig. 15. Fig. 20 shows the conveyance amount per unit of time as function motion interval obtained by the least squares fitting of the data presented in Fig. 19. We find that for the motion interval values of 1.5, 1.0, 0.5, 0.3, and 0.1 s, conveyance rates of 1.5, 1.9, 4.4, 1.7, and 0.0 g / s, respectively, are obtained. This shows a carrier conveyance amount close to the results described in previous section. The photograph of the conduit immediately after the experiment is presented in Fig. 21, showing that little carrier is found in the conduit. Increasing the amount of the carrier in the conduit leads to improvement in the conveyance amount. We will investigate the construction of a system for supplying a sufficient amount of powder in future work. As described in the previous section, it was found that an applied pressure of 40 kPa and motion interval of 0.5 s gave the highest conveyance amount. We therefore used these parameters to examine conveyance of the developer used in the actual printer by peristaltic conveyor in order to calculate the conveyance amount per unit time" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002210_j.robot.2020.103670-Figure19-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002210_j.robot.2020.103670-Figure19-1.png", "caption": "Fig. 19. Distance between 0\u03a3B and \u03a3B .", "texts": [ " When the robot center of mass shifts behind the contact point etween its middle wheels and the ground, the angles of the links f the manipulators, 0\u03c61, 0\u03c612, and 0\u03c6123, shift to \u03c61, \u03c612, and \u03c6123, respectively. Thus, the position vector of hand rim of the wheelchair, Bp4 = [X4 Y4] T in \u03a3B, is as given below. Bp4 : [ X4 Y4 ] = [ cos\u03c61 \u2212 sin\u03c61 sin\u03c61 cos\u03c61 ][ lLB hLB ] + l2 [ cos\u03c612 sin\u03c612 ] + l4C [ cos\u03c6123 sin\u03c6123 ] + [ 0 RB ] (3) he basic coordinate system 0\u03a3B moves to \u03a3B in Stage 3. \u2206x is the orizontal movement of coordination as the robot\u2019s angle inclines the distance between 0\u03a3B and \u03a3B, Fig. 19). The wheelchair emains stopped until the system finishes lifting the robot\u2019s front heels. p4 does not change. Thus, x = \u2212(X4 \u2212 0X4) (4) ubstituting (2) and (3) for (4), x = l4C (cos 0\u03c6123 \u2212 cos\u03c6123) + l2(cos 0\u03c612 \u2212 cos\u03c612) +lLB(1 \u2212 cos\u03c61) + hLB sin\u03c61 (5) The position vector of \u03a3B in 0\u03a3B is described below (Fig. 19). p\u2206x = [\u2206x 0]T (6) 4.2. Robot front wheel height needed to climb a step The angle of the robot necessary to place the front wheels on a step is clarified in Stage 3. In Fig. 20, q1 is the position of the robot\u2019s middle wheel axes (Here, q1 = p1, Fig. 18), q2 is the position of the robot\u2019s front wheel axes, and q3 is the tread of the robot\u2019s front wheels. \u03a3I and \u03a3II are the local coordinate system. (Here, \u03a3I = \u03a31, Fig. 18.) \u03c11 and \u03c12 are the angles of \u03a3I and \u03a3II formed by \u03a3B and \u03a3I , respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001066_1.a34565-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001066_1.a34565-Figure2-1.png", "caption": "Fig. 2 Interlayer friction prevents blossoming.", "texts": [ " The photograph also shows acrylic guards installed on each side of the hub. These guards ensure that the boom does not slide into the springs, and they make it easier to keep the boom straight as it is rolled onto the hub. The mechanism relies on interlayer friction to prevent slip that would lead to blossoming. The sprung roller creates a localized normal force Fn in one location on the coiled boom, thereby increasing the force of friction Ff between adjacent layers in the coiled boom. This is shown in Fig. 2: Fn \u2212kx (1) Ff \u03bcsFn (2) If the boom is used to hold an instrument away from the spacecraft, the instrument can be integrated into themechanism. If there is no tipmounted instrument, the mechanism can be retained or jettisoned, depending on debris requirements. If the mechanism is jettisoned, an undamped mechanism may be sufficient. The opening shock will be minimal because only a small portion of the boom is traveling at the end of deployment. The collapsible tube mast (CTM) boom was selected for use with this mechanism" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001836_j.jfranklin.2020.03.017-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001836_j.jfranklin.2020.03.017-Figure1-1.png", "caption": "Fig. 1. Prescribed funnel set F \u03c8 \u03b5 .", "texts": [ " Moreover, Property 1 and Assumption 1 imply that system (1) allows more serious nonparametric uncertainties than those of [21,22] since the bounds of general mass matrix and disturbance are unknown rather than known. For this, certain compensation mechanism for the serious nonparametric unknowns should be developed, under which to develop a novel control design scheme. 2.2. Control objective For a given reference trajectory y r \u2208 R n which is continuously differentiable and satisfies Assumption 2 below, we define the tracking error e = q \u2212 y r and a funnel set (see Fig. 1 above) F \u03c8 \u03b5 = { (t, e ) \u2208 R + \u00d7 R \u2223\u2223\u03c8 \u03b5 (t ) \u2016 e \u2016 < 1 } , (2) where \u03c8 \u03b5 \u2208 S \u03b5 is defined as S \u03b5 = { \u03c8 \u03b5 (t ) \u2223\u2223\u2223\u03c8 \u03b5 (t 0 ) = 0;\u03c8 \u03b5 (t ) > 0, \u2200 t > t 0 ; lim inf t\u2192 + \u221e \u03c8 \u03b5 (t ) \u2265 2 \u03b5 ; \u2223\u2223 \u02d9 \u03c8 \u03b5 (t ) \u2223\u2223 \u2264 c(1 + \u03c8 \u03b5 (t )) , \u2200 t \u2265 t 0 ; sup t\u2265t 0 \u03c8 \u03b5 (t ) < + \u221e } , (3) with t 0 being initial time, \u03b5 (called tracking accuracy parameter) and c being known positive constants. Assumption 2. There exists an unknown constant y\u0304 r such that \u2016 y r \u2016 \u2264 y\u0304 r and \u2016 \u0307 yr \u2016 \u2264 y\u0304 r . Please cite this article as: J" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001619_tcyb.2020.3034697-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001619_tcyb.2020.3034697-Figure1-1.png", "caption": "Fig. 1. Two-mass\u2013spring system as an agent.", "texts": [ " However, it is difficult to establish a clear relationship between kmi and the mean-square consensus error, though some discussion will be given in the simulation example. V. EXAMPLES In this section, two examples are presented to illustrate the effectiveness of the proposed control strategy. Example 1: In practice, a large class of mechanical systems can be modeled by mass\u2013spring systems [35]. Consider a multiagent system consisting of four two-mass\u2013spring agents. The agent structure is shown in Fig. 1, which consists of two masses that are free to slide on a frictionless surface. The agent is controlled by a force ui(t), which acts on mass M1i. The displacement of mass M1i is regarded as the output of the agent. The displacement of mass M2i is unmeasurable, and the spring constants k1i and k2i and the damping constants b1i and b2i are unknown. The dynamics of each two-mass\u2013spring agent is described by x\u0307i(t) = \u23a1 \u23a3 0 1 \u2212 k1i M1i \u2212 b1i M1i \u23a4 \u23a6xi(t)+ \u23a1 \u23a3 0 1 M1i \u23a4 \u23a6ui(t) + \u23a1 \u23a3 0 1 M1i \u23a4 \u23a6 [ k1i b1i ] qi(t) q\u0307i(t) = \u23a1 \u23a3 0 1 \u2212k1i + k2i M2i \u2212b1i + b2i M2i \u23a4 \u23a6qi(t) + \u23a1 \u23a3 0 0 k1i M2i b1i M2i \u23a4 \u23a6xi(t) yi(t) = [1 0 ] xi(t) (49) where xi(t) = [y1i(t), y\u03071i(t)]T is the state of the model part of agent i, qi(t) = [y2i(t), y\u03072i(t)]T is the unmeasurable state of the unmodeled dynamics, y1i(t) and y\u03071i(t) are the position and the velocity of mass M1i for agent i, respectively, and y2i(t) and y\u03072i(t) are the position and the velocity of mass M2i for agent i, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000586_j.ifacol.2016.10.501-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000586_j.ifacol.2016.10.501-Figure2-1.png", "caption": "Fig. 2. Course command near a course changing point", "texts": [ " However, near the turning point, the fuzzy system will decide to choose the appropriate course defined by the next two WPs as following equation: CDH I *)( 121 (2) where, I is order of course change, 1 is course of the shortest path to the next WP, 2 is course of the shortest path to the second next WP and CDH is the reference degree to the second next WP ( 10 CDH ), calculated by fuzzy controller. 2016 IFAC CAMS Sept 13-16, 2016. Trondheim, Norway 606 Yaseen Adnan Ahmed et al. / IFAC-PapersOnLine 49-23 (2016) 604\u2013609 Fig. 2 shows the course changing command near a course changing point (WP). In this research, to judge the nearness of the waypoint, TCPA (time to closest point of approach) and DCPA (distance of the closest point of approach) are used for fuzzy reasoning. Fig. 3 shows the bearing relationship between the ship and waypoint. According to the figure, the distance between the ship and nearest waypoint is calculated as follows: 22 )()( YtYoXtXoD (3) Then, the following calculations are done to get the bearing angle of waypoint from the ship" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002269_lra.2020.3039732-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002269_lra.2020.3039732-Figure3-1.png", "caption": "Fig. 3. Surgical FOV control based on tracking motion without RCM constraints.", "texts": [ " The tracking motion outside the nasal cavity is not restricted by RCM, but for the tracking motion inside the nasal cavity, in order to protect the nostril from being pulled, the motion of the endoscope is restricted by RCM. Therefore, two kinds of tracking motion control are analyzed separately in this section and then combined with the transition process from outside the nasal cavity to inside the nasal cavity, to establish the overall framework of tracking motion control for the entire surgical process. Finally, to improve the safety of the robot motion, VFs constraints and contact force control are performed on the tracking motion process. Fig. 3 shows the schematic diagram of surgical FOV control based on tracking motion without RCM constraints. First, through the navigation system, the poses of optical markers 1 and 2 can be obtained as NDI ndi1 T and NDI ndi2 T (see Fig. 2). Then the poses of the endoscope and instrument tips in Fig. 3 can be calculated as NDI End T and NDI Ins T (see Fig. 3). Since OFocus is the camera focal point of the endoscope, the coordinate system at the focal point can be expressed in the coordinate system of the endoscope tip. NDI FocusT = NDI End T \u00b7 \u23a1 \u23a2\u23a2\u23a2\u23a3 1 0 0 0 0 1 0 0 0 0 1 R 0 0 0 1 \u23a4 \u23a5\u23a5\u23a5\u23a6 (1) We keep the orientation of End FocusT at the focal point unchanged and move its origin position OFocus to the instrument tip OIns (NDI Ins T (1 : 3, 4) is the origin point of NDI Ins T ) NDI NewFocusT = [ NDI FocusR NDI Ins T (1 : 3, 4) 0 1 ] (2) Then, we keep the orientation of the coordinate system unchanged and move the origin of the obtained coordinate system to point O1 End; then we obtain: NDI O1 End T = NDI NewFocusT \u00b7 \u23a1 \u23a2\u23a2\u23a2\u23a3 1 0 0 0 0 1 0 0 0 0 1 \u2212R 0 0 0 1 \u23a4 \u23a5\u23a5\u23a5\u23a6 (3) At this time, the Z-axis vector of NDI O1 End T and the Z-axis vector of NDI Ins T can be extracted as follows: NDIzO1 End = NDI O1 End T (1 : 3, 4) (4) NDIzIns = NDI Ins T (1 : 3, 4) (5) Based on the two extracted Z-axis vectors, the current spatial angle \u03b81 between the endoscope and the instrument and the rotation axis NDIzaxis(kx, ky, kz) from the endoscope to the instrument can be calculated as follows: \u03b81 = acos \u239b \u239d NDIzIns \u00b7 NDIzO1 End \u2016NDIzIns\u2016 \u00b7 \u2225\u2225\u2225NDIzO1 End \u2225\u2225\u2225 \u239e \u23a0 (6) NDIzaxis = NDIzIns \u00d7 NDIzO1 End \u2016NDIzIns\u2016 \u00b7 \u2225\u2225\u2225NDIzO1 End \u2225\u2225\u2225 (7) According to the space required between the surgeon\u2019s hand and the end-effector of the robot, we can set the desired tracking angle \u03b82, and then the desired endoscope pose can be obtained as: NDI O2 End T = RK(\u03b8) \u00b7 NDI O1 End T (8) RK(\u03b8) = \u23a1 \u23a2\u23a2\u23a2\u23a3 kxkxv\u03b8 + c\u03b8 kxkyv\u03b8 \u2212 kzs\u03b8 kxkzv\u03b8 + kys\u03b8 0 kxkyv\u03b8 + kzs\u03b8 kykyv\u03b8 + c\u03b8 kykzv\u03b8 \u2212 kxs\u03b8 0 kxkzv\u03b8 \u2212 kys\u03b8 kykzv\u03b8 + kxs\u03b8 kzkzv\u03b8 + c\u03b8 0 0 0 0 1 \u23a4 \u23a5\u23a5\u23a5\u23a6 (9) where \u03b8 = \u03b82 \u2212 \u03b81,c\u03b8 = cos\u03b8,s\u03b8 = sin\u03b8,v\u03b8 = 1\u2212 cos\u03b8, and RK(\u03b8) is the equivalent axis angle rotation matrix [38]" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000723_00207721.2019.1567865-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000723_00207721.2019.1567865-Figure4-1.png", "caption": "Figure 4. The topology of multi-agent system with six agents.", "texts": [ " Simulation illustrates that with \u03c4 = 0.35, each agent of the system is reaching consensus, but the states are diverging with \u03c4 = 0.38. The time trajectory of each agent of system (4) is shown in Figure 3. Furthermore, with Equations (16) and (17), we can obtain the value \u03c9\u2217 = 4.2505 and \u03c4 \u2217 = 0.3608 exactly with the same parameters mentioned above. Hence, it confirms that system (4) is consensusable as \u03c4 \u2208 (0, \u03c4 \u2217). Example 5.2: Consider a multi-agent system consisting of six agents, with digraph topology as shown in Figure 4. The Laplacian matrix of the topology L and weighting matrix D\u22121L are L = \u239b \u239c\u239c\u239c\u239c\u239c\u239c\u239d 2 \u22121 0 0 \u22121 0 0 1 \u22121 0 0 0 0 0 2 \u22121 \u22121 0 0 0 0 1 \u22121 0 \u22121 0 \u22121 0 3 \u22121 \u22121 0 0 0 0 1 \u239e \u239f\u239f\u239f\u239f\u239f\u239f\u23a0 , Figure 3. (a) The time trajectory of system (4) as \u03c4 = 0.35. (b) The time trajectory of system (4) as \u03c4 = 0.38. D\u22121L = \u239b \u239c\u239c\u239c\u239c\u239c\u239c\u239d 1 \u22120.5 0 0 1 \u22121 0 0 1 0 0 0 \u22120.3333 0 \u22120.3333 \u22121 0 0 0 \u22120.5 0 0 0 0 \u22120.5 \u22120.5 0 1 \u22121 0 0 1 \u22120.3333 0 0 1 \u239e \u239f\u239f\u239f\u239f\u239f\u239f\u23a0 , where D\u22121 is diag{0.5, 1, 0.5, 1, 0.3333, 1}. By simple calculation, the eigenvalues ofD\u22121L noted as \u03bb1 = 0, \u03bb2,3 = 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001574_j.compstruct.2020.113204-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001574_j.compstruct.2020.113204-Figure1-1.png", "caption": "Figure 1. Schematic view of bending-dominated periodic lattice (a) schematic of the lattice and (b) schematic of the unit cell.", "texts": [], "surrounding_texts": [ "structural failure caused by the bending deformation of slender parts when they are subjected to compression force. Several definitions of the critical buckling load can be found in the literature. In [22], the buckling load is considered as the maximum value of force in the force-displacement curve. However, Russel et al [23] defined the first non-uniformity as buckling load in the stress-strain curve. In this paper, we adopt the former definition, i.e. the highest compressive load.\nElastic buckling of BDPLs can serve as an important mechanism in designing\nlattice metamaterials. In recent years, pattern transformation, due to in-plane buckling of lattice structures have received lots of attention. Buckling-based pattern transformation has been used to achieve many programmable mechanical properties such as negative Poisson\u2019s ratio [24-26] and so on [27-29]. These structures usually buckle elastically before yield or fracture when they are subjected to compression exceeding some critical value. It is generally caused by local elastic instability and is therefore reversible and repeatable [24, 30, 31]. For example, buckling of struts triggers a dramatic transformation of circular holes in a flexible matrix into a periodic pattern of alternating and mutually orthogonal ellipses [32-34]. However, if out-of-plane buckling takes place in advance, the lattice material cannot show pattern transformation and corresponding programmable properties. Therefore, it is very helpful to systematically investigate buckling behaviors of BDPLs, in order to broaden their usages in designing potential extraordinary properties.\nThe paper is structured as follows. In Section 2, the settings of finite element\nsimulation and experiment used to study the buckling behaviors are briefly represented. In Section 3, based on a series of simulation results and experimental data, we analyze the influences of relative density, slenderness ratio, cell numbers and central angle on the instability performance of BDPLs. The possibilities of adjusting both critical buckling load and buckling modes are demonstrated. It opens up a broader design space for achieving particular effective properties by adjusting lattice geometries.", "The schematic diagram of the bending-dominated periodic lattice considered in\nwhere N is the compressive load in y direction per unit width, and E represents the Young\u2019s modulus of the matrix material that the BDPL is made of. The matrix material", "is thermoplastic polyurethanes (TPU), whose Young\u2019s modulus and Poisson\u2019s ratio are 30MPa and 0.45, respectively.\nSolidWorks software (Ver. 2016) has been used to establish the lattice geometry.\nIn this study, we adopted the 5\u00d75 cells to present our results unless otherwise" ] }, { "image_filename": "designv11_5_0000734_b978-0-12-814062-8.00012-1-Figure10.9-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000734_b978-0-12-814062-8.00012-1-Figure10.9-1.png", "caption": "Figure 10.9 Samples obtained using a wire as a substrate for deposition: (A) deposition along substrate wires and (B) deposition across substrate wires.", "texts": [ " A Ti 6Al 4V wire with diameter of 3 mm was used as consumable material for deposition and CP Ti wire with diameter of 3.2 mm was used as the substrate in this experiment accordingly. Setup process parameters were the following: power of the electron beam gun\u20143 kW at an accelerating voltage of 15 kV, the wire feed rate 14 mm/s, and the substrate translation speed 14 mm/s. Preliminary heating of the substrate was not applied. In one experiment, the single-bead and the doublebead walls were built along the substrate wires (Fig. 10.9A). In another experiment, two single-bead walls were built across substrate wires and also one cylinder was built upon the same substrate wires (see Fig. 10.9B). These figures (walls and cylinder) have similar main dimensions (width and height of the walls, layer thickness) and the general view of side surfaces with the same figures built upon the thick solid substrate under the same process parameters. But, in this case, the base wiresubstrate looks practically straight. It means formation of very low residual stresses and distortions in the substrate that is almost unavoidable when using a massive plate [14]. The application of such an approach can be very effective in cases of production of 3D objects in which the base plate is not a part of the final product and must be completely removed by machining" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002356_978-3-030-50460-1-Figure24.8-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002356_978-3-030-50460-1-Figure24.8-1.png", "caption": "Fig. 24.8 Distribution of damage measure \u03c9 over the sample volume for load q = 100MPa.", "texts": [ " The material functions of the SCC models, the values of which were used to obtain the results presented below, were obtained on the basis of experimental data on the dependence of the fracture time on the stress level and a number of assumptions 424 Gorokhov, Zhegalov, Kazakov, Kapustin, Churilov Fig. 24.7 Variant of FE discretization of a sample fragment in a spatial formulation. about the duration of the various stages of the SCC process, which were further refined in the process of numerical simulation. The distribution pattern of the \u03c9 damage measure over the sample volume for a moment of time close to the moment of failure for the load case q = 100 MPa is shown in Fig. 24.8 A similar picture of the evolution of SCC processes was observed for other load cases. Based on the analysis of the results of numerical studies, it was found that the destruction of the sample for all the considered load cases occurs in the region of the interface between the corrosive medium and the air. The nature of the zones destroyed as a result of the SCC and the sequence of their development in time for all considered load cases qualitatively coincide. Moreover, the calculation results obtained on various types of grids within each load case are in good agreement with each other both qualitatively and quantitatively" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001403_j.ymssp.2020.107075-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001403_j.ymssp.2020.107075-Figure5-1.png", "caption": "Fig. 5. Interaction between the ball and the cage.", "texts": [ " 4 shows the contact condition of the ball and the inner-right raceway. From this figure, the reverse extending line of rrbcr and the circle of raceway outline are intersected atm. The vector from the origin of the ring fixed frame to pointm can be expressed as rrrm \u00bc rrbr D 2 rrbcr k rrbcr k \u00f09\u00de thus, the contact condition of the ball and the inner-right raceway can be written as dir > 0 rrrm1 < 0 \u00f010\u00de where rrrm1 is the component of rrrm on xr axis. The geometrical relationship between the ball and the cage is shown in Fig. 5, define the cage fixed frame Ocaxcaycazca and the cage pocket frame Ocpxcpycpzcp. In the cage pocket frame, the contact deformation can be given as dcpb \u00bc Dcpb ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi rcpcpb1 2 \u00fe rcpcpb2 2 r \u00f011\u00de where Dcpb is the clearance of the cage pocket, rcpcpb1, r cp cpb2 are the components of the vector from the geometric center of the cage pocket to the ball center on the xcp axis and the ycp axis in the pocket frame" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000489_s00170-016-9165-4-Figure9-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000489_s00170-016-9165-4-Figure9-1.png", "caption": "Fig. 9 Contour of the temperature distribution for case 1", "texts": [ "9, and 1.53 W in each case. The friction heat is actually distributed along the ball contact area while the LMB is reciprocally moved, so the thermal load acts as a moving heat source. However, this simulation focuses on a steady-state condition, so it is assumed to be stationary. For steady-state conditions, the temperature distribution of the LM guide system was analyzed using ANSYS Workbench software. The same nine points as in the experiments were used in the FE model to compare the temperature. Figure 9 shows the steadystate temperature distribution of case 1 using the FEM. Figure 10 compares experimental and simulation results for temperature rise for all cases. Among nine points, the maximum temperature appeared at channel 5 or channel 8 positions for all cases, in agreement with the experimental results. In this research, temperature rises were compared at those positions only because these channels had the highest thermal response (Table 5). A maximum error of 13 % was determined for case 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000718_978-1-4471-7278-9-Figure6.10-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000718_978-1-4471-7278-9-Figure6.10-1.png", "caption": "Fig. 6.10 Self-assembling M-Block robots", "texts": [ " The face magnets join cubes together, while the magnets on the corners help keep the cubes in contact as they experience changes in position, like the connected legs of fire ants. The magnets on the faces of the cubes can be turned on or off. Because the robots are cubeshaped, they are stable and relatively easy to stack into multiple configurations, such as a tower, a bridge, or a wall. Cubes can also slide alongside each other for alignment without any external forces. As a result, M-Block can form almost any desired shape.8 (Fig. 6.10) Collective tooling can be found in software development. The Open Source movement has introduced crowdsourcing platforms such as GitHub, Mathworks 7Emspak J, Cubic Robots Build Themselves. http://news.discovery.com/tech/robotics/cubicrobots-build-themselves-131004.htm 8Romanishin J, M-Blocks Modular Robots, YouTube. https://www.youtube.com/watch?v= mOqjFa4RskA Instinctive Computing 111 Central, and Kickstarter, and thousands of sites for sharing, such as Instructables and Pinterest. However, usually collective tooling is more about sharing than purposeful development" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001788_j.ijmecsci.2020.105516-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001788_j.ijmecsci.2020.105516-Figure1-1.png", "caption": "Fig. 1. (colour online) (a) Representation of a body in undeformed and deformed configurations. (b) Geometry of a circular floating ring bearing.", "texts": [ " The lastic motions are described by a vector of generalised coordinates = ( \ud835\udc2a \ud835\uddb3 1 , \u2026 , \ud835\udc2a \ud835\uddb3 \ud835\udc56 , \u2026 , \ud835\udc2a \ud835\uddb3 \ud835\udc56 )\ud835\uddb3 which contains all nodal displacements u i,j and odal rotations \ud835\udf11 i,j : \ud835\uddb3 \ud835\udc56 = ( \ud835\udc62 \ud835\udc56, 1 \ud835\udc62 \ud835\udc56, 2 \ud835\udc62 \ud835\udc56, 3 \ud835\udf11 \ud835\udc56, 1 \ud835\udf11 \ud835\udc56, 2 \ud835\udf11 \ud835\udc56, 3 ) = ( \ud835\udc2e \ud835\uddb3 \ud835\udc56 \ud835\udf4b \ud835\uddb3 \ud835\udc56 ) . (2) An exact formulation of the global motions depends on the selecion of an origin of the reference frame. In the presented work, a floatng frame of reference formulation is used: the reference frame moves ogether with the body but is not attached to it firmly [26\u201328] . The ituation is shown in Fig. 1 a where \ud835\udc31 \ud835\udc35 \u2208 \u211d 3\u00d71 is the positional vector f reference frame \u03a3B relative to base frame \u03a30 , vector \ud835\udc1c \ud835\udc56 \u2208 \u211d 3\u00d71 repesents the time-invariant position of the i th node of the flexible body n the undeformed configuration, and vector \ud835\udc2e \ud835\udc56 \u2208 \u211d 3\u00d71 contains local isplacements of the i th node. The orientation of reference frame \u03a3B is escribed by four Euler parameters which were introduced by Wasfy nd Noor in Ref. [29] . The Euler parameters are arranged in the form of uaternion \ud835\udf3d\ud835\udc35 \u2208 \u211d 4\u00d71 , which satisfies normalisation condition [29] \ud835\uddb3 \ud835\udc35 \ud835\udf3d\ud835\udc35 = 1 ", " The HD forces acting on the joural can be evaluated using the following integrals hd ,\ud835\udc66 = \u2212 \ud835\udc59\u22152 \u222b \u2212 \ud835\udc59\u22152 2 \ud835\udf0b\ud835\udc5f \u222b 0 \ud835\udc5d cos ( \ud835\udc60 \ud835\udc5f ) d \ud835\udc60 d \ud835\udc65, (8) \ud835\udc39 \ud835\udc47 w t s F u t d c t 2 t n F i H t t r d w E i p a i t i \u2200 2 r f d c h p \ud835\udc5d w i fi fl t f \u0394 T fl s o w w m b c a 2 a f i \ud835\udc0c hd ,\ud835\udc67 = \u2212 \ud835\udc59\u22152 \u222b \u2212 \ud835\udc59\u22152 2 \ud835\udf0b\ud835\udc5f \u222b 0 \ud835\udc5d sin ( \ud835\udc60 \ud835\udc5f ) d \ud835\udc60 d \ud835\udc65, (9) hd = \u2212 \ud835\udc59\u22152 \u222b \u2212 \ud835\udc59\u22152 2 \ud835\udf0b\ud835\udc5f \u222b 0 [ \u210e 2 \ud835\udf15\ud835\udc5d \ud835\udf15\ud835\udc60 + ( \ud835\udc62 in \u2212 \ud835\udc62 out )\ud835\udf07 \u210e ] d \ud835\udc60 d \ud835\udc65, (10) here F hd, y , F hd, z are forces acting in the vertical and horizontal direcions, respectively, T hd is a torque resulting from the HD friction, and , x are circumferential and axial coordinates, respectively, see Fig. 1 b. unction \u210e = \u210e ( \ud835\udc60, \ud835\udc65, \ud835\udc61 ) characterises a gap between bearing surfaces, in , u out are surface velocities of the inner and outer surfaces, respecively. If the bearing gap h is small enough, elastic forces due to structural eformations are produced in the bearing. The Greenwood and Tripp ontact model [30] is used here when the bearing gap is smaller then han 5\u22c510 7 m . .3. Hydrodynamic lubrication We assume that the oil film is thin, the oil is an incompressible Newonian fluid with constant viscosity \ud835\udf07 and the surfaces of both joural and bearing shell are smooth", " [25] and where p i , p o are pressures at the nner side and the outer side of the feed hole, respectively, \ud835\udf01o is a coefcient which characterises the pressure drop due to the change of the ow direction at the outer side of the feed hole, \u27e8q \u27e9 is the average flow hrough the feed hole, and \u0394p is the total pressure drop due to inertial orces which reads \ud835\udc5d = \ud835\udf0c [ ?\u0308? \ud835\uddb3 \ud835\udc5f ( \ud835\udc31 \ud835\udc56 \u2212 \ud835\udc31 \ud835\udc5c ) \u2212 1 2 (||\ud835\udec0\ud835\udc5f \u00d7 \ud835\udc31 \ud835\udc56 ||2 \u2212 ||\ud835\udec0\ud835\udc5f \u00d7 \ud835\udc31 \ud835\udc5c ||2 ) + ?\u0307?\ud835\uddb3 \ud835\udc5f ( \ud835\udc31 \ud835\udc5c \u00d7 \ud835\udc31 \ud835\udc56 )] . (14) he first term in Eq. (14) is caused by the lateral acceleration of the oating ring, the second term is caused by the centrifugal force which retrains the radial flow through the feed holes, and the third term depends n the angular acceleration of the floating ring. Vectors x r , x i , x o , \ud835\udec0r , hich appeared in Eq. (14) , are depicted in Fig. 1 b. Numerical solution of Eq. (14) is implemented so that all nodes hich are situated at the outer side of the feed hole are combined into a aster node with pressure p o , and the nodes on the inner side are com- ined into a slave node with pressure p i . Furthermore, the cavitation ondition (12) is applied to both the master and the slave node. This pproach is described in detail in Refs. [23,25] . .4. Computation strategy Eq. (1) is a set of nonlinear second-order differential equations, and s such it has to be transformed into a first-order system" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002356_978-3-030-50460-1-Figure17.9-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002356_978-3-030-50460-1-Figure17.9-1.png", "caption": "Fig. 17.9 3D printed specimen after tensile testing; left: design a) with indication of simultaneous rupture locations; right: design b) optimized with the SLD scheme with indication of fail-safe rupture sequence.", "texts": [ " An AlMgSc aluminum alloy is used with approximated mechanical values described in the last section. At the current stage, the analysis does not consider that the mechanical properties for the 3D printed material can be inhomogeneous with possibly included material defects, and can depend on the cross sectional dimension. Each specimen is fixed on the left side of Fig. 17.8 and loaded axially with a uniform traverse speed of 1.0 mm/s on the other side. The applied force raised by the machine for impinging the uniform speed on the specimen is measured in the experiment. In Fig. 17.9 and the graphs of Fig. 17.10, the tensile test results are illustrated. The 17 Multimodal approach for automation of mechanical design 319 rupture areas and sequence of rupture are marked in white and by numeration. The SIMP design a) shows a collapsing failure directly when the first rupture occurs, due to the simultaneous breaking of the two marked beams. On the contrary, the configuration b) optimized by the proposed design scheme shows the intended structural ductility and fail-safe behavior" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001914_j.triboint.2020.106390-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001914_j.triboint.2020.106390-Figure1-1.png", "caption": "Fig. 1. Lubricant performances against scuffing, micro-pitting and pitting damages were evaluated on a) FZG back-to-back gear test rig [27] and b) shows a pair of test spur gears mounted on the rig, which are rotated against each other.", "texts": [ " After each load stage, the gear flanks are inspected visually for scuffing marks and the gears are weighed to determine their weight loss. When cumulative scuffing marks observed from all teeth of the pinion exceed the width of one tooth then the critical load stage is achieved. The test parameters and the evaluation of micro-pitting and pitting results based on the standard test methods are described in Ref. [27]. The FZG back-to-back test gear rig consists of two test gears and two slave gears as shown in Fig. 1. The test gears have a center distance of 91.5 mm and are connected to the respective slave gears by two shafts. A load clutch is in place on one of the shafts that allows the shaft to fix the half of the shaft to the base with the locking pin and the other half is twisted to apply the desired torque load, e.g., by means of a lever and weights. Thus, a static torque is applied between the test gears and the load used to apply the torque is removed after securing the clutch together. The test gears are splash lubricated, and the lubricant temperature is monitored using a temperature controlling unit" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002700_s12541-021-00510-4-Figure32-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002700_s12541-021-00510-4-Figure32-1.png", "caption": "Fig. 32 Schematic diagram of a conjugated straight-line internal gear pump", "texts": [ " Involute internal gear pairs have a characteristic of separability, implying that the center distance between the gear and the ring gear changes within a certain range and the transmission ratio remains constant [13]. In contrast, meshing points of a conjugated straight-line internal gear pair correspond to each other. If the center distance changes, the gear pair fails to mesh effectively. This characteristic leads to a different radial structure design. The involute internal gear pump shown in Fig.\u00a031 can be designed as a floating crescent plate by introducing a compensation structure and dynamically adjusting the center distance. However, the conjugated straightline internal gear pump shown in Fig.\u00a032 can only use fixed crescent plates. This fixed clearance design cannot dynamically adjust the center distance to compensate for the wear of friction pairs. Therefore, this design is more complicated, and the requirement for materials is high. 2. The involute internal meshing gear pair shown in Fig.\u00a031 meshes as a straight line, and the meshing characteristic changes linearly with the gear rotation angle. In addition, the flow characteristic changes with the gear rotation angle in a quadratic relationship. On the contrary, the meshing line of the conjugated straight-line internal gear pair shown in Fig.\u00a032 is a curve. The meshing curve changes nonlinearly with the gear rotation angle. Thus, the trapped oil volume gradually increases dur- 1 3 ing transmission, and the formation of negative pressure between gear pairs is more conducive to stable transmission (Fig.\u00a033). In order to verify the effectiveness of the gear pair design method, the structure of the pump was designed according to previous literatures [14, 15] (Fig.\u00a01). The gear and the ring gear were manufactured by the linear cutting method and then processed by certain heat treatments to improve their anti-friction and lubrication properties" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002614_j.compositesa.2021.106449-Figure15-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002614_j.compositesa.2021.106449-Figure15-1.png", "caption": "Fig. 15. Application of z-offset rule in complex tow-tow intersection case.", "texts": [ " This is named neighbour rule which is used to check the zpositions of each tow point after offsetting half tow thickness from the corresponding intersections. If the adjacent point (top, bottom, left or right) has a higher z-position, then raise the current point to the same height, as shown in Fig. 14c. This can break the ideal triangular tow drop (Fig. 14a) for the regular and well-intersected case so a small correction is added to each point lying directly above a tow edge to preserve the computed tow drop. A complex tow-tow intersection example using the z-offset rule is presented in Fig. 15. In this case, there are three height levels including one, two, and three tow thickness, represented by different colours. The outcome of the z-offset manipulation (including the neighbour rule) for this complex tow drop and overlap is presented in Fig. 16. As the defects have a significant knockdown effect on the mechanical properties as shown in various studies [4\u20137], precise defects representation is critical in predicting the local and global material behaviours. However, the 3D shape of the resin-rich pocket is generally irregular and hard to define a priori, such as for the case with multiple tows intersections as mentioned in Section 3" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000414_1.4947440-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000414_1.4947440-Figure1-1.png", "caption": "FIG. 1. (a) Schematic of experimental setup and surface charge distribution of the liquid metal when the electrodes are powered. The blue line represents outline of the liquid metal in unelectrified state; (b) and (c) real object snapshots of the liquid metal, respectively, when t\u00bc 0 s and t\u00bc 0.04 s as the voltage between the inner and the outer electrodes is 1.7 V. (d) The real object snapshot of the liquid metal when the voltage between the inner and the outer electrodes is 18.78 V; (e)\u2013(g) motion diagrams of the liquid metal corresponding to (b)\u2013(d), respectively. The positive sign \u201c\u00fe\u201d and the negative sign \u201c \u201d represent the graphite anode and the cathode, respectively. The scale bars in (b)\u2013(d) are all 10 cm.", "texts": [ " It is immersed in 1 mol/l NaOH electrolyte solution whose main role is to remove oxide on the liquid metal surface and reduce its kinetic friction. Both the liquid metal and the solution are located between a pair of concentric ring graphite electrodes which is described in detail in the authors\u2019 previous work.19 All the liquid metal, NaOH solution, and the electrodes are placed in a petri dish below which there is a permanent magnet to produce vertical magnetic field. The schematic diagram of the vertical crosssection of the experimental setup is presented in Fig. 1(a). As the liquid metal is placed in the NaOH electrolyte, an electrical double layer (EDL), which is equal to a parallel plate capacitor, will be created at the metal-electrolyte interface. Such a layer isolates the liquid metal from the surrounding electrolyte, and the liquid metal is negatively charged while the electrolyte is positively charged.20 The liquid metal can be regarded as an equipotential body due to its high electric conductivity.21 And there is a uniform charge distribution of the potential difference across the EDL in the case of no external voltage", " The Marangoni flow on the liquid meal/electrolyte interface and the electrocapillary motion of the liquid metal are thus generated20,25 due to surface tension gradient. The driving force of the flow exerted upon the interface can be expressed as23 f eVE kD ; (3) where E is the electric field intensity, e and kD are electric permittivity and the Debye screening length of the electrolyte, respectively. Such electric capillary force is the reason causing various liquid metal patterns. The surface charge distribution and outline variation (from dotted line to solid line) of the liquid metal in an electric field are presented in Fig. 1(a). Figs. 1(b) and 1(c), respectively, display the gear pattern (t\u00bc 0 s) and the fan blade pattern (t\u00bc 0.04 s) induced on the free surface of the liquid metal when the voltage between the inner and the outer electrodes is 1.7 V. Fig. 1(d) shows the glitch pattern on the liquid metal surface when the voltage between the inner and the outer electrodes is 18.78 V. The pattern-forming reasons can be interpreted as follows: The deformation of the liquid metal is mainly influenced by its physical properties such as gravity, viscosity, and surface tension. On the basis of the related research results, the surface tension acts as the dominant factor in determining the deformation and breakup of the liquid metal in an electric field compared to gravity and viscosity", "32\u201334 Meanwhile, the liquid metal surrounding the inner anode has a spreading trend35 due to the oxidative potential imposed on it.36 And an inner ring is thus generated close to the trapezium-shaped parts. Figs. 1(e) and 1(f) present motion diagrams of the liquid metal, which correspond to Figs. 1(b) and 1(c), respectively. As the inner electrode works as cathode, however, the surface tension of the liquid metal near the outer anode is too weak to form gaps on the liquid metal surface. And the generated dense folding pattern shown in Fig. 1(d) is quite different from that in Fig. 1(c). Fig. 1(g) presents clutter direction of the perturbations imposed on liquid metal corresponding to Fig. 1(d). As a magnetic field combined with the electric field is applied on the liquid metal, the latter will be subjected to a perturbative Lorentz force F which can be expressed as37 F \u00bc J B; (4) where J is the electric current density directed from the anode to the cathode, and B is the magnetic induction strength. Assuming the direction of B is normal to the substrate, then F is along the clockwise or counterclockwise direction viewed from the overlook to the circular ring-shaped liquid metal according to the right-hand rule" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001857_j.knosys.2020.105863-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001857_j.knosys.2020.105863-Figure7-1.png", "caption": "Fig. 7. Experimental equipment of the performance degradation of rotating machinery.", "texts": [ " , x\u0303b+1, x\u0303b+2 ] is taken as the input of the welltrained QREDNN to obtain the output value x\u0303b+3 at b + 3 time. By that analogy, the newest testing sample vector[ x\u0303b\u2212n+N , x\u0303b\u2212n+1+N , . . . , x\u0303b\u22121+N ] is input to the well-trained QREDNN to obtain the output value x\u0303b+N at b + N time. Finally, we can get N predicted output values. In this section, the effectiveness of our performance trend prediction method based on QREDNN is verified by the full-life vibration acceleration data of double-row roller bearings collected by University of Cincinnati [32]. Fig. 7 shows the experimental equipment of the vibration acceleration data of RM. There are four ZA-2115 double-row roller bearings made by Rexnord, and they are installed on the rotating shaft of the bearing test bed. The rotating shaft is driven by an AC motor with rub belts, and rotates at a constant speed of 2000 r/min. A radial load of 6000lbs is applied onto the rotating shaft and bearings during the experiment. The vibration acceleration data of the four bearings are collected every 10 min, and each bearing keeps running until a serious failure occurs" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001241_j.ijfatigue.2020.105632-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001241_j.ijfatigue.2020.105632-Figure3-1.png", "caption": "Fig. 3. The structural diagram of one stage of planetary gear train.", "texts": [ " The features of this failure mode were examined in detail by comparing standard chemical compositions and the microtopography of the failed planetary gear and bearing. https://doi.org/10.1016/j.ijfatigue.2020.105632 Received 3 January 2020; Received in revised form 3 March 2020; Accepted 25 March 2020 \u204e Corresponding author. E-mail address: weijing_slmt@163.com (J. Wei). International Journal of Fatigue 136 (2020) 105632 0142-1123/ \u00a9 2020 Elsevier Ltd. All rights reserved. T The structural diagram of one stage of the planetary gear train investigated in this study is shown in Fig. 3. In this planetary gear train, the planet carrier d is the input and the sun gear c is the output. The ring gear a is fixed, the three planetary gears b rotate around the sun gear c under the action of the planet carrier d, and each planetary gear meshes with the sun gear and the ring gear simultaneously. The inner hole of each planetary gear and corresponding bearing outer ring were fitted with an interference fit assembly. The stresses of the interference-fit assembly unit of the planetary gear and bearing outer ring are analyzed in the following section" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001800_978-3-030-40410-9-Figure4.16-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001800_978-3-030-40410-9-Figure4.16-1.png", "caption": "Fig. 4.16 Different orientations for the same position", "texts": [ "44) where Ci denotes the instantaneous position vector and De is the desired position vector. In other words, Eq. (4.44) is the function that has to be minimized as much as possible, and it is just the distance between the end-effector and task point. f = \u221a (xCi \u2212 xt )2 + (yCi \u2212 yt )2 + (zCi \u2212 zt )2. (4.45) where t refers to the task point coordinates given for the inverse kinematics problem. If Eq. (4.45) has been used alone as an objective function, we may get the endeffector in the task point but with many choices of orientations. Figure 4.16 reveals an example of two options of orientation for the same task point, and in manipulators with a high degree of freedom, there can be even an infinite number of orientations. In this section, we shall model the objective function for the inverse kinematics of any robot manipulator. Figure 4.17 shows a schematic diagram for the inverse problem; it is more descriptive to explain the procedure by a set of notes as follows: 3. Forward function contains all the forward kinematics equations of the robot arm, and by substituting the candidate solution into those equations, we can get the overall homogenous transformation matrix by a repeated call for the HTM function" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000672_lra.2018.2890433-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000672_lra.2018.2890433-Figure2-1.png", "caption": "Fig. 2. Construction of Reeds-Shepp path from (0, 0,\u2206\u03b8) to (0, 0, 0)", "texts": [ " Then, Costlb = n\u22121\u2211 i=1 cost(vim\u2217 i , v(i+1)m\u2217 i+1 ) \u2264 n\u22121\u2211 i=1 cost(vimoi , v(i+1)moi+1 ) \u2264 n\u22121\u2211 i=1 costo(si, si+1). The following lemma is a simplification of a more general bounding result proved for the Reeds-Shepp vehicle in [27]. Lemma 5. The length of the Reeds-Shepp path between two oriented points (0, 0,\u2206\u03b8) and (0, 0, 0) is at most 3\u03c1\u2206\u03b8. Proof. Without loss of generality, assume \u2206\u03b8 \u2208 [0, \u03c0]. Consider a feasible path for the Reeds-Shepp vehicle between (0, 0,\u2206\u03b8) and (0, 0, 0) passing through points A,B and C as shown in Fig. 2. The length distrs(\u2206\u03b8) of this path is equal to the sum of length of the four segments OA,AB,BC and CO. From Fig. 2, one can deduce that \u03b1 = \u2206\u03b8 2 . The length of the segment joining points B and C is denoted by BC. distrs(\u2206\u03b8) = \u03c1\u2206\u03b8 + \u03c1\u03b1+BC + \u03c1\u03b1 = \u03c1\u2206\u03b8 + 2\u03c1\u03b1+ \u221a (\u03c1 sin \u2206\u03b8)2 + (\u03c1\u2212 \u03c1 cos \u2206\u03b8)2 \u2264 3\u03c1\u2206\u03b8. 2377-3766 (c) 2018 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. Lemma 6. (Upper bounding result) The length of the feasible solution obtained by Approx is at most equal to Costlb + 3\u03c1(n\u2212 2)\u2206\u03b8" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002259_j.measurement.2020.108723-Figure15-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002259_j.measurement.2020.108723-Figure15-1.png", "caption": "Fig. 15. Sketch map of external load on the front bearing of induction motor (Ignoring the bending moment).", "texts": [ " In order to further verify the robustness of FHD and the obtained relation curve (FHD curve) in different operating conditions, this section will further simulate and calculate FHD considering the following factors: external load (including torque load and radial load), power supply (including frequency and unbalance of supply), and different motors. In practice, the external mechanical load which mainly includes torque and radial loads would be transmitted to the induction motor when the motor is working (Fig. 15). Torque load directly increases the amplitude of stator current; Radial load may cause some misalignment of the rotor when the load is significant high. On the other hand, the motor usually has some inherent eccentricity due to the assembly accuracy, which is inevitable in a real motor. Therefore, it is necessary to discuss about the influence of initial static eccentricity to FHD in the motor. According to Hertzian contact theory, when radial load is added on the bearing, the bearing inner ring, outer ring and rolling elements bring forth plastic deformation in the contact area as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000720_tmag.2019.2894739-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000720_tmag.2019.2894739-Figure5-1.png", "caption": "Fig. 5. Simulated (a) deformation of the stator scaled by 5 \u00d7 104 and (b) displacement of an outer stator boundary node.", "texts": [ " To understand how the stator vibrations are influenced by the rotor eccentricity along with the additional winding excitation, we simulated the machine under different operation conditions and with both models. It is worth to remember that the eccentricity-related harmonics have also p \u00b1 1 pole pair components, similar to the one generated by the additional winding. Furthermore, the contribution of the magnetostriction to these vibrations has not been clarified up to now. The computed stator deformation at a given time step and the displacement of a node on the outer stator boundary as a function of time are shown in Fig. 5. These results are for the rated operation where the rotor is at the center position. Considering the magnetostriction in the core changes the stator deformation by almost 8% with respect to the linear elasticity case. The vibrations of the stator are further investigated for the three following operation conditions. 1) Only the main winding is supplied, no eccentricity. 2) Both windings are supplied, no eccentricity. 3) Both windings are supplied, and the rotor is eccentric in the positive x-direction" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002594_s10409-021-01079-x-Figure8-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002594_s10409-021-01079-x-Figure8-1.png", "caption": "Fig. 8 Simulation of indentation: a a shape-memory hydrogel substrate (at a temperature of 50 \u2103) compressed by a spherical rigid indenter; b cooling the substrate from 50 \u00b0C to 5 \u00b0C; c rebounds after releasing", "texts": [ " The simulation results of the gel strip in its (I) original state, (II) prescribed deformed state, (III) shape-fixing state, and (IV) recovered state are shown in Fig.\u00a07b. 1 3 As experimentally demonstrated in previous work, the shape-memory effect of the hydrogel can be used to print surface features [27] and also to make grippers to pick up objects with arbitrary shapes [26]. In these applications, there is a common mechanical process, i.e., indentation. In this subsection, we perform a simulation of indentation on a shape-memory double network hydrogel, as shown in Fig.\u00a08. The cylindrical gel; and spherical indenter are modeled with axisymmetric elements. The indenter is first compressed onto the hot hydrogel and then held in position while the gel is cooled (from 50 \u2103 to 5 \u2103). The stress in the gel decreases due to the reduced effective modulus of the primary network. After releasing, the indent is mostly retained with little rebound. In the applications of using the shapememory hydrogel as a gripper pad [26], it is the rebound deformation that contributes to the gripping force to hold the subject to be picked up" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002540_j.triboint.2021.106983-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002540_j.triboint.2021.106983-Figure1-1.png", "caption": "Fig. 1. Working principle of four ball tester.", "texts": [ " It also reduces the dependability of oil viscosity to change with temperature variations. At higher temperature oil will have higher viscosity which will provide required film thickness for protecting moving parts and bearings. At lower temperature it allows fluid to flow easily and reach the bearings. PIB VII increased moving part durability and performance. In the present studies, the experimental conditions were designed such that in non-conformal point and line contact we get boundary/ mixed lubrication regimes. Fig. 1 shows testing rig schematic diagram and working principle. A point contact is formed between a stationary three steel balls with testing oil sample and rotating steel ball. The lubricated contact formed at desired load, speed and temperature generates tangential friction force on three stationary balls, this force can be sensed through load cell when arm of ball pot touched to it. Input parameters like normal load, speed, temperature could be set to desired condition and out put parameters like friction force, vertical displacement; metal contact can be recorded online" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002426_j.engfailanal.2020.105211-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002426_j.engfailanal.2020.105211-Figure2-1.png", "caption": "Fig. 2. A six-dimensional force transducer fixed on the right front wheel.", "texts": [ " According to references [27\u201329], when the vehicle is traveling at low speed, the additional mass of the six-dimensional force transducer has no significant impact on the strain of relevant parts. The wheel center load signal of the test vehicle was measured in Tongxian proving ground under different load conditions. Taking the full load state of the vehicle as an example, the life of the hub bearing was analyzed. The installation position and wheel coordinate system of the sixdimensional transducer force were shown in Fig. 2. The load history of the original load spectrum data collected by the data acquisition instrument after preprocessing, such as separation and extraction, zero drift removal, singular point elimination, filtering and trend term removal, is shown in Fig. 3. Mx is the driving torque of the wheel. The hub unit in this paper is located in the front wheel of the vehicle, so there is no driving moment. My is offset by the brake torque. Mz is the aligning torque of the wheel, its value and frequency are low, so it can be ignored" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002579_j.jmapro.2021.03.032-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002579_j.jmapro.2021.03.032-Figure7-1.png", "caption": "Fig. 7. FEM thermal model for 8L10 P WAAM sample: (a) and (b) are the geometric model after meshing; (c) actual WAAM sample during cooling stage; (d) and (e) are the transient temperature fields at 710 s during deposition and 30 min after cooling respectively.", "texts": [ " In this work, 8-layers & 10-passes (8L10 P) WAAM sample prepared under typical reciprocating cyclic deposition path was taken as a calculation object. Based on the temperature field of the WAAM sample obtained through transient finite element thermal simulation, STK modeling was carried out by using an in-house code, and the thermal history data of WAAM sample was input into the calibrated STK model to predict volume fraction of precipitates, to study the spatial distribution of precipitates related to hardness distribution inside the WAAM sample. As shown in Fig. 7a and b, a well-tested 8L10 P WAAM finite element model (FEM) with dimensions of 30 mm \u00d7 48 mm \u00d7 27.7 mm is X. Lu et al. Journal of Manufacturing Processes 65 (2021) 258\u2013270 established, which is as close as possible to the actual dimensions of the multi-layer & multi-passes sample in Fig. 7c. The 8-node linear hexahedral heat transfer element (DC3D8) provided by the ABAQUS software was used, and the whole model included 27,420 elements and 32,538 nodes. In the modeling process of FEM model, the representative volume element (RVE) was used as the basic molten pool element, and the step-by-step filling simulation of materials was realized by using the element birth and death method, and the double ellipsoid heat source model and dynamic thermal boundary conditions were also applied. Besides, the process parameters of FEM simulation and WAAM experiment were the same, which were optimized according to our previous experimental design. Based on good cladding quality and low input energy (P < 1000 W), the optimum process parameters were determined, that is, deposition speed (3 mm/s), wire feeding speed (3.8 m/ min), voltage (18 V), current (65 A), layer thickness (3.8 mm); overlapping distance (3 mm). Fig. 7d and e are the transient thermal simulation results of the WAAM process of the 8L10 P sample. The transient temperature field at 710 s during deposition process is shown in Fig. 7d. Besides, the heat flow load was removed after the deposition process was simulated, the cooling process was also simulated for a total of 30 min in the final analysis step. The final transient temperature field is shown in Fig. 7e. As shown in Fig. 8 a, the mid-section of the WAAM sample (actual offset 2 mm) was taken on the FEM transient thermal model, and the face marked with a green dashed line corresponds to the real midsection. The diagram also shows the location of the node from which the temperature history was extracted. Multiple points were selected on the whole mid-section for the extraction of temperature history, which X. Lu et al. Journal of Manufacturing Processes 65 (2021) 258\u2013270 can be divided into 4 categories: a Yellow dots, located in the center of the mid-section of each deposition pass" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002700_s12541-021-00510-4-Figure17-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002700_s12541-021-00510-4-Figure17-1.png", "caption": "Fig. 17 Meshing process with the overlap factor of less than 1", "texts": [ " (19)R1 = M1( 1)R (20)R2 = M2( 1z1\u2215z2)Mg1R1 In order to ensure a smooth transmission, the overlap coefficient must be greater than 1; hence, before the demeshing of the target gear pair, the next gear pair must come under meshing. According to Fig.\u00a016, singularity results in the loss of the tooth profile of the ring gear, consequently, the loss of the corresponding meshing part of the gear. In other words, some points at the gear root have no corresponding meshing point, and this part is transformed into a root clearance region. The meshing line in Fig.\u00a017 is cut short because of the singularity. Therefore, point T , the intersection point of the ring gear addendum circle and the tooth profile, should be taken as the starting meshing point instead of point S , the intersection point of the gear dedendum circle and the tooth profile. Subsequently, a part of the meshing line before point T should also be removed. When the overlap is less than 1, the corresponding meshing process is presented in Fig.\u00a017. In this case, the first gear pair has reached its limit position and is about to be 1 3 detached. According to the traditional analysis, when point S enters the meshing line, the latter gear pair is considered to be in the meshing state. Clearly, this analysis is completely wrong. In fact, when point T comes into the meshing line, the latter gear pair begins to mesh. Therefore, in order to satisfy an overlap of greater than 1, point T must enter the meshing zone when the tooth tip of the previous gear pair is out of meshing" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001397_1350650120942329-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001397_1350650120942329-Figure4-1.png", "caption": "Figure 4. Illustration of jet impingement depth on pinion (t0\u00bc 0): (a) schematic diagram; (b) projection line of oil streamline on surface OI p.", "texts": [ " To achieve better lubrication, a positive value for is recommended. The value of is limited to min4 4 max, where min \u00bc tan 1 S hag . RI iv h i max \u00bc tan 1 S\u00fe hap . RI iv h i 8>< >: \u00f06\u00de and RI i \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi RI2 av RI2 s q , RI av \u00bc 1 2 NI gv mI n cos I \u00fe hIag, and RI sv \u00bc RI v \u00fe S: where RI i is the normal offset distance. 1. At the initial moment (t0\u00bc 0): the position parameters of the spiral bevel gears and the jet streamline are as illustrated in Figure 4(a); at this moment, the oil jet starts to pass the gear addendum and approach the pinion meshing surface. The outside radius RI a of the gear at the impingement point is calculated as RI2 a \u00bc x0 \u00fe z0 zI tan cos \u2019 2 \u00fe y0 z0 zI tan sin \u2019 2 \u00f07\u00de where xI \u00bc x0 \u00fe z0 zI tan cos \u2019 yI \u00bc y0 z0 zI tan sin \u2019: ( And Akp is the apex to crown. For the spiral bevel gear RI a \u00bc R Akg Ag zI \u00bc R Ag Hg Ag zI \u00bc Ag zI tan ag \u00f08\u00de The unit direction vectors of the gear axis and the oil jet stream are np, ng, and nVj where np \u00bc \u00f0 sin , 0, cos \u00de, ng \u00bc 0, 0, 1\u00f0 \u00de, and nVj \u00bc sin \u00f0 cos \u2019, sin sin\u2019, cos \u00de. Then, the arbitrary elevation angle between the jet stream and pinion axis is \u00bc arc cos nVj np . Based on the geometric relations of the spiral bevel gear pair, the radius RI e and diameter dI at the impingement point are RI e \u00bc Ag z I\u00f0 \u00de cos ag cos ag dI \u00bc 2rI \u00bc d Re RI e \u00bc 2RI e sin p 8< : \u00f09\u00de As shown in Figure 4, the surface OI p perpendicular to the pinion axis and passing the impingement point on the pinion meshing surface is defined; then on this surface, a line perpendicular to the projection line of the jet stream from the central point OI p is drawn. The perpendicular length r is r \u00bc rz cos l. As shown in Figure 4(b), the center coordinates of the surface OI p are (R I e cos p, 0, Ag RI e sin p); the point VE xE, 0, zE\u00f0 \u00de for the oil flow jet on the symmetry plane is (x y tan \u2019, 0, z y tan sin \u2019). The projection of point VE on surface OI p is V pro E zE b\u00f0 \u00de tan \u00fexE 1\u00fetan2 , 0, zE b\u00f0 \u00de tan \u00fexE\u00bd tan 1\u00fetan2 \u00fe b , where is the shaft angle, b \u00bc Ag RI e sin p RI e cos p tan . The line VyV pro E is perpendicular to line OI pV pro E and parallel to the y-axis. Based on the geometric relationship, l is equal to the angle between line VyV pro E and line VSV pro E ,rz is equal to the length of line OI pV pro E ", " The homogeneous coordinate of the new coordinate system is E \u00bc Trans Ap sin , 0,Ag Ap cos Rot y, 90 U \u00f021\u00de That is x y z 1 2 6664 3 7775 \u00bc cos 2 \u00f0 \u00de 0 sin 2 \u00f0 \u00de Ap sin 0 1 0 0 sin 2 \u00f0 \u00de 0 cos 2 \u00f0 \u00de Ag Ap cos 0 0 0 1 2 6664 3 7775 X Y Z 1 2 6664 3 7775 \u00bc cos 0 sin Ap sin 0 1 0 0 sin 0 cos Ag Ap cos 0 0 0 1 2 6664 3 7775 X Y Z 1 2 6664 3 7775 \u00f022\u00de The outside radius rIIa of the gear at the impingement point is rIIa 2 \u00bc X0 \u00fe Z0 ZII tan sin l 2 \u00fe Y0 Z0 ZII tan cos l 2 \u00f023\u00de where XII \u00bc X0 \u00fe Z0 ZII tan sin l YII \u00bc Y0 Z0 ZII tan cos l and XII,YII,ZII represent the impingement point position in the pinion coordinates. For the spiral bevel pinion rIIa \u00bc ra Akp Ap ZII \u00bc ra Ap Hp Ap ZII \u00bc Ap ZII tan ap \u00f024\u00de where the cone distance RII e at the impingement point on the gear meshing surface is RII e \u00bc Ap Z II\u00f0 \u00de cos ap cos ap . As shown in Figure 4, the surface OI p perpendicular to the pinion axis and passing the impingement point on the pinion meshing surface is defined; then on this surface, a line perpendicular to the jet stream projection line from the central point OI p is drawn. The perpendicular length r is r \u00bc rz cos l. A plane OII g perpendicular to the gear axis and passing the impingement point on the gear meshing surface is drawn; then on plane OII g , a line perpendicular to the projection line of the jet stream from the central point OII g is drawn, and the perpendicular length R is R \u00bc Rx cos \u00f025\u00de where: \u00bc 3 2 \u2019 Rx \u00bc x0 y0 sin \u00bc x0 \u00fe y0 cos \u2019 \u00f026\u00de The initial angular positions of the pinion and gear at the initial time t0 are p4 and g2, respectively, thus g2 \u00bc p4=u\u00fe inv sph , b\u00f0 \u00deg\u00fe \u00f027\u00de p4 \u00bc cos 1 XII =rIIa inv a sph a, b\u00f0 \u00dep \u00fe inv sph , b\u00f0 \u00dep \u00f028\u00de where inv a sph a, b\u00f0 \u00dep denotes the declination angle a sph at the intersection point between the outside circle of the pinion and the involute, and inv a sph a, b\u00f0 \u00dep\u00bc 1 sin bp arccos cos ap cos bp arccos tan bp tan ap " ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000798_j.neucom.2019.04.056-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000798_j.neucom.2019.04.056-Figure1-1.png", "caption": "Fig. 1. A suspension cable system of a helicopter.", "texts": [ " sign ( \u00b7) represents the symbolic function. ln ( \u00b7) denotes the nature logarithm of ( \u00b7). | \u00b7 | stands the absolute value. Throughout the paper, \u02d9 s(z, t) = \u2202s (z,t) \u2202t , \u02d9 s\u2032 (z, t) = \u2202s 2 (z,t) \u2202 z\u2202 t , s\u0308 ( z, t) = \u2202 2 s (z,t) \u2202t 2 , s \u2032 ( z, t) = \u2202s (z,t) \u2202z , s \u2032\u2032 (z, t) = \u2202 2 s (z,t) \u2202z 2 , p \u2032 (z, t) = \u2202 p(z,t) \u2202z , \u03c2 \u2032 (z) = \u2202\u03c2(z) \u2202z , R \u2032 (z) = \u2202R (z) \u2202z , \u03c6\u2032 (z) = \u2202\u03c6(z) \u2202z , and \u03b1\u2032 (z) = \u2202\u03b1(z) \u2202z . 2. Problem formulation and preliminaries In this paper, the schematic diagram is described by Fig. 1 for a suspension cable system of a helicopter. Moreover, the system variables of the helicopter suspension cable system are given by Table 1 . The transverse deflection for the suspension cable system of a helicopter is only considered in the sequel section. Due to the less impact on the helicopter\u2019s center of gravity by using a singlepoint suspension, thus, this kind of lifting manner is adopted in this paper. Since the payload is connected with the boundary of the suspension cable, it has p(0 , t) = 0 with respect to the local reference frame x - y ", " For the C 1 continuous Lyapunov function ( t ) > 0, ith the initial condition (0), if \u03c9 2 1 (\u03b8 (t)) \u2264 (t) \u2264 \u03c9 2 2 (\u03b8 (t)) and \u02d9 (t) \u2264 \u2212\u03c61 (t) + \u03c62 with \u03c61 and \u03c62 being positive constants , 1 (\u03b8 (t)) and \u03c9 2 (\u03b8 (t)) being K\u2212 class functions, then \u03b8 ( t ) is uni- orm bounded. emark 1. The term \u222b L 0 [ \u03c2(z) \u0308s (z, t) \u2212 f (z, t) + c \u0307 s(z, t)] dz in (12) is a ariable only with respect to time t , thus it can exist in the boundry Eq. (12) . In addition, according to Assumption 1 and the bounddness of \u03c7 ( w ( t )), it can be easily to obtain the following concluion that D ( t ) is bounded, i.e., | D (t) | \u2264 D\u0304 with D\u0304 being a positive onstant. emark 2. According to Fig. 1 , one can be easily to obtain that p(0 , t) = 0 ; moreover, assuming that the small segment of the ini- ial end for suspension cable is tightly fixed to the payload for a ingle point hanging safety, thus, s \u2032 (0 , t) = 0 . emark 3. In order to perform the lifting operation secuity, the state of the suspension cable is usually vertically own in the beginning period. Then, according to the fact hat R ( z , 0) := R ( z ), we have R (L ) = m p g + \u222b L 0 \u03c2(z) dzg, where m p nd g are the mass and the gravitational acceleration, repectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001994_lra.2020.3003862-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001994_lra.2020.3003862-Figure7-1.png", "caption": "Fig. 7: Engagement Geometry in 3 dimensions.", "texts": [ " These acceleration commands are then converted into required forces on the robotic fish as follows. The required vertical force is created by varying the gas generation/gas release rate V\u03073. The required horizontal plane forces are created by an appropriate combination of the amplitude Am, bias angle b and flapping frequency \u03c9 governing the sinusoidal motion of the flapping angle \u03b1 of the fish tail. In particular, varying Am creates the required forward force on the robotic fish, and varying b creates the required lateral force. Fig 7 shows the engagement geometry between the robotic fish A and an orifice B, moving in 3-D space, with speeds of VA and VB , respectively, at heading angle pairs of (\u03b2A, \u03b1A), and (\u03b2B , \u03b1B) respectively. Here, \u03b2A and \u03b1A represent, respectively, the azimuth and elevation angles of the velocity vector of A, and a corresponding definition holds for \u03b2B and \u03b1B . When the orifice is stationary, we have VB = 0. We assume that the shape of the robotic fish A is approximated by a bounding sphere of radius RA" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001538_jestpe.2020.3029802-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001538_jestpe.2020.3029802-Figure3-1.png", "caption": "Fig. 3. Longitudinal diagram of EDS Maglev.", "texts": [ " In the guidance system, the figure-eight-shaped coils on the left and right sides are connected by the zero-flux cables. As shown in Fig.2(a), when the maglev deviates laterally to the right, the guidance currents will be induced in unit coils, and each unit coil will generate a guidance component force exerted on the vehicle coils, then a left guidance force opposite to the lateral offset direction is synthesized. This guidance system can realize the independent guidance function, and make the car body recover to the left and right centering position. As shown in Fig.3, when the vehicle coils move at the speed of vx, the discontinuous distribution of the vehicle coils in the running direction will cause the end effect at the entrance end and exit end. Therefore, in the process of EDS maglev operation, the output parameters of the suspension system and the guidance system fluctuate, thus affecting the quality of maglev operation [18]. III. DERIVATION OF ANALYTICAL METHOD In Fig.4(a), it is assumed that the array of the actual vehicle coils is distributed in the xoz plane and their M" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001800_978-3-030-40410-9-Figure4.1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001800_978-3-030-40410-9-Figure4.1-1.png", "caption": "Fig. 4.1 PUMA robot manipulator", "texts": [ " Forward kinematics is a straightforward operation; in contrast, the inverse kinematic problem is more complicated. By using optimization methods, we will simplify the inverse problem to make it direct for all types of robot manipulators, even those with a high degree of freedom. The robot manipulator is a series of mechanical components called links, joined to each other by joints. Usually, manipulators are designed formulti-purpose operations and canbe governeddirectly by a humanoperator or through a logic device. Figure 4.1 illustrates a virtual model for the famous PUMA manipulator. Sometimes, the manipulator is called the robot arm or just arm because it consists of many combined pieces in one mechanism, just like the human arm. A robotic prosthetic arm for patients is a clear example of why the manipulator simply arms. There are two main categories of manipulators: the serial manipulator or sometimes called an open-chain robot, where the terminal link ends with an end-effector (e.g. the robots in Figs. 4.1 and 4" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001075_ecce.2019.8912286-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001075_ecce.2019.8912286-Figure2-1.png", "caption": "Fig. 2. (a) Magnetically active portion of the base design and (b) assembled prototype with a penny for a size reference.", "texts": [ " Additionally, due to the lack of magnetic containment provided by the back irons, the flux from Rotor 1 may extend beyond Rotor 3 [16]. This may cause losses in nearby conductive objects and attract nearby magnetic objects. Also, the use of plastic instead of steel can exacerbate thermal challenges, such as removing heat from eddy currents in the PMs and stresses from uneven thermal expansion. This paper systematically addresses the impacts of various tradeoffs involved in solving some of these fabrication challenges for an example design described in Table I and Fig. 2. This base design operates with Rotor 1 as the high speed rotor, Rotor 2 as the low speed rotor, and Rotor 3 fixed, which results in a 4.67:1 gear ratio. The PMs are NdFeB N52, and the modulators are made of 26 gauge M19 laminations. All simulation sweep results in this paper were produced by sweeping only one or two of the base design parameters from their base values at a time. II. MODULATORS SUPPORT To simplify handling and fabrication, the modulator stack was bonded. However, additional modulator support is required", " For this design, these rods are made of G10 fiberglass-epoxy laminate to provide high strength with high electrical resistivity. Additionally, the slots between adjacent modulators are filled with glass-filled nylon spacers. Circular arc shaped holes are cut out of the modulators\u2019 inner corners to allow the spacers to interlock well with the modulators. Both the circular holes for the rods and the arc shaped holes remove magnetically permeable material from the flux paths in the modulators, as depicted in Fig. 2(a). Fig. 3 illustrates how the radii of these holes impact the design\u2019s slip torque and full load electromagnetic efficiency at the rated Rotor 2 speed of 400 rpm based on finite element analysis (FEA). Generally, these holes do not have much effect on the performance unless they become large enough that the area between them is thoroughly saturated, worsening gear performance. For this design, the holes have radii of 1.2 mm. Fig. 3(b) also indicates that this design has a very high electromagnetic efficiency (neglecting mechanical losses), which mitigates thermal concerns", " In a conventional magnetic gear without Halbach arrays and with back irons, the Rotor 3 PMs are attracted to the Rotor 3 back iron. However, the Halbach array results in forces pushing the radially magnetized PMs inward towards Rotor 2. Thus, it is necessary to retain the Rotor 3 PMs, but it is highly undesirable to do this in a way that increases the effective magnetic air gap. One alternative is to reduce the PM fill factor and use nonmagnetic walls between the PMs to hold them in place, particularly if the walls are wider at the radial inside than they are at the outside, as shown in Fig. 2(a). On Rotor 1, adding walls between the magnets allows the sleeve to be formed as a single piece with the plastic core and strengthens the sleeve. For both rotors, these walls also help to position the PMs. Fig. 7 illustrates the impact of changing the widths of the walls between adjacent PMs. Comparing Fig. 6(a) and Fig. 7(b) reveals that reducing the Rotor 3 PM fill factors to accommodate this retention strategy lowers the slip torque less than increasing the effective outer air gap to insert a Rotor 3 PM retention sleeve", " This enabled the PMs to be aligned axially beyond most of the magnetic fields and then pushed into place. Similar tooling was created for inserting the Rotor 3 PMs. Also, the radially magnetized PMs were inserted before the tangentially magnetized PMs to avoid strong forces pushing the PMs being inserted away from the plastic cores. With this strategy, all of the PMs were successfully inserted, and each of the authors still has 10 unmaimed fingers. Fig. 9 shows the resulting assemblies. The finished prototype is illustrated in Fig. 2(b). End supports were included to mount the prototype to the testbed. The overall mass of the prototype was 4.4 kg with the end supports and 3.6 kg without the end supports. Table II provides a breakdown of the measured masses of some of the different prototype components. The measured Rotor 2 slip torque was 31.2 N\u2219m, which is 2.6% higher than the simulated slip torque of 30.4 N\u2219m. This yields GTD values of 25.2 N\u2219m/kg considering only the active material, 8.7 N\u2219m/kg considering the total mass of the prototype without end supports, and 7" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001099_j.mechmachtheory.2019.103739-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001099_j.mechmachtheory.2019.103739-Figure3-1.png", "caption": "Fig 3. Manufacturing machine coordinate system.", "texts": [ " Moreover, the parameters \u03b4s , R, \u03b2v /( \u03b2c ) and \u03b2 are offset angle of cutter blade, cutter plate radius, the initial assembly angle of outer blade/(inner blade) and the rotate angle of cutter plate, respectively. The initial assemble angle of outer blade and inner blade is fixed. In this paper, the initial assemble angles of outer blade and inner blade are 10.5882 \u25e6 and 21.1765 \u25e6, respectively. Further, the transform matrix of machine from S t to S p are proposed in order to get the cutting track of blade on the gear blank. Fig. 3 represents the machine coordinate system from the center of cutter plate to gear blank for hypoid gears. The generator coordinates have nine parameters including the tilt angle i , swivel angle j , radical distance S R , cradle angle q 0 , \u03d5c ( q 0 is initial cradle angle and \u03d5c is cradle rotation angle), work offset E , sliding base B , machine root angle \u03b3 m and horizontal A . The transformation matrixes of manufacturing machine from S t to S p can be obtained as follow M pt = M pk ( \u03d5 1 ) \u2217 M kh \u2217 M hg \u2217 M gm \u2217 M mc ( \u03d5 c ) \u2217 M ce \u2217 M f e \u2217 M et (11) In order to solve this equation, the processing methods and the principle of gear meshing must be applied" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000473_j.proeng.2016.06.307-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000473_j.proeng.2016.06.307-Figure1-1.png", "caption": "Fig. 1. Experimental determination of the restitution coefficient CR: (a) raw image showing traces of the ball before, during and after contact with the rigid surface, (b) incident and reflected velocities are determined for different positions with respect to the impact location and extrapolated to determine the coefficient of restitution.", "texts": [ " Set-up and protocol A home-made apparatus has been designed to launch table tennis balls yet controlling the ball\u2019s initial velocity and spin. The target is fixed normal to ball\u2019s trajectory onto the rigid frame, 260 mm away from the launcher\u2019s end. A camera coupled with stroboscopic lighting is set to acquire pictures which display the target as well as numerous traces of the ball before and after impact. Marks are drawn onto the target to locate its center and onto the ball to ascertain negligible spin and locate the joint equator. Two examples of recorded pictures are provided in Fig. 1a&b with clear identification of the drawn marks. Using the aforementioned setup, normal impact experiments were performed at room temperature for launching velocities ranging from 5 to 30 m/s, consistent with the speed of the ball in game conditions. Twenty consecutive impacts were repeated for a given set of {target; launching velocity}. The first ten shots aimed at \u201cwarming up\u201d the ball and the target materials. Out of the last ten shots, three pictures were visually selected so that: \u2022 The ball spin was negligible: spin would cause noticeable deviation of the meridional marks drawn on the ball between two consecutive traces", " \u2022 The impact was normal: in the xy-plan of the pictures, this is confirmed when all the traces of the ball lie on a unique straight line. Now considering a potential tilt perpendicular to the picture, the perspective would cause the ball\u2019s size to decrease or increase as it moves away from the camera or towards the camera respectively. \u2022 The contact zone did not contain the joint of the ball: this is ascertained by locating the dashed-line disk almost perpendicular to the ball\u2019s trajectory. With the aid of the remaining filtered pictures, the coefficient of restitution was calculated following the procedure depicted in Fig. 1b and detailed hereafter: incident and reflected traces of the ball were separated assuming that the distance will be lower for two successive reflected traces compared to two successive incident ones. The centers of the ball\u2019s traces were then calculated by means of the image processing free software ImageJ\u00a9. Using the {xc,yc} coordinates of two consecutive centers, the ball velocity was determined with respect to its position from the target (xtarget=0), knowing that subsequent traces are separated by a time increment \u0394ts=1/fs with fs the stroboscopic frequency", " Furthermore, frictionless simulations tend to over predict the experimental measurements and do not seem to evolve with speed so that the discrepancy between the experiments and the simulations increases with increasing velocity. Ultimately, an important cause of energy dissipation comes from the friction process. This major role can be explained by the sample/ball contact zone which is identified as a ring with its mean radius increasing till the maximum crushing. The ball buckling phenomena has been largely studied and reported in the literature [8], observed experimentally in this work (see the contact zone in Fig. 1a) and simulated numerically (see Fig. 5a&b). The greater the incident velocity, the larger the maximum radius of the ring and the more important the friction dissipation is. Assigning the same constitutive equations to the foam and the compact is pertinent when looking at the DMA measurements presented in Fig. 3b. Besides, it permits to study of the effect of the targets architecture. This is done by predicting the restitution coefficients of three targets: a 2 mm thick foam, a 4 mm thick foam and a 2+2 mm thick foam+compact" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000888_j.promfg.2019.06.223-Figure19-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000888_j.promfg.2019.06.223-Figure19-1.png", "caption": "Fig. 19. An AM system with two robotic arm actuators for fabrication of a single-material fixture.", "texts": [ " Moreover, a 20mm safety distance is assigned to each robotic arm according to the dimensions of their end-effectors. After generating concurrent toolpaths, the multi-material fixture can be digitally fabricated and visualized, as shown in Fig. 18. The total build time is about 12.05 hours with the concurrent toolpaths, and 13.26 hours with sequential toolpaths. For some applications, a single-material fixture may be good enough and cost-effective. In this case, a monochrome STL model of the fixture can be loaded, as shown in Fig. 19. The region-based constraint, the slicing parameters and the hatch width are the same as before. The total build time of the single-material fixture is about 11.82 hours with concurrent toolpaths, and about 13.26 hours in a traditional AM system with only one actuator. To illustrate the concurrent toolpaths, a layer of the single-material fixture model is selected. Fig. 20 shows the CFs with their safety envelopes, while Table 2 shows the deposition groups and the toolpath sequence for this layer" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000716_s00170-019-03372-3-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000716_s00170-019-03372-3-Figure7-1.png", "caption": "Fig. 7 Temperature field and velocity profile at cross-section (PL = 2 kW) t = 0.61 s (a), t = 0.88 s (b), and t = 1.11 s (c)", "texts": [ " Figures 6 and 7 offer the temperature filed and fluid flow in hybrid fillet welding at 2 kW laser power. It is observed that, due to the relatively lower laser power, keyhole phenomenon is not very obvious. In this case, the laser has minor effect on weld pool dynamic behavior. The basic flow pattern of liquid metal is similar to that in MIG welding, and the sagging of molten pool is still evident. But, owing to recoil pressure, the peak flow velocity of liquid metal in hybrid fillet welding is higher than that in MIG welding, which reaches 2.8 m/s. In Fig. 7, it is seen that, in hybrid fillet welding, the weld pool depths at vertical and horizontal plates tend to be close owing to the addition of laser energy with high heat density. Besides, the compaction of electric arc by laser beam also provides some contributions to this change of weld pool geometry, which decreases the arc distribution area and helps to reduce the difference in heat inputs acting on the vertical and horizontal plates. When laser power increases to 4 kW, a deep keyhole appears in hybrid weld pool because of the further enhancement of evaporation-induced recoil pressure" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002615_s10853-021-06012-y-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002615_s10853-021-06012-y-Figure1-1.png", "caption": "Figure 1 a Outside view of TA15B part; b deposition strategy; c the layout of the specimens; d dimension of the tensile specimen. The directed lines in b indicate the track of the deposition.", "texts": [ " Two block parts with multilayer multi-bead structure (about 65 mm long, 65 mm wide, and 40 mm high) were built by using the same deposition parameters (Tables 1, 2). Equation (1) is used to calculate the average arc heat input \u00f0Hi\u00de [24, 25]: Hi \u00bc gIavVavv 1 s \u00f01\u00de where g is the welding efficiency (assumed to be 0.83 [24]), vs is the scan speed, and Iav and Vav are the average current and average voltage of GTAW, respectively. The part built by using TA15 wire without boron addition was called TA15 part. The other part built by using TA15B wire with boron addition (0.1wt.%) was called TA15B part (Fig. 1a); i.e. the B-modified part was built by using B-modified feedstock wire directly. The chemical analysis results of the main alloying elements of the wire and the built parts are shown in Table 2. The Z-direction was denoted as the build direction (BD). The two parts were built in an argon atmosphere to avoid oxidation. A zigzag deposition strategy was adopted, and the deposition track between layers was same but with opposite direction, as shown in Fig. 1b. The two parts were deposited continuously without inter-layer or inter-bead dwelling time. The microstructure was observed by optical microscope (OM), scanning electron microscope (SEM) and electron backscatter diffraction (EBSD). The preparations of the metallographic and EBSD specimens were performed on the X\u2013Z sections. The metallographic specimens were mechanically ground, polished, and etched with Kroll\u2019s reagent (2% HF ? 4% HNO3 ? 94% H2O). The EBSD specimens were prepared by electropolishing", " The EBSD data were processed by using TSL OIM Analysis 6 software, and the results here were depicted with inverse pole figure (IPF) coloration with respect to the Z direction (BD). Tensile specimens were machined from the two parts in Y direction and Z direction. Room-temperature tensile testing was performed by using an Instron 5966 electronic universal material testing machine at a strain rate of 0.001 s-1. Five tensile specimens of each group were tested. The fractography and the longitudinal section near the fracture surface were observed by SEM. The layout of the specimens is shown in Fig. 1c. The dimension of the tensile specimen is shown in Fig. 1d. The volume fraction of the retained b phase in TA15 and TA15B alloys at room temperature was less than 3% in the present work. To determine the morphology and texture of the parent b grains, the b reconstruction process was performed by using the measured EBSD data of the a phase. This process was performed by using software ARPGE. ARPGE can automatically reconstruct the parent grains from the EBSD data obtained on phase transition materials with or without residual parent phase [26]. For Ti alloys, the phase transition between the body-centered cubic b phase and hexagonal-close-packed a phase follows the Burgers orientation relationship (BOR)" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000123_s1560354719050071-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000123_s1560354719050071-Figure6-1.png", "caption": "Fig. 6. Trajectories for K3 = 0.3 on the Poisson sphere, its involute ( \u03d5 = \u03c0 4 + arctan \u03b31 \u03b32 ) and the", "texts": [ " The isolated equilibrium points of the second type are \u0398 = {\u03b31 = sin\u03d5, \u03b32 = 0, \u03b33 = cos\u03d5}, \u03b81 = { \u03d5 = arcsin(\u03ba2)\u2212 \u03bd2 2 } , \u03b82 = { \u03d5 = arcsin(\u03ba2)\u2212 \u03bd2 2 + \u03c0 } , \u03b83 = { \u03d5 = \u2212arcsin(\u03ba2) + \u03bd2 + \u03c0 2 } , \u03b84 = { \u03d5 = \u2212arcsin(\u03ba2) + \u03bd2 2 + 3\u03c0 2 } , \u03ba2 = (a1 \u2212 a3)\u03a93 + 2a1a3K3\u221a \u03a92 1 +\u03a92 3(a3 \u2212 a1) , \u03bd2 = arcsin ( \u03a92 3\u221a \u03a92 1 +\u03a92 3 ) , which exist if |\u03ba2| < 1. Depending on the system parameters, the conditions for stability of these equilibrium points are represented in the form of extremely cumbersome relations. Therefore, we illustrate the evolution of the phase portrait of the system (4.1) with increasing |K3| for the fixed parameters \u03a9 = (4, 2, 1), A = diag(0.5, 0.4, 0.3) and a = 1 (see Figs. 6\u20138). If 0 < \u03ba2 < \u03ba1 < 1, then on the Poisson sphere there are ten isolated fixed points of different types (see Fig. 6): \u2014 four saddles (\u03b81, \u03b82, \u03b43, \u03b44); \u2014 two stable foci (\u03b83, \u03b84); \u2014 two unstable nodes (\u03b41, \u03b42); \u2014 two slow stable foci (\u03935, \u03936). We note that, in this case, the unstable and stable manifolds of equilibrium points \u03b43, \u03b44 and the central manifold at \u03b41, \u03b42 form a closed contour. Let \u03ba2 < 1 < \u03ba1. Then the equilibrium points \u03b41, \u03b42, \u03b43 and \u03b44 disappear, and an unstable limit cycle arises from the closed contour (see Fig. 7). If 1 < \u03ba2 < \u03ba1, then of all the equilibrium points the system (4.1) has only \u03935 and \u03936 and the unstable limit cycle (see Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001538_jestpe.2020.3029802-Figure14-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001538_jestpe.2020.3029802-Figure14-1.png", "caption": "Fig. 14. 3-D model in Magnet R2014.", "texts": [ " This means that if the maglev has a large lateral deviation after being disturbed, the guidance force will be reduced, so that the maglev may not be able to recover to the centering position, which will affect the operation quality of maglev. Fig.13 shows the trend of levitation and guidance stiffness with \u0394z when h is 0.09m and running speed is 500km/h. The levitation stiffness decreases and the guidance stiffness increases with the \u0394z increase. In the rated operation state (\u0394z=0), the levitation stiffness is 1.86 times of the guidance stiffness. E. 3-D FEM Simulation and Experimental Verification The 3-D model is shown in Fig.14. For the convenience of illustration, only one side of vehicle coils and ground eight-shaped coils are drawn. Fig.15 shows the comparison Authorized licensed use limited to: Carleton University. Downloaded on November 04,2020 at 22:40:21 UTC from IEEE Xplore. Restrictions apply. 2168-6777 (c) 2020 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. results of the levitation force on one side of each bogie obtained by numerical calculation and 3-D FEM simulation when h is 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000442_tsmc.2016.2560528-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000442_tsmc.2016.2560528-Figure5-1.png", "caption": "Fig. 5. Example of the structural controllability for a network system with four identical nodes. (a) Network is controllable and the directed chain path is shown in bold arrows. (b)\u2013(d) Networks are not controllable due to a lack of directed chain path between identical nodes.", "texts": [ " (13) Using Lemma 4, the matrix expressed in (13) is full rank if and only if a(i+1)i = 0 for i = 1, . . . , n \u2212 1. Hence, the system is structurally controllable if and only if a(i+1)i = 0 for i = 1, . . . , n \u2212 1. Therefore, we can state the following corollary. Corollary 2: The network system with a single driver node and identical nodes is structurally controllable if and only if the network topology contains a simple directed chain path. For illustration, consider a network system with four identical nodes as represented in Fig. 5. It is shown that, in the network system with all identical nodes, existence of a simple directed chain path in the associated digraph is a crucial condition for structural controllability. Hence, for a network system with n identical nodes, if an identical node is added to the network, to preserve the structural controllability, the node must be connected to the last node as depicted in Fig. 6. For clarity of presentation, we first consider a network system with two identical nodes and then similarly we generalize to a network system with multiple identical nodes" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002444_j.measurement.2021.109021-Figure15-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002444_j.measurement.2021.109021-Figure15-1.png", "caption": "Fig. 15. Health status of the low speed stage components after 345 h of run time (a), (b), (c) & a gear tooth damage observed in an actual wind turbine industry (d), (e) [34].", "texts": [ " 14 depicts the \u2018R2\u2019 value for all the sensor channels. It can be noted that, the A3 channel has yielded highest \u2018R2\u2019 values and has registered better precision. The LSS (planetary 1) comprises of six 6004 2RSH bearings, three planet gears, one sun gear, one ring gear and two SYJ40TF bearings. As described in the Fig. 2, a tri-axis accelerometer was mounted on the LSS bearing. The horizontal component (X-axis) and vertical component (Yaxis) of vibration data were chosen and designated as axis A1 and A2, respectively. Fig. 15 illustrates the health condition of LSS stage at the end of the runtime. It was observed that, the LSS has suffered a catastrophic damage and stopped the working of the wind turbine gearbox. The teeth of all the planet gears were deformed (plastic flow), oil seal of multiple 6004 2RSH bearings and one SYJ40TF bearing failed, refer Fig. 15 (a). In addition to that, the sun gear teeth were stripped (tooth fracture) and excessive metal fragments were deposited on the planet carrier, due to which the LSS had become non-operational, refer Fig. 15 H.M. Praveen et al. Measurement 174 (2021) 109021 (b) & (c). Fig. 16 describes the predicted vs actual health condition for the LSS stage using A2 at scale 5 (A2-5-1, A2-5-2 & A2-5-3) and A1 scale 5 (A1-5-1, A1-5-2 & A1-5-3). It was observed that, the prediction capability for channel A2 was better than A1 and the RUL prediction of A2 channel could distinguish all the four health zones of LSS of the gearbox. The green zone was until ~120 h, blue zone was until ~190 h and the orange zone was until ~275 h of the operational time of gearbox respectively, refer Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000173_01691864.2019.1680316-Figure15-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000173_01691864.2019.1680316-Figure15-1.png", "caption": "Figure 15. Geometric relationshipswhen shifting the elongating part. (a) Before shifting. (b) While shifting.", "texts": [ " If (5) and (6) have a solution without imaginary parts, the following condition has to be satisfied: l2x + l2z \u2264 4L2. (9) Figure 14 shows the relationship between d and the maximum h which is numerically obtained in the case where l = 0.01m, L = 0.181m, r = 0.075m, and lb = 0.025m. Let the start time of the shift be t = 0 and the end time of the shift be t = tshift. If the shift from Figure 11(a,b) is carried out, we design the joint angles according to the following geometric relationships as Figure 15: \u03b8i+2 = (\u03b21 \u2212 \u03b12)t\u2032 + \u03b12, (10) \u03b8i+3 = (\u03b22 \u2212 \u03b11)t\u2032 + \u03b11, (11) \u03b8i = 2 atan2 ( 2L z\u2212 \u221a ( x2+ z2)(4L2\u2212 x2\u2212 z2) x2+ z2+2L x ) , (12) \u03b8i+1 = 2 atan2 ( 2L z+ \u221a ( x2+ z2)(4L2\u2212 x2\u2212 z2) x2+ z2+2L x ) , (13) where t is time, t\u2032 = 6t\u03035 \u2212 15t\u03034 + 10t\u03033 is the quintic curve, given as a cam curve, t\u0303 = t/tshift, 0 \u2264 t \u2264 tshift, and x = d + L(cos\u03b21 + cos\u03b22) \u2212 L(cos \u03b8i+2 + cos \u03b8i+3), (14) z = h + L(sin\u03b21 + sin\u03b22) \u2212 L(sin \u03b8i+2 + sin \u03b8i+3). (15) If (12) and (13) have a solutionwithout imaginary parts, it is necessary to satisfy the condition 4L2 \u2212 x2 \u2212 z2 \u2265 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002827_s11694-021-01098-z-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002827_s11694-021-01098-z-Figure1-1.png", "caption": "Fig. 1 Schematic representation of the process of graphene wrapping the real sample microparticles and the subsequent electrochemical detection steps", "texts": [ " Then 10\u00a0mg of crushed real sample was added into 10 mL of diluted graphene dispersion and then sonicated for 3\u00a0min. A certain amount of mixture was then directly drop coated on the glassy carbon electrode (GCE) surface. The modified GCE was then dried naturally. The electrochemical experiments were carried out on a CHI760E electrochemical workstation (Shanghai Chenhua Instruments, China) with a three-electrode system: a GCE (3 mm diameter) and its modified electrode as the working electrode, a platinum wire electrode as the counter electrode, and an Ag/AgCl electrode as the reference electrode. Figure\u00a01 shows a schematic diagram of the real sample microparticles wrapped with graphene and then immobilized on the electrode surface. This method of coating samples with graphene has two benefits. The first is that the excellent conductivity of graphene can increase the response current of molecules during electrochemical reactions [30]. The second is that the coating of sample particles with graphene can well improve the stability of the immobilization. This can keep the sample particles from falling off during contact with the electrolyte" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000614_1.4035407-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000614_1.4035407-Figure1-1.png", "caption": "Fig. 1 (a) Chaplygin sleigh with a balanced rotor. The center of mass of the arrangement is at a distance b from the rear contact. The picture on the right shows an illustration of a physical cart to realize the Chaplygin sleigh with castors at the front.", "texts": [ " We show that the reduced state space of the system can be partitioned into sets of distinct dynamics and that the stick\u2013slip transitions can be explained in terms of transitions of the state of the system between these sets. [DOI: 10.1115/1.4035407] Nonholonomic constraints in mechanical systems are often encountered due to the friction imposed no-slip condition between a body and some surface. Even very simple systems that are subject to nonholonomic constraints, such as disks and spheres moving on a rough surface, exhibit very rich and often nonintuitive dynamics [1\u20134]. One such classical example of a nonholonomically constrained system is the Chaplygin sleigh [1,2] shown in Fig. 1. This system has also been one of the canonical examples in the study of the control of nonholonomic mechanical systems [3,5\u20138]. The Chaplygin sleigh has assumed additional importance recently in the area of aquatic robots [9], as the constraints encountered in the swimming motion of an idealized fish-shaped body have been shown to be similar to those on the Chaplygin sleigh [10,11]. An interesting feature of the Chaplygin sleigh, as with many nonholonomic systems, is that its reduced equations of motion are dissipative in the sense of decreasing phase space volumes while conserving kinetic energy and its asymptotic motion is along a straight line", " This is a useful idealization to investigate the utility of the relaxation of the nonholonomic constraint in controlling the motion of the Chaplygin sleigh. The idealized Chaplygin sleigh discussed in Refs. [1\u20133] consists of a sleigh with a sharp runner at the rear that is in contact with the ground at a single point. The runner and hence the sleigh can slip freely along the longitudinal axis of the sleigh but the normal velocity of the runner is constrained to be zero. However, a more physically plausible realization of the Chaplygin sleigh consists of a cart with a large heavy wheel at the rear in contact with the ground, as shown in Fig. 1. The point of contact of the wheel is denoted by P. The wheel is assumed to roll without slipping in the longitudinal direction (along Xb) but is constrained to not slip in the lateral direction (along Yb). At the front of the sleigh are castors that allow rotation and motion of the sleigh in any direction. This classical model is modified with the addition of a balanced rotor of moment of inertia J, placed at the center of mass of the sleigh. The distance of the center of mass of the combined system from the wheel is denoted by b", " A representative set of initial conditions \u00f0ux\u00f00\u00de \u00bc 5;x\u00f00\u00de \u00bc 2; h\u00f00\u00de \u00bc 0\u00de are chosen. The results of this simulation are shown in Fig. 3. During the motion, the angular velocity of the cart converges to zero (Fig. 3(a)), i.e., cart moves asymptotically along a straight line, see Fig. 3(b). We relax the nonholonomic constraint on the wheel of the cart by imposing a limit on the friction force that enforces the constraint. When friction exceeds the maximum allowed value, denoted by Fc, the sleigh is allowed to slip in the normal direction at the rear wheel, i.e., point P (Fig. 1) is allowed to move in the Yb direction. In this slip mode of motion, friction does work on the sleigh and dissipates the kinetic energy. When the normal motion of the contact point becomes zero and the required friction force is below the maximum allowed value friction, the motion of the sleigh transitions back to the stick mode, one where the nonholonomic constraint holds. During the stick mode of motion, the kinetic energy of the sleigh remains constant. Depending on the initial conditions, ux\u00f00\u00de and x\u00f00\u00de, the Chaplygin sleigh can transition between stick and slip modes of motion multiple times" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001934_j.mechmachtheory.2020.103895-Figure12-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001934_j.mechmachtheory.2020.103895-Figure12-1.png", "caption": "Fig. 12. Stiffness analysis. with multiple input/output nodes. The input nodes are represented with blue circles with crosses inside, the output inner nodes are represented in red color with empty circles, the output boundary nodes are represented in red color with filled circles. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)", "texts": [ " Therefore, it cannot be an internal node but must belong to the set of the boundary nodes. An output node is a node used to acquire information. For example, it can be a node subject to measurements with accelerometers or optical measurement systems. The output nodes can include both inner and boundary nodes. If a reduction method is adopted, the displacements of the inner nodes can always be recovered through the transformation matrix of the reduction method. For the numerical simulations, the MD1 model described in the Section 4.1 will be used. Regarding Fig. 12 , four input nodes for the application of external forces and three output nodes for detecting displacements will be considered. In particular, two of the three output nodes, i.e. E and V 1 belong to the set of the boundary nodes while the third, T , is an inner node of M . Using the above layout, the Fig. 13 reports four load cases and the corresponding deformed shapes of the robot. Table 13 shows the values of the displacements of the output nodes. The load cases and the configuration p E = [0 , 0 , \u22120 " ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001769_s11668-020-00812-1-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001769_s11668-020-00812-1-Figure2-1.png", "caption": "Fig. 2 Presence of flaking on inner race of the bearing", "texts": [ " Since the MCD located in this gearbox was found with metal debris, more care was taken for thorough investigation of every component in this assembly. Oil pump unit, oil tank, oil pumps and filters in the oil circuit were also thoroughly examined for any abnormality. Visual and Macro-examination All the bearings of the reduction gear box were subjected to serviceability check. One of the ball bearings was found with flaking on the balls and outer race. It was cut open in two halves for binocular examination. Binocular examination was carried out at 16X. Light flaking was noticed on the entire circumference of inner race as shown in Fig. 2 (Fig. 3). Sever flaking was noticed on the surface of few of the balls as shown in Fig. 4. Sever flaking was also noticed on the outer race of bearing track as shown in Fig. 5. One half of outer race was found with coarse flaking while other half of inner race showed finer flaking as evident from Fig. 5. No damage was noticed on the roller cage. Another roller bearing was found to have minor frettage on rollers. Detailed examination of the parts of ball bearing revealed that there was discoloration on the inner diameter of the outer ring of the bearing to some extent" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002656_tmech.2021.3082935-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002656_tmech.2021.3082935-Figure1-1.png", "caption": "Fig. 1. Schematic diagram of the SbW system.", "texts": [ " Compared with SbW control systems [6]\u2013[13], the unnecessary sensor can be removed and CAN communication resources can be saved. Compared with event-triggered output feedback control [24]\u2013[26], [35], both the observation error and tracking error can converge to the neighborhood of the origin in finite time rather than exponentially, such that the transient performance is improved. Notations: For a matrix A, \u2016A\u2016 denotes the Euclidean norm of A, and \u03bbmax(min)(A) denotes its maximum (minimum) eigenvalue. bxe% = |x|%sign(x) with % \u2265 0 Fig. 1 shows a schematic diagram of the SbW system for automatic vehicles. To describe the dynamics model clearly, Tab. I is given to clarify each of the components in Fig. 1. According to the researches [6], [7], [36], [37], the dynamics model of the steering motor can be established as Jm\u03b8\u0308m + Bm\u03b8\u0307m + \u03c412 = \u03c4m + \u03c4d. (1) The rotation of the front-wheels around their vertical axes can be modeled as Jf \u03b8\u0308f +Hf (\u03b8f , \u03b8\u0307f ) = \u03c4s (2) whereHf (\u03b8f , \u03b8\u0307f ) = \u03c4e+\u03c4f denotes the uncertain nonlinearity. The transmission ratio between the steering motor and frontwheels is \u03b8f \u03b8m = \u03b8\u0307f \u03b8\u0307m = \u03b8\u0308f \u03b8\u0308m = \u03c412 \u03c4s = 1 \u00b5 (3) which together with (1)-(2) gets Je\u03b8\u0308f + \u00b52Bm\u03b8\u0307f +Hf (\u03b8f , \u03b8\u0307f ) = \u00b5(\u03c4m + \u03c4d) (4) Authorized licensed use limited to: BOURNEMOUTH UNIVERSITY" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001549_j.ymssp.2020.107328-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001549_j.ymssp.2020.107328-Figure6-1.png", "caption": "Fig. 6. Slider-crank mechanism and the driving system.", "texts": [ " [56,53]. By generating wear depth increment of each joint and calculating the motion output until the time point when the motion output exceeds its threshold, the time interval between the latest observation time point and the time point to reach the threshold is the RUL of a mechanism. Using all the samples, the necessary statistics of RUL for the mechanism of interest can be acquired. A slider-crank mechanism is used to illustrate and validate our method. The slide-crank mechanism is shown in Fig. 6. The slider can yield a reciprocating motion using a motor-driven system. Driven by a motor, the transmission shaft rotates anticlockwise. Then linkage EF pushes Slider 2 downward along Rail 2, and linkage L2 drives Slider 1 to move left along Rail 1. Once Slider 2 reaches the maximummovement, it begins to move upward along Rail 2. Meanwhile, Linkage L2 drives Slider 1 to move right along Rail 1. The angular velocity of the motor is 15 RPM. The load applied on the mechanism is 500 N. In the slider-crank mechanism, there are three revolute joints: joint A, joint B and joint C" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000539_s00170-016-9337-2-Figure39-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000539_s00170-016-9337-2-Figure39-1.png", "caption": "Fig. 39 \u201c3 + 1 axis roughing\u201d method", "texts": [ " For each section\u2019s milling, we record the machining time and observe the overload indicator of the machine tool. When the feed rate increases, the machining time decreases accordingly, while the overload of the machining tool remains at a stable value (\u22487 %). This means that there is room to increase the feed rate. When the spindle speed is 3200 r/min and feed rate is 4000 mm/min, the total machining time is about 60 min. The machining time consumed by the conventional \u201c3 + 1 axis roughing method\u201d generated by software UG NX 8.5 (as shown in Fig. 39) with the same machining parameters is about 90 min. So, in regard of the machining efficiency, our algorithm has better performance. As shown in Fig. 35, both the color distribution and the size of the chips are uniform, which are desired by the blade manufacturer. In this section, we use simulations to verify that hybrid method can further improve the machining efficiency of the tool path. We simulate the 3D platform surface machining process using software VERICUT\u00ae. The machining parameters are the same with the ones in the machining experiment (in Table 2)" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002604_lra.2021.3071673-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002604_lra.2021.3071673-Figure2-1.png", "caption": "Fig. 2. Quad-rotor coordinate system.", "texts": [ " In particular, we only consider the case that the morphing rate is relatively low, which make the first item in (2) have less contribution to the system dynamics. Authorized licensed use limited to: Carleton University. Downloaded on June 04,2021 at 23:06:27 UTC from IEEE Xplore. Restrictions apply. Dynamics for Conventional Quad-rotor: To achieve the motion dynamics of a morphing quad-rotor, we start the step from a conventional quad-rotor\u2019s dynamics. The coordinate system of the quad-rotor is illustrated in Fig. 2, and the related parameters are defined in Tab. I. Note that the origin of the body fixed frame B coincides with the center of gravity (COG) of the quad-rotor. The dynamic model of a quad-rotor can be expressed as p\u0308 = 1 M R\u0304T \u2212 ge3 (3) \u03b3\u0308 = f + J\u03c4 (4) where e3 = [0, 0, 1]T, R\u0304 = [cos\u03c6 sin \u03b8 cos\u03c8 + sin\u03c6 sin\u03c8, cos\u03c6 sin \u03b8 sin\u03c8 \u2212 sin\u03c6 cos\u03c8, cos\u03c6 cos \u03b8]T, and f = [f\u03c6, f\u03b8, f\u03c8] T = [I\u22121 x (Iy \u2212 Iz)\u03b8\u0307\u03c8\u0307, I \u22121 y (Iz \u2212 Ix)\u03c6\u0307\u03c8\u0307, I \u22121 z (Ix \u2212 Iy)\u03b8\u0307\u03c6\u0307] T. According to [16], we can achieve a single matrix (5) to describe the mapping of force and torque for a conventional quad-rotor" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001549_j.ymssp.2020.107328-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001549_j.ymssp.2020.107328-Figure5-1.png", "caption": "Fig. 5. Three distinct motions of the journal inside the bearing.", "texts": [ " If the motion output is over a pre-defined threshold, the mechanism fails. The performance function is given as: g \u00bc G\u00f0L; c; y; r;u\u00de \u00f011\u00de where L, c, y, r, and u are the inputs of the performance function, and g is the motion output. Linkage dimensions and the initial joint clearance can be obtained from the design specifications of a mechanism. For the relative position between the journal and the bearing, there are three distinct motions: free flight, impact and contact motions [48], as is shown in Fig. 5. The free flight motion occurs when the journal moves freely inside the bearing without contact with the bearing. This mode is common in high-velocity mechanisms. The impact mode is realized when the journal moves in contact with the internal surface of the bearing. The contact motion is observed when the penetration occurs between the bearing and the journal. It is obvious that the motions of the journal inside the bearing comprehensively influence the accuracy of the mechanism. As compared with the free flight motion, the impact and contact motions have a more severe influence on the motion accuracy of a mechanism" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002192_toh.2020.3029043-Figure8-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002192_toh.2020.3029043-Figure8-1.png", "caption": "Fig. 8. (a) 3-axis linear servo Cartesian robot. Three blue arrows denote three motion axis separately. (b) End of robot arm.", "texts": [ " With the aid of the rules, the fuzzy logic controller is capable of imitating the human intelligence in decision making. The center of area defuzzification method is selected to convert the degrees of membership of the output linguistic into precise displacement. A schematic diagram of the motion control strategy for the robot is illustrated in Fig. 7, which indicates that the fuzzy logic controller is employed to regulate the position. Since the spine is a complex and important structure, and the artificial spine surgery is often high-risk, this study takes the robot-assisted spine surgery as an example. In Fig. 8, a 3-axis linear servo Cartesian robot is used to perform the surgical milling. Each driving unit of the robot includes a servo motor and an optical encoder. The motion of the robot arm is controlled by the self-developed multi-axis motion controller based on a digital signal processor (DSP) (TMS320C6713B, Texas Instruments, United States). Control signals transmitted by the controller are sent to three motor drivers, and then three motors are separately driven by these motor drivers to finish milling task. The position feedback device is the encoder. The end of the robot arm is mounted by SPT (GD676, B. Braun, Germany), a single axis accelerometer (352C33, PCB Piezotronics, United States) and an irrigation tube. The ultraviolet disinfection method or the medical instrument isolation cover is suited for the accelerometer. The accelerometer is mounted on the metal housing of SPT by a worm-drive jubilee clip (Fig. 8), and the clip is adjusted as tight as possible to prevent loss of high-frequency vibration information. There is no request in the location and orientation accuracy for mounting the accelerometer, and the reason is that the proposed learning mechanism and control strategy only depend on the relative values of acceleration. Authorized licensed use limited to: Auckland University of Technology. Downloaded on November 05,2020 at 14:51:39 UTC from IEEE Xplore. Restrictions apply. 1939-1412 (c) 2020 IEEE" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002611_j.robot.2021.103784-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002611_j.robot.2021.103784-Figure3-1.png", "caption": "Fig. 3. An overview of the internal components of CHAD. a. The geared DC motors, the bearing, and the cable lock. b. The PCB board. The spring contacts shown in the figure are short-circuited and connect the carbon strips to the conductive pads. The cover, the base, and some of the internal parts are 3D printed.", "texts": [ " The use of rubber bands to support finger extension is inspired by the SaeboGlove (Saebo Inc, US). To reduce the size of the actuation unit, linear actuators are used in CHAD. Three linear actuators are used to pull the five tendon cables connected to the fingers. The linear actuators consist of a 12 V brushed DC motor with a gear reduction of 100:1, which is connected to a four-start lead screw with a 1.25 mm pitch or equivalently a lead of 5 mm. A cable lock is fixed on each lead screw nut and used to connect the cables to the actuators, Fig. 3a. The cable lock slides back and forth in accordance with lead screw rotation. The specifications of the overall mechanism are: a maximum load of 40 N; an unloaded speed of 20 mm/s; a stall current of 800 mA; and a stroke length of 100 mm (experimentally measured). The lead screw, gearbox, and lead screw nut can be obtained from internal parts of an off-the-shelf linear motor (L12-100-100-12-S, Actuonix Motion Devices Inc., Canada). The power screw provides a self-locking mechanism that can maintain the gripping force without continuously powering the motors. This significantly extends the battery life of the device. A printed circuit board (PCB) was designed to hold both the mechanical and electronic components of CHAD, Fig. 3b. The main electronic components are an Arduino Nano microcontroller, motor drivers (L293D, Texas Instruments Inc., US), and a Wi-Fi module (ESP-8266-12E module). The Wi-Fi module is the primary interface for controlling the device; it creates a soft access point, acts as a user datagram protocol (UDP) server, and handles received command data for controlling the device (extend/flex fingers). The UDP protocol was chosen as it allows multiple devices to send command data simultaneously. In addition, client devices being used to control CHAD do not need to be continuously connected to the device and can connect and reconnect freely; for example, a client device is allowed to go idle for power saving when inactive", " A client device, such as a gadget or mobile application, listens to voice commands from the patient and sends UDP command messages to CHAD (see an example in movie s1). In order to obtain the position of each cable lock, we have incorporated a resistive carbon layer printed on the top side of the PCB board. Position feedback for cable locks is obtained using a voltage-divider circuit. A resistive carbon layer printed on the top side of the PCB board is incorporated. The carbon layer consists of three resistive carbon strips located directly under each lead screw, as shown in Fig. 3b. Furthermore, a voltage is applied at the two ends of each carbon strip. The position of the cable lock is obtained by measuring the voltage between the cable lock contact and one end of the carbon strip via a conductive pad located next to the carbon strip, which is connected to an analog channel of the micro-controller. The contacts of the cable lock create a connection between the conductive pad and the carbon strip at the current location of the cable lock. The voltage across the conductive pad is linearly proportional to the cable lock position" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000798_j.neucom.2019.04.056-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000798_j.neucom.2019.04.056-Figure2-1.png", "caption": "Fig. 2. A serial connection of n -links model for suspension cable.", "texts": [ " Due to the less impact on the helicopter\u2019s center of gravity by using a singlepoint suspension, thus, this kind of lifting manner is adopted in this paper. Since the payload is connected with the boundary of the suspension cable, it has p(0 , t) = 0 with respect to the local reference frame x - y . Assuming that the small segment of the initial end for suspension cable is tightly fixed to the payload for the single point hanging safety, thus, s \u2032 (0 , t) = 0 . t In order to understand more intuitively that how the provided ontrol scheme in this paper to reduce the swing of the payoad, the following analysis will be given. From Fig. 2 , the sus- ension cable is supposed to connect via a series of n -links, the rst link is connected with payload and the last link is connected ith helicopter. The m i and z i are the mass and length of link , i = 1 , 2 , . . . , n, moreover, m i = \u222b z i z i \u22121 \u03c2(z) dz when i = 1 , 2 , . . . , n \u2212 , with z 0 being the endpoint which is connected with payload, nd m n = \u222b z n z n \u22121 \u03c2(z) dz + m h ; The external disturbance acting on he link i is represented by f i ; R i \u22121 and R i are the internal tensions Table 1 System variables of the helicopter suspension cable system" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000984_s12555-017-9768-z-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000984_s12555-017-9768-z-Figure2-1.png", "caption": "Fig. 2. Schematic of four-corner leveling system.", "texts": [ " \u25a1 To verify the effectiveness of the proposed strategy in this paper, the simulations are carried out for the fourcorner leveling system of an advanced hydraulic press in [7]. The comparative results are demonstrated for three different control strategies, which are the proposed dynamic control allocation (PDCA) algorithm in this paper, the quadratic programming control allocation (QPCA) algorithm in [21], and the conventional DCS control without control allocation (WCA). 5.1. System model The four-corner leveling system shown in Fig. 2 contains a slider and four hydraulic cylinder subsystems (denoted by Sub-1, Sub-2, Sub-3 and Sub-4) with the same nominal parameters [7]. The slider is driven by the four hydraulic cylinders to keep its surface in horizon. The four cylinders are physically distributed at the four corners of the slider, and controlled independently by servo proportional valves which are connected through field-bus network. Thus the overall leveling system is a typical DCS. The linearized model of each subsystem is{ x\u0307(t) = Acx(t)+B1u(t), y(t) =Cx(t), (33) where Ac = [ \u2212Bp m , A2 m ; \u22122\u03b2e(1+\u03bb 3 v )Ap \u03bbvVt , \u22122\u03b2e(1+\u03bb 3 v )Ct\u22122\u03b2e(1+\u03bb 2 v )Kc \u03bbvVt ], B1 = [0; 2\u03b2e(1+\u03bb 3 v )KqKv \u03bbvVt ], C = [0, 1]; x(t) = [vxp(t), P(t)]T is the system state, vxp(t) and P(t) are the operating speed of piston rod and the load pressure on the hydraulic cylinder (leveling pressure), respectively; u(t) is the input voltage command of the proportional valve; Ap = (A1 +A2)/2 is the effective area of hydraulic cylinder piston; \u03bbv =A1/A2" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002033_j.amc.2020.125493-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002033_j.amc.2020.125493-Figure1-1.png", "caption": "Fig. 1. The coordinates of an spacecraft.", "texts": [ " [37] is based on the modified Rodriguez parameters. Since the first step in this paper is to plan attitude trajectory which is based on the Euler angle model, there is no need to do the model conversion, which implies that the design is simpler. Finally, by conducting a real experimental platform for quadrotor, the proposed attitude control strategy is verified by experimental results. For a spacecraft, the two coordinates of the inertial frame and the body frame are used to describe the attitude, as shown in Fig. 1 . Rotation matrix R \u2208 R 3 \u00d73 =: SO (3) can take advantage of the conversion between two different poses. To give a parametric description of spacecraft attitude, the Euler angles are often used [38] . Define = (\u03bb, \u03bc, \u03bd) T \u2208 R 3 (1) that represents the attitude of the spacecraft relative to the inertial coordinate system, i.e., the pitch angle \u03bc, the yaw angle \u03bd , and the roll angle \u03bb. Based on the Euler angles, the rotation matrix is defined as: R = [ cos \u03bc cos \u03bb sin \u03bd sin \u03bc cos \u03bb \u2212 cos \u03bb sin \u03bd cos \u03bb sin \u03bc cos \u03bb + sin \u03bd sin \u03bb cos \u03bc sin \u03bb sin \u03bd sin \u03bc sin \u03bb + cos \u03bd cos \u03bb cos \u03bd sin \u03bc sin \u03bb \u2212 sin \u03bd cos \u03bb \u2212 sin \u03bc sin \u03bd cos \u03bc cos \u03bd cos \u03bc ] " ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001738_j.mechmachtheory.2019.103629-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001738_j.mechmachtheory.2019.103629-Figure6-1.png", "caption": "Fig. 6. Workspace of the manipulator under study.", "texts": [ " When this happens, all links of the rhombus becomes parallel to each other and the rhombus changes it shape and becomes a line ( Fig. 5 (c) and Fig. 5 (d)). The reachable workspace of a manipulator is the region in the space that every point can be reached by the manipulated platform in at least one orientation [45] . It consists of the points that the inverse kinematic problem has solution. However, a workspace that includes singularities or a workspace position that can only be reached with some orientation is typically not useful for a robot manipulator. Fig. 6 , illustrates the largest singularity-free box-shaped workspace that can be reached with any orientation of tool. In this figure, h is the distance between the center of the inner guide-ways, n is the thickness of the guide-ways, and t is the distance between the centre of the end-effector and the outer part of the purple tool section. Having computed all the kinematic relations, in this section, we proceed to establish the dynamic model of the manipulator under study. In this paper, the dynamic model is formulated by means of d\u2019Alembert\u2019s form of the principle of virtual work" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000614_1.4035407-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000614_1.4035407-Figure7-1.png", "caption": "Fig. 7 Trajectories of the cart for \u00f0ux \u00f00\u00de;x\u00f00\u00de\u00de5 \u00f00;5:6\u00de (shown purple) and \u00f0ux \u00f00\u00de;x\u00f00\u00de\u00de5 \u00f00; 5:7\u00de (shown red). In both cases, the no-slip trajectory confined to the plane uy 5 0 is shown in green.", "texts": [ " Suppose two trajectories with initial conditions, \u00f0ux\u00f00\u00de;x\u00f00\u00de\u00de 2 S1 and \u00f0ux\u00f00\u00de \u00fe dux;x\u00f00\u00de \u00fe dx\u00de 2 S1, respectively, transition into the set S2. In each case, when the cart ceases to slip, the two trajectories are still close to each other but the energy associated with one trajectory could be less than Ec and the energy associated could be greater than Ec. In other words, the trajectories are mapped to distinct sets, S1 and S3, respectively. The velocities for two such samples nearby initial conditions are shown in Fig. 7. The stick\u2013slip behavior for the two trajectories in Fig. 7 is similar initially. The first transition is from S1 ! S2. But one of the trajectories (shown in red in Fig. 7) transitions to set S3 and behaves as a nonholonomic system without further transitions to the slip phase. On the other hand, the other trajectory (shown in purple) transitions to set S5 and ceases to slip temporarily. This trajectory then makes a transition to set S2 again and once again slips. After a series of such transitions, the trajectory eventually maps to the set S3 and ceases to slip without further transitions. The transitions in each case are S1 ! S2 ! S3 and S1 ! S2 ! S5 ! S2 ! S1 " ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000457_1.4033888-Figure9-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000457_1.4033888-Figure9-1.png", "caption": "Fig. 9 Film boundary conditions for the Reynolds equation and the energy equation", "texts": [ " The 3D oil temperature distribution is calculated by solving the energy equation in the following equation, where u and w are the oil velocity components in the circumferential and axial directions: qc u @T @x \u00fe w @T @z \u00bc k @2T @x2 \u00fe @ 2T @y2 \u00fe @ 2T @z2 ! \u00fe l @u @y 2 \u00fe @w @y 2 \" # (16) The FEM is employed to solve the Reynolds equation and the energy equation with 2D triangle elements for the former and 3D eight-node isoparametric elements for the latter. To suppress the spatial oscillation caused by the convection term on the left side of Eq. (16), the quadratic up-winding scheme [32] is utilized. The boundary conditions for the pressure and temperature distribution for both equations are shown in Fig. 9. 6.1 Fully Nonlinear Double Overhung Method. The ME prediction comprises three coupled subproblems. (1) The transient rotor and bearing dynamics and the transient lubricant temperature distribution. In the algorithm, one single orbit will be divided into dozens of segments and within each segment, the Reynolds equation (15) will be utilized for hydrodynamic pressure calculation; Eqs. (11)\u2013(14) will be used for transient rotordynamics prediction; Journal of Tribology JANUARY 2017, Vol. 139 / 011705-5 Downloaded From: http://tribology" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000718_978-1-4471-7278-9-Figure16.15-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000718_978-1-4471-7278-9-Figure16.15-1.png", "caption": "Fig. 16.15 The robot Tessellator for tile re-waterproofing and its workspace", "texts": [ " additional weight might cause tiles to bend and fall off as well as reduce flight efficiency. Also frozen water in space might cause damage to the silica-based tiles, yielding a loss in thermal protection. While substantial productivity gains were expected through the use of the autonomous re-waterproofing robot, an efficient work strategy both within the robot workspace and over the entire Orbiter bottom was necessary. For example, a one-second improvement in title processing time would yield about a four-hour reduction in total re-waterproofing time (Fig. 16.15). The robot Tessellator contains a mobile base with omni-directional wheels, which are small wheels incorporated into the large wheels and oriented at a 45\u0131 angle. Omni-directional wheels allow the robot to rotate around its own center without extra space for rotation. The objective was to determine the effective shape of the workspace according to the structure of the robot manipulator, to determine the layout for the work-area, the minimal number of the robot workspaces and their layout, as well as the \u201csatisfactory\u201d route for the robot base movements, i" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002437_s00366-020-01236-z-Figure10-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002437_s00366-020-01236-z-Figure10-1.png", "caption": "Fig. 10 Schematic of the 582- bar tower truss", "texts": [ "95% and the maximum inter-story drift is 47.92. Table 8 Performance comparison for the 3-bay 24-story frame problem Element group Optimal W-shaped sections SBO [36] CS [35] TLBO [35] WEO [35] CBO [23] ECBO [23] MWQI-CBO Best weight (lb) 202,422 202,482 202,626 203,058 215,874 201,618 201,906 Average optimized weight (lb) 209,560 230,342 218,853 222,880 225,071 209,644 206,025 Standard deviation on average weight (lb) 7052 65,703 37,979 66,839 N/A N/A 3202 1 3 The 582-bar tower truss is schematized in Fig.\u00a010. The members are divided into 32 groups, because of structural symmetry. A single-load case is considered consisting of lateral loads of 1.12\u00a0kips (5.0\u00a0kN) applied in both x- and y-directions and vertical loads of \u2212\u00a06.74\u00a0kips (\u2212\u00a030\u00a0kN) applied in z-direction to all free nodes of the tower. Crosssectional areas of elements are selected from a discrete list of W-shaped standard steel sections based on area and radii of gyration properties. Cross-sectional areas of elements can vary between 6.16 and 215\u00a0in2 (i" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002212_tie.2020.3031535-Figure36-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002212_tie.2020.3031535-Figure36-1.png", "caption": "Fig. 36. 12S22P ST2 prototype machine.", "texts": [ " However, if the demagnetizing current is larger than 4pu, the output torque decreases significantly. Fig. 35 shows that the torque decreases by 63% after demagnetization under Id=-6pu at 200\u00b0C. For the ST2, even though the PM corners are demagnetized, the output torque is not affected. The PM magnetization length is decreased due to demagnetization, but it can still maintain its original magnetic path at rated condition. Hence, the ST2 exhibits much better performances than the SPMM under the same demagnetization fault. VII. EXPERIMENT VALIDATION A 12S22P ST2 prototype, as shown in Fig. 36, is manufactured to verify the previous analyses. Fig. 37 shows that the measured back EMF is around 11% lower than the 2D FEA calculated one due to the end effect. It can also be observed that measured one is close to the 3D FEA calculated one. The measured back EMF is around 5% lower than the 3D FEA calculated one. The static torque versus rotor position is measured by injecting dc current (Ia=-2Ib=-2Ic=Irated) into the three phase windings. The torques show similar tendency with the back EMFs, as shown in Figs" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002356_978-3-030-50460-1-Figure16.11-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002356_978-3-030-50460-1-Figure16.11-1.png", "caption": "Fig. 16.11 ANSYS model of horizontal cylindrical shell partially filled with fluid (d/2r = 0.50).", "texts": [ " It must also be realized that the fundamental in-vacuo natural frequency does not correspond to the mode shape with the lowest number of waves around the circumference (n = 2). The order of modes depends on the internal strain energy and geometrical characteristics of the cylindrical shell under study. In the second part of the analysis, the hydroelastic vibration characteristics of the cylindrical shell are investigated for three different filling depth-to-diameter ratios, d/2r = 0.2, 0.5 and 0.8. In ANSYS, the cylindrical shell is discretized with four-node quadrilateral SHELL181 elements, and fluid is modeled with FLUID30 elements (see Fig. 16.11). 286 Ardic, Yildizdag, Ergin In order to check the convergence of the wet natural frequencies of cylindrical shell, for each filling ratio, three different idealizations are adopted, and the results are compared with those obtained by ANSYS. The number of hydrodynamic panels (boundary elements) over the wetted surface is 1350, 2195 and 3190 for d/2r = 0.20; 952, 1710 and 3752 for d/2r = 0.50; and 1360, 2430 and 5360 for d/2r = 0.80, respectively. As can be seen from Table 16.5, for all the filling ratios, results exhibit monotonic convergence, and the differences between the results of the 2nd and 3rd idealizations are negligibly small" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001790_lcsys.2020.2972908-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001790_lcsys.2020.2972908-Figure6-1.png", "caption": "Fig. 6. Measurement of electrically conductive part. (a) Measurement conditions. (b) Clamp fastening plates.", "texts": [ " Hereinafter, the zone measuring an electric potential distribution is referred as the \u201czone measuring potentials.\u201d The area of the zone is 110 mm by 35 mm. The lid measuring potentials overlaps the four flanges of the metal box shown in Fig. 4. One flange is metal and the other flanges are insulated. The width, length, and thickness of the metal flange are 100, 30, and 1.2 mm, respectively. As shown in Fig. 5, the lid measuring potentials overlaps the metal box. When a voltage is applied to the lid and the box as shown in Fig. 5, a current can flow only through the metal flange. Fig. 6(a) shows measurement conditions. In case A, a 1 kg weight is on the lid. In case B, the lid and the metal flange are fastened by two clamps shown in Fig. 6(b). Fig. 7 is a schematic illustration of the electric potential distribution measurement method. The terminals of the lead wires attached to the grids are connected to relays, which are controlled by a personal computer (PC), and the connection between an amplifier and the grid is changed by the relays. The electric potential between a reference for potentials and the grid is amplified and input to an analogue digital converter in the PC. Based on the input data, the PC displays the potential distribution of the grids. An arbitrary grid is assigned as the reference for potentials. A dc current (10 A) is applied to the lid and the metal flange as shown in Fig. 6(a). Figs. 8 and 9 show the observed electric potential distribution for the cases A and B. On the basis of these distributions, current density distributions are calculated by (9). Figs. 10 and 11 show the distribution of the extracted current density Ie/(\u0394x\u0394y) (\u03c3 of (9) is the electric conductivity of steel [104 \u03a9\u00b7mm)] and electrically conductive parts. The x and y coordinates of Figs. 8\u201311 are shown in Fig. 3. The distributions of the extracted current density show that current flows through a comparatively small area (about 10 \u00d7 10 mm2)" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000567_s00339-016-0480-2-Figure10-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000567_s00339-016-0480-2-Figure10-1.png", "caption": "Fig. 10 Thin-walled metal structure produced by hybridizing fusedcoating process and subtractive manufacturing technologies", "texts": [ " Another consideration is range of applicable materials is limited presently. However, this process is especially suitable for some casting alloys, such as aluminium alloys and magnesium alloys. The layered nature of the process produces the ripples on part surfaces. The surface morphology of the sample could be improved by a milling process, and the milling depth changes from 400 to 150 lm (from up to down). The thinwalled metal structure produced by the combination of fused-coating process and milling is shown in Fig. 10. It could be seen from Fig. 10 that when the milling depth reaches 400 lm, the surface roughness of the samples can be improved significantly. The fusion zone between the layers cannot always be seen clearly, and no obvious metallurgical or forming defects were observed. A new metal 3D printing process, called metal fusedcoating additive manufacturing (MFCAM), was proposed in this paper, and the forming mechanism of the process was analysed by numerical simulations. A number of single-track multi-layer specimens of Sn\u2013Pb alloy and 7075 Al alloy were prepared by MFCAM" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001846_j.mechmachtheory.2020.103877-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001846_j.mechmachtheory.2020.103877-Figure2-1.png", "caption": "Fig. 2. Schematic picture of the variator and related geometrical quantities for double-idler (a) and single-idler (b) geometries.", "texts": [ " Hereafter, the theory is derived for the architecture comprising two inner idlers (double-idler architecture), and then it will be adapted to the single-idler architecture. The mechanism includes the following main parts: one input and one output disc; a number N of equal spherical rollers of radius r S , the axes of which are supported by a carrier that in the present analysis is held stationary; two equal idle discs (idlers), radially and axially supported by a common shaft which is free to slide in the axial direction. Fig. 2 shows the geometrical quantities of the mechanism. Here the model is derived in the general case of unequal input and output discs. Thus, the quantities r 0 and r 2 are the radial distances of the points of contact between the roller and input and output discs respectively, from the axis of the device (in short, the input and output discs radii). Analogously, r a and r b are the radial distances of the points of contact between the roller and the two idlers from the main axis of the device (in short: idlers radii). In this analysis, the idlers are considered equal, i.e r a = r b . Input and output disc radii r 0 and r 2 and also the idlers radius r i are constant quantities because the roller-discs and the roller-idlers contact points do not change location, even when the speed ratio varies. For the given sizes of the idlers, input and output discs, and the roller, three constant angles are defined, namely \u03b80 , \u03b82 and \u03b1, as shown in Fig. 2 . The tilt angle of the roller axis \u03b3 is the one formed between the roller spin axis and the axial direction (horizontal in the Fig. 2 ): tilt angle can be varied to change the speed ratio. The distances between the spin axis of the roller and all the points of contact between the roller and the input, the output and the idlers (namely, r S 0 , r S 2 , r Sa and r Sb ) are variable quantities with \u03b3 . In particular, from the Fig. 2 it can be found that: r S0 = r S cos (\u03b80 + \u03b3 ) r Sa = r S cos (\u03b1 \u2212 \u03b3 ) r Sb = r S cos (\u03b1 + \u03b3 ) r S2 = r S cos (\u03b82 \u2212 \u03b3 ) (1) Similarly, r 0 , r 2 , r i can be also written as: r 0 = r 1 + r S + r S cos \u03b80 = r S [1 + k + cos \u03b80 ] r a = r b = r 1 + r S \u2212 r S cos \u03b1 = r S [1 + k \u2212 cos \u03b1] r 2 = r 1 + r S + r S cos \u03b82 = r S [1 + k + cos \u03b82 ] (2) where r 1 is the distance between the point of the roller closest to the axis of the device (see Fig. 2 ) and the axis itself, and k = r 1 /r S is the aspect ratio. To complete the geometrical description of the device, we here define the quantities r 02 , r a 2 , r b 2 and r 22 , which are the second principal radii of curvature along tangential direction for the points of contact, respectively of: input disc, the idlers, and output disc. In the case of perfect rolling contact between discs and rollers, given the angular speed of the input disc \u03c9 0 , the following kinematical relations hold: \u03c9 0 r 0 = \u03c9 id S r S0 \u03c9 id 2 r 2 = \u03c9 id S r S2 (3) where \u03c9 0 , \u03c9 2 , \u03c9 S are the angular speeds of the input disc, the output disc and the rollers, respectively, and the superscript id means ideal, i", " (1) \u2013(3) gives the ideal speed ratio \u03c4 ID as a function of the roller axis tilt angle \u03b3 : \u03c4ID = \u03c9 id 2 \u03c9 0 = r S2 r 0 r S0 r 2 = cos (\u03b82 \u2212 \u03b3 )[1 + k + cos (\u03b80 )] cos (\u03b80 + \u03b3 )[1 + k + cos (\u03b82 )] = \u03c40 cos (\u03b82 \u2212 \u03b3 ) cos (\u03b80 + \u03b3 ) (4) where \u03c40 = [1 + k + cos (\u03b80 )] / [1 + k + cos (\u03b82 )] is a constant quantity for a given geometry. If \u03b80 = \u03b82 = \u03b8, i.e. input and output discs are equal, then the ideal speed ratio is: \u03c4ID = cos (\u03b8 \u2212 \u03b3 ) cos (\u03b8 + \u03b3 ) (5) Because of geometrical constraint, the tilt angle has a range of admissible values. Specifically, referring to Fig. 2 it follows that: \u03b3min = max [ \u2212\u03b80 , (\u03b82 \u2212 \u03c0/ 2)] \u03b3max = min [ \u03b82 , (\u03c0/ 2 \u2212 \u03b80 )] (6) with \u03b3 min \u2264 \u03b3 \u2264 \u03b3 max . The two limiting values are derived from purely geometrical constraints and it can be noted that the limit values of \u03b3 are those making \u03c4ID = 0 or \u03c4ID = \u221e . In practice, it will also be necessary to consider the actual encumbrance of mechanical parts, like the actuation system, to define the actual limits of \u03b3 . To represent the ideal speed ratio \u03c4 ID as a function of the tilt angle \u03b3 ( Fig", " A speed efficiency \u03b7speed is thus defined: \u03b7speed = \u03c4R \u03c4ID = (1 \u2212 C R 0 )(1 \u2212 C R 2 ) (9) For a given contact pair, the relative angular velocity vector of the two contacting bodies has a spin component and a tangential component. The former is obtained by projecting the relative angular velocity vector onto a direction orthogonal to the contact surfaces at the contact spot. Spin velocity causes losses in torque transmission and can generate considerable side forces on the rotating bodies [16] . By Fig. 2 , the relevant relative angular velocity vectors can be evaluated as follows: \u03c9 S0 = \u03c9 S s \u2212 \u03c9 0 j = \u03c9 S ( cos \u03b3 j \u2212 sin \u03b3 k ) \u2212 \u03c9 0 j \u03c9 Sa = \u03c9 S s + \u03c9 a j = \u03c9 S ( cos \u03b3 j \u2212 sin \u03b3 k ) + \u03c9 a j \u03c9 Sb = \u03c9 S s + \u03c9 b j = \u03c9 S ( cos \u03b3 j \u2212 sin \u03b3 k ) + \u03c9 b j \u03c9 S2 = \u03c9 S s \u2212 \u03c9 2 j = \u03c9 S ( cos \u03b3 j \u2212 sin \u03b3 k ) \u2212 \u03c9 2 j (10) where the unit vectors are those defined in Fig. 4 . The spin velocity of two bodies in contact is the component of the relative angular velocity vector onto the normal direction evaluated at the contact point", " The analysis on the power flow between the roller and the inner disc leads to the conclusion that, at least when the inner disc bearing losses are negligible, there is a certain amount of creep which is due to the axial component of spin. Therefore, a geometry of the contact between the roller and the inner disc which vanishes the axial component of the spin could be much better in terms of power dissipation and efficiency, since it will necessarily reduce the creep too. Thus, we analyzed the geometry shown in Fig. 2 b, where the inner discs are replaced by one single annulus giving radial support to the rollers. In this conditions, there is already a spin torque and, consequently, a spin dissipation at the contact between the roller and the inner annulus; however, the creep is expected to be limited since the spin has no axial component on the inner annulus. Further, also bearing losses on inner annulus are much smaller since there is no radial or axial load to be supported by bearings, and the loss is only due to bearing rotational speed" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002120_j.ijfatigue.2020.105894-Figure8-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002120_j.ijfatigue.2020.105894-Figure8-1.png", "caption": "Fig. 8. (a) Boundary conditions and loading for a fatigue model analyzed using ABAQUS for static stress/strain, and (b) the meshed model using tetrahedron elements. The mesh is finer near the surface incorporated pore.", "texts": [ " The plane with the shortest life was defined as the plane of crack initiation. This algorithm was recommended for most of the metals in FE_SAFE, mostly for ductile metals. For brittle metals maximum principal strain algorithm is recommended. All parameters are needed for the simulations to be conducted. And FE-SAFE was using tensorial stresss. The boundary conditions were set such that one end of the fatigue model was restricted at all degrees of freedom (U1 = U2 = U3 = UR1 = UR2 = UR3 = 0) and a uniform tensile loading of 400 MPa was set at the other end as shown in Fig. 8a. The meshed model shown in Fig. 8b is the assigned mesh by ABAQUS with the auto-refined mesh near the surface pore region. Tetrahedron elements were used to mesh the fatigue test samples with quadratic shape functions. The current FEM analysis has focused on the trend analysis. Further refining of the mesh may increase the accuracy of the results. Overall, the printed EH36 steel samples demonstrate a porosity dominated fatigue behavior. The samples printed at higher scanning speed show larger sized and amount of lack of fusion (LOF) resulted from deficit laser penetration, which has been reported elsewhere for the SLMed EH36 steel [24]" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000843_s11661-019-05309-7-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000843_s11661-019-05309-7-Figure2-1.png", "caption": "Fig. 2\u2014Schematics of an as-built surface comprising a series of valleys with depth Rvi and spacing 2ki and an idealization of s surface notch in terms of valley depth and spacing: (a) a series of surface notches in a rough surface comprising surface valleys and peaks, (b) an idealization of a surface notch in terms of valley depth (Rvi) and valley spacing (2ki).", "texts": [ "[24] Figure 1(a) illustrates the fabrication of an AM part by layer-by-layer deposition and solidification of powders by a laser or electron beam. As shown in Figure 1(a), the scan direction is in the horizontal (X) direction, while the build direction is in the vertical (Y) direction. The orientation of the grain structure of the heat-treated AM part is elongated in the scan (X) direction, as schematically shown in Figure 1(a) and demonstrated in the micrograph shown in Figure 1(b).[25] Layer-by-layer deposition is expected to result in a rough surface normal to the scan (X) direction, which is shown schematically in Figure 2(a). It is envisioned that the rough surface comprises a series of valleys and peaks. METALLURGICAL AND MATERIALS TRANSACTIONS A The valley depth is denoted as Rvi and the valley spacing is designated as 2ki, as shown in Figure 2(a). The valley spacing may also be considered as the valley width. The two terms will be treated as equivalent in this paper. It is further stipulated that the valley depth and spacing from a surface notch consisted of an arc height (Rvi) and a chord length (2ki) of a circular segment of an elliptical notch with a radius q, as shown in Figure 2(b). Based on this 2D valley geometry, the stress concentration factor, kt, can be obtained as[26,27] kt \u00bc 1\u00fe 2 ffiffiffiffiffiffi Rvi q s ; \u00bd1 which leads to kt \u00bc 1\u00fe 2 2 1\u00fe ki Rvi 2 2 6 4 3 7 5 1=2 \u00bd2 when the notch radius q is expressed in terms of Rvi and ki is based on the geometrical relations between chord length, arc height, and the radius of a circular segment. It is anticipated that the ratio of valley depth to spacing (Rvi/2ki) can be measured by surface profilometry. Once obtained, the surface roughness data can be used to compute the stress concentration factors associated with surface notches of an AM part" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000069_012061-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000069_012061-Figure6-1.png", "caption": "Figure 6. Test specimen under compressive test", "texts": [ "7 x 12.7 x 25.4 mm. 3. Results & Discussion- The results obtained through the experimental process were analyzed with the help of ANOVA test. The Tensile and compressive strength values for the specimen based on the Taguchi experimental layout is shown in Table 3. Sample Layer thickness Infill % Print Speed Compressive Strength (Mpa) Tensile Strength (Mpa) NFEST IOP Conf. Series: Journal of Physics: Conf. Series 1240 (2019) 012061 IOP Publishing doi:10.1088/1742-6596/1240/1/012061 As we see from Figure 6, the highest Tensile strength, i.e., 30 MPa occurs for the minimum layer thickness, i.e., 0.1mm. And lowest Tensile strength, i.e., 18 MPa occurs for the maximum layer thickness, i.e., 0.3 mm. Primarily reason for tensile test failure is principal stress. Since a low value of layer thickness imparts more compact specimen. Layers are closely deposited over each other. This help in better bonding between the layers. It results in an increase of tensile strength. In the same way, a higher value of layer thickness, affect the compactness of the specimen" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000018_s2301385020500016-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000018_s2301385020500016-Figure2-1.png", "caption": "Fig. 2. Two collision avoidance modes. White regions denote the convex set of allowed positions.", "texts": [ " The constraint marked by n\u0302 r;head ijj is replaced by a linear constraint derived from the Voronoi Cell and not necessarily at the same position as that shown. Using this method, the constraint obtained is on position space and not limited to a time horizon. We coordinate agent movement to enable them to reach their assigned targets by careful selection of the linearization. Selection is done immediately after information exchange between the communicating agent pair and is constant until the next communication instance. The linearization is split into two modes as shown in Fig. 2. The first and default mode, called the Buffered Voronoi Partition (BVP) mode, consists of the head-on Voronoi Cell-based linear constraint and repulsion. The repulsion is used to expand the space between agents and ensures that the agent remains within the feasible set when transitioning between modes. The second mode, called the Collision Cone Partition (CCP) mode, consists of either of the clockwise or counterclockwise constraints from Fig. 1(b). While the first mode ensures that agents can approach each other up to some distance without being deadlocked, the second is used to enable them to pass each other as close as rs" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002544_012013-Figure6-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002544_012013-Figure6-1.png", "caption": "Figure 6. Bearing specimens used for wear (left) and fatigue tests (right)", "texts": [ " The fatigue experiments were performed with centrifugally casted sliding bearings (Gleitlagertechnik Wei\u00dfbacher, Alpen, Germany) with 2 mm thick babbitt sliding layer made of SnSb8Cu4. The shaft was made from temper-hardened 42CrMo4 with and hardness of 56 HRC. The bearing diameter and width were 70.01 mm and 28 mm. The shaft diameter was manufactured to 69.96 mm, which results in a relative clearance of 2 \u2030. The shaft and bearing were cleaned with Acetone prior to each test. In this study, a commercial turbine oil (Teresstic\u2122 T 100, Exxon Mobil) was used. The specimens are exemplarily shown in Figure 6. 19th Drive Train Technology Conference (ATK 2021) IOP Conf. Series: Materials Science and Engineering 1097 (2021) 012013 IOP Publishing doi:10.1088/1757-899X/1097/1/012013 During the tests, AE measurements were conducted with a NANO30-sensor (MISTRAS, Princeton Jct, NJ, USA). The sensor has a resonance frequency of 300 kHz and a bandwidth of 125 kHz \u2013 750 kHz. To ensure a coupling between the sensor head and the bearing housing, thread seal paste (Neo-Fermit Universal 2000/1, Fermit, Vettelscho\u00df, Germany) was used" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000451_j.robot.2016.05.012-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000451_j.robot.2016.05.012-Figure4-1.png", "caption": "Fig. 4. General planar robot on irregular terrain.", "texts": [ " However, if there are two such axes with the same set of normal foot forces at the same time, the measure of MFFSM could be different for them since bPi\u03b2 chi\u03b3 varies. In this case, a lower MFFSM indicates a higher tip-over potential about the tip-over axis and vice versa. In a very specific case, when the MFFSM is the same for two axes, tip over will occur about the single foot contact point with maximum foot force instead of an axis. One example is a tripod robot with a lateral force applied along the bisector of the supporting isosceles triangle. Consider a planar robot on irregular terrain as shown in Fig. 4. Eq. (7) is applicable with minor changes to the parameter definitions. In the planar case, Pi = Ci \u2212 Cg is the ith tip over vector with magnitude Pi = \u2225Pi\u2225, and hi is the height of the CG with respect to ith ground contact point in the gravity direction. In Eq. (7), the choice of subscript i, i \u2208 1, . . . , n, is based on the foot with the largest measured normal foot force. For example, if f1 > f2, then i = 1, if f1 < f2, then i = 2, and if f1 = f2, then i could be either 1 or 2. In the latter case, FFSM is 1 butMFFSM for i = 1 and i = 2 could be different since Pi\u03b2 hi\u03b3 could be different" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001981_s10846-020-01213-0-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001981_s10846-020-01213-0-Figure3-1.png", "caption": "Fig. 3 Geometric parameter of the object", "texts": [ " It is worth remarking that, assuming that a rigid object is tightly grasped by the N vehicles, ro k and Ro ek are constant quantities describing the kinematics of the grasp. Thus, the above equations define a kinematic model of the grasp, i.e., a set of mechanical constraints (closedchain constraints) on the variables describing the motion of each end effector Tk = {xek , x\u0307ek , x\u0308ek } (k = 1, 2, . . . , N) and those describing the object\u2019s motion To = {xo, x\u0307o, x\u0308o}. It is assumed that the dynamic parameters of the grasped object are not known by the aerial vehicles: thus, a procedure to cooperatively estimate them needs to be devised. By referring to Fig. 3, the parameters to be estimated are: \u2013 the mass mo; \u2013 the position of the center of mass, Oo, with respect to the contact points; \u2013 the inertia matrix I o. Following the approach proposed in [9], the estimation of the unknown parameters will be achieved on the basis of the data collected during a preliminary identification stage, in which the aerial vehicles perform simple motions before starting the execution of the assigned task. It is assumed that the parameter identification stage occurs in the free space, i", " The mapping from force and moment, expressed in the rotor frame, to the propellers\u2019 thrusts \u03c3 can be written as [30] [ \u2225\u2225f vk \u2225\u2225 (RkR k p)TT (\u03c6k) \u22121\u03bcvk ] = \u23a1 \u23a2\u23a2\u23a3 1 1 1 1 l 0 \u2212l 0 0 \u2212l 0 l c \u2212c c \u2212c \u23a4 \u23a5\u23a5\u23a6 \u23a1 \u23a2\u23a2\u23a3 \u03c31 \u03c32 \u03c33 \u03c34 \u23a4 \u23a5\u23a5\u23a6 = N\u03c3 , (36) where l > 0 is the distance from each motor to the origin of rotor frame, c = \u03b3d/\u03b3t , with \u03b3d and \u03b3t the drag and thrust coefficients, respectively. Then, the rotor thrusts are computed as \u03c3 = N\u22121 [ \u2225\u2225f vk \u2225\u2225 (RkR k p)TT (\u03c6k) \u22121\u03bcvk ] . (37) The proposed cooperative control scheme has been tested via realistic simulations by using the MATLAB\u00a9 SimMechanics environment. The geometric and dynamic parameters of the ODQuad have been set via CAD software and are reported in Table 1. A team of three robots is commanded to move the object depicted in Fig. 3, whose dynamic parameters are supposed to be unknown to the robots, along a planned path. The control gains are reported in Table 2. In order to simulate a realistic scenario the following assumptions have been done: \u2013 Measurements of position and orientation of the aerial robots are affected by noise with mean 1 \u00b7 10\u22122 m for position and 0 rad for the orientation, and a standard deviation of 10\u22123 m for position and 10\u22123 rad for the orientation. \u2013 The platform linear velocity are obtained via derivative filter, while angular velocity is assumed measurable with a 0 mean measurement noise and standard deviation 10\u22123 rad/s" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002356_978-3-030-50460-1-Figure15.2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002356_978-3-030-50460-1-Figure15.2-1.png", "caption": "Fig. 15.2 Corrugated laminate reference coordinates. Source: Kress and Filipovic (2019)", "texts": [ " Discussion and conclusion are placed in Sections 15.12 and 15.13, respectively. In oder to be able to predict the flat-laminate design for creating a corrugated laminate by the thermal curvature effect, where corrugation amplitude and periodic length must meet specified values, it is necessary to consider the classical theory of laminated plates (CTLP) and some geometric relations for describing the geometry of corrugation shapes consisting of circular sections. 15 Manufacturing and Morphing Behavior of High-Amplitude Corrugated Laminates 237 Fig. 15.2 indicates the reference coordinates in which the corrugated laminate is described and the curvilinear coordinates along the laminated sheet mid-plane. The relation between circular-segments radius R, corrugation amplitude c, and periodic length of the corrugation pattern P is given with (Kress and Winkler, 2010): R = 16c2 + P 2 32c . (15.1) This equation is rearranged to calculate the periodic length of one corrugationpattern unit cell for given reference-shape curvature R and the ratio c\u0304 that is the corrugation amplitude c normalized with respect to the periodic length P : P = R 32c\u0304 16c\u03042 + 1 ; c\u0304 = c P , (15", " Note that extreme morphing stretches, where the corrugated laminate tends to become flat, will create less bending curvatures for high-amplitude corrugations than for corrugation shapes composed of semi-circles (c\u0304 = 0.25). The condition, that the laminate with sheet thickness t must not penetrate itself, gives the upper bound of the corrugation amplitude c (Filipovic and Kress, 2018): 238 Kress, Filipovic 0 \u2264 c \u2264 P \u2212 t+ \u221a (P \u2212 t) 2 \u2212 P 2 4 2 . (15.3) The curved-sheet length Ls, covering one period of corrugation, is given with (Kress and Winkler, 2010): Ls = 4\u03c80R , (15.4) where the opening angle \u03c80 is defined in Fig. 15.2 and is calculated with (Kress and Filipovic, 2019): \u03c80 = acos ( 1\u2212 c R ) . (15.5) If the corrugated-shape unit cell with undeformed length P is stretched to flatness, its length will increase from P to Ls. This thought allows the estimate of maximum morphing capacity in terms of stretch \u03bb (Kress and Filipovic, 2019): \u03bblim = Ls P , (15.6) where Fig. 15.4(b) illustrates that the morphing stretch limit increases progressively 15 Manufacturing and Morphing Behavior of High-Amplitude Corrugated Laminates 239 with normalized corrugation amplitude c\u0304", " The geometric-unit-cell is point symmetric with respect to its center point, and the two circular segments are in themselves symmetric with respect to vertical lines. Therefore, the point marked with a solid circle cannot rotate. Because of these circumstances, and even though the stacking sequence of the first circular section, say [0n/90m], is reversed in the second one, the structural behavior of the geometric unit cell is fully de- 252 Kress, Filipovic scribed by the mechanical unit cell shown in the sketch to the right of the figure. The relation between stress resultants and deformations is described in the local system shown in Fig. 15.2, where the coordinate mapping between the moving trihedral x \u03be2 \u03be3 and the reference system xyz is given with\u23a7\u23a8\u23a9 x \u03be2 \u03be3 \u23ab\u23ac\u23ad = \u23a1\u23a3 1 0 0 0 cos \u03b8 sin \u03b8 0 \u2212 sin \u03b8 cos \u03b8 \u23a4\u23a6\u23a7\u23a8\u23a9 x y z \u23ab\u23ac\u23ad . (15.21) For writing convenience of subscripts, we prefer to replace the symbol \u03be2 with s. Fig. 15.2 indicates points in the reference configuration, y(0)(\u03be2) = Rsin\u0394\u03c8 ; z(0)(\u03be2) = R [1\u2212 cos\u0394\u03c8]\u2212 c , (15.22) where \u0394\u03c8 = \u03c8 \u2212 \u03c80 = \u03be2 R \u2212 \u03c80 = \u03ba0\u03be2 \u2212 \u03c80 , (15.23) Point B in the reference configuration shown in Fig. 15.17 is given with: y (0) B = P 4 ; z (0) B = \u2212c . (15.24) Cylindrical bending implies that curvature deformations vanish, \u03b51x = \u03b51xs = 0. We also assume that the midplane strain transverse to the corrugations and the shearing 15 Manufacturing and Morphing Behavior of High-Amplitude Corrugated Laminates 253 strain vanish, \u03b50x = \u03b50xs = 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0003137_1.1752269-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0003137_1.1752269-Figure1-1.png", "caption": "FIG. 1. Apparatus to determine the non viscous friction of thin lubricating layers. W, steel axle; K, steel sphere; H, heating element.", "texts": [ " Both apparatus have (1) the advantage of simple equations of motion and (2) that the motion as a function of time itself indi cates by the kind of damping whether there is viscous or nonviscous friction. For the sake of simplicity both of method and design, a p.endu lum apparatus was built first, the operatIOn of which, incidentaBy, simulates the motion of the piston rings in engines. It was considered at that time (January, 1934), in accordance wi th ~he general view, that piston rings ~ere. workmg under conditions of boundary lubncatIOn for a considerable part of their path. Fig. 1 represents the friction pendulum, Fig. 2 explains a few of its essential dimensions. The suspension consists of a ground and finely polished journal which is free to oscillate o~ fo~r fixed steel balls. The support of the suspenSIOn 1S provided with an electric heater. T~e pend.ulum oscillates above a scale with a releasmg dev1ce at its end. The ends of the journal are provided with points opposite adjustable stops on the pend~lum support which permit accurate lateral adJust ment by means of a feeler gauge", " To set up the pendulum, the journal is lubri cated without particular precautions and the pendulum adjusted approximately above the zero point of the scale, carefully adjusted, how ever, to prevent it from sliding sideways. For the measurement proper the sliding surfaces are repeatedly rinsed with chemically pure benzene and rubbed with a smalI swab of cellucotton wrapped in tissue paper. Any contamination by the hands should be avoided. The hands should be carefully washed before with soap and water as in the experiments of Hardy. At the bottoms of the short supports of the steel balls there are small flat dishes (Fig. 1) from which the liquid from the cleaning process is removed by blotting with cell uco tton. After these preparations a swab of cellucotton is wetted with the test liquid, wrapped in tissue paper and, avoiding con tamination by the hands, the sliding surfaces are rubbed, and subsequently cleaned by means of a dry swab of cellucotton. In the same way the surfaces are \"lubricated\" with a freshly \"wetted\" roll of cellucotton. This final lubrica tion is best achieved when the cellucotton is so little wetted that the liquid barely penetrates the wrapping paper" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002460_lra.2021.3056349-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002460_lra.2021.3056349-Figure3-1.png", "caption": "Fig. 3. The setup for the loading test. The sample is rigidly clamped on the left, top, and right sides and under a dead load of 200 g on the bottom side.", "texts": [ " As the performance of the dielectric is improved through prestraining, we opted to displace a dead load that maintained prestrain along one direction. However, prestrain could be maintained through other techniques, such as using interpenetrating polymer networks, or IPNs [28]. To maintain prestrain, we first clamped the elastomer at four points to a rigid frame. Then, weight was applied to the bottom clamp, and the clamp was detached from the rigid frame. Finally, a set of linear stages was attached to the rigid frame to constrain the motion of the bottom clamp to one dimension. A schematic of the test setup is shown in Fig. 3. Similar to the technique used in Section III-B, a separate fiber-reinforced encapsulation layer was attached to the elastomer after CNT electrodes were applied to the dielectric layer. The fiber reinforcement was extended to both vertical clamps to transfer a majority of the deformation to the load. The materials used were the same as those used in Section IIIB, though samples were prestretched to \u03bbPS = 3.5. Before prestraining, the elastomer was laser cut into a cross shape. The electrodes were applied to take up a majority of the cross-section" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000187_s10846-019-01113-y-Figure8-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000187_s10846-019-01113-y-Figure8-1.png", "caption": "Fig. 8 a Configurations with limited difference. b Tunneling problem for configurations with larger difference", "texts": [ " However, this problem can be eliminated by limiting the difference between configurations properly. In this paper, threshold of collision margin is set for different collision situations based on the size of links, stations and obstacles. Maximum angle differences of each joint are calculated accordingly so that the displacement of any point on the manipulator during the transition is less than these thresholds if the differences of all joint angles are within their limits. It ensures that the feasibility of ending configurations can fully indicate the safety of their transition. Figure 8a demonstrates this idea. Two red spheres serve as obstacles. If the manipulator transfers from the solid configuration to the transparent one, collision will happen and can be predicted. The detailed limits of angle differences used in this paper will be list in Section 7. Yet, if joint angle differences between configurations are beyond their limits, more instances must be sampled to enforce this limitation so that collision during the transition can be predicted. Figure 8b shows an example of tunneling problem. Difference between the upper configuration and lower one is beyond limit. So the collision happened between the transfer will not be predicted unless an intermediate instance is sampled. The tunneling problem is thus eliminated. Moreover, if the collision margin of a configuration is greater than the corresponding thresholds, it indicates that the neighboring configuration within angle interval limits is also collision-free for that specific situation. The ratio between collision margin and the threshold can be recorded as supplementary collision information to considerably reduce the time-consumption and enhance efficiency" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002356_978-3-030-50460-1-Figure21.1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002356_978-3-030-50460-1-Figure21.1-1.png", "caption": "Fig. 21.1 Bearing crosssection and geometry over which flow is modelled.", "texts": [], "surrounding_texts": [ "1. Begins with writing down the Reynolds equation for one\u2019s case and formulate FEM for the differential equation. 2. Discretize the problem according to geometry. 3. Apply proper boundary conditions. 4. Apply Galerkin criteria or Variational principle to get an integral functional that has to be minimized. 5. Minimize the functional or residual in local coordinate system. 6. Get local matrix equations using fluidity matrices for individual elements. 7. Write a global equation that can be solved by an iterative procedure." ] }, { "image_filename": "designv11_5_0001062_s00202-019-00874-x-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001062_s00202-019-00874-x-Figure1-1.png", "caption": "Fig. 1 a IM two-dimensional FE model geometry and mesh, S/R = 48/40, p = 4; b Coupled circuit schematic of the IM", "texts": [ " The PWF model can be solved using any known numerical method for the integration of the set of nonlinear differential equations that describes the IM. The execution times stated in this section are based on Euler\u2019s method of integration with a preset constant time step. The program code was written and executed in MATLAB\u00ae. The FE model of the examined IM designs was built using the FLUX2D electromagnetic software package from Cedrat\u00ae [19]. Detailed machine geometry/design specifications were used to develop the FE models of the examined eight-pole IMs, as illustrated in Fig.\u00a01a. The two-dimensional FE models were developed to reduce computational time but still maintain sufficient accuracy by choosing an appropriate mesh. A pseudo-threedimensional FE model was created by coupling an external circuit to the FE model, to include the values of end effects, as shown in Fig.\u00a01b. Figure\u00a01b shows the coupled circuit with phase voltage sources, which feed each phase winding modelled in a single block called a \u2018coil conductor\u2019 that represents 16\u00a0series connected coils in each phase. An inductor is also included in each phase to take account of the value of the end-winding inductance. The total phase resistance was included in the \u2018Resistance\u2019 blocks of each phase. Rotor skew was not considered during this investigation. The FE model assumes a balanced three-phase supply, and only the fundamental phase voltage was considered in the transient simulations" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001191_j.asoc.2020.106194-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001191_j.asoc.2020.106194-Figure1-1.png", "caption": "Fig. 1. The schematic of the bilateral underwater tele-operation system.", "texts": [ " The problem formulation is briefly illustrated in Section 2. The bilateral underwater teleoperation dynamic model is introduced in Section 3. In Section 4, the bilateral impedance control of the underwater tele-operation system is explained in detail with stabilized analysis. Numerical simulations are showed in Section 5. Some conclusions are given in the conclusion part. Normally, in the underwater tele-operation system, the human operator is located in the manned submarine above the water, as shown in Fig. 1. Driving the master manipulator, the human operator may control the underwater slave manipulator under to achieve the desired operations. Normally, an umbilical cord is used to supply the power and data to the master side. The communication between the master side and the slave side is implemented by cables or wireless between the surface and the deep of the water [20]. The communication among three parts (the human operator, the master and the slave) may occur the time delay in the complex underwater environment" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000403_0954405416640171-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000403_0954405416640171-Figure3-1.png", "caption": "Figure 3. (a) Solution of the generating line and (b) determination of the initial points.", "texts": [ " They are the basic geometric parameters. In the formation theory of spherical involute tooth flank, the numerical expression to the generating line k0kt and the initial motion points are the keys to the formation of the tooth flank. Considering tooth flank characteristics of the spiral bevel gear, generating line can be presented by a certain mathematical function based on the base cone spiral curve which is an Archimedean spiral curve named equidistant spiral curve. And then, the function of base cone spiral cure is represented in Figure 3(a) FG = f \u00bdRn,bn \u00f03\u00de where bn represents spiral angle at any point on the toe and heel. It can be obtained by following formula4 bn = arc sin 1 2Rn Rn 2 +2rcRp sin b Rp 2 \u00f04\u00de where rc means the radius of the cutter head, Rn indicates cutter location and Rp signifies the cone distance in correspondence with nominal spiral angle b. Finally, as shown in Figure 3(b), base cone apex O as the ball center is made a serial of variable radius ball to not only intersect with the spiral lines at one point kn but also with the circular flank intersect at one point Kn, and the last intersection points can constitute the initial movement points. Recently, the tooth flank of spiral bevel gear is required a high precision. In the actual manufacture, it needs dedicated machine tools to complete the complex machining through repeated trial-to-error process to determine the target modification machine settings" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000286_b978-0-08-102884-1.00001-7-Figure1.13-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000286_b978-0-08-102884-1.00001-7-Figure1.13-1.png", "caption": "FIG. 1.13 Kinematic design of three jointed legs used in robotic applications. (A) Robotic leg that can be used in walking robots inspired by mammals. (B) Robotic leg that can be used in walking robots inspired by insects.", "texts": [ " Secondly, the actuators of the robot must not be damaged during impact steps or falls and must maintain stability following an impact. Finally, the actuators may need to be force controllable because the in some approaches the algorithms used for robot locomotion are force based. A large number of walking robots have been developed with two, four or six legs, as discussed by Zhou and Bi (2012) for a variety of applications ranging from pure research through to load carrying for military or humanitarian logistics requirements in areas where wheeled vehicles could not be used. Fig. 1.13 shows two possible approaches to the design of a leg. One of two possible approaches are possible dependant of the design of the joints that replicate the motion of the hip. The placement of the two coincident hip joints has the advantage of placing the respective actuators on or close to the vehicle body, so that their mass is not carried by the leg during motion. In addition, the design of the hip joint provides optimum workspace and simplifies the kinematics. There are two variations of this approach to leg design leg: mammal or insect inspiration. In the leg based on a mammal, the knee joint is placed under the hip, Fig. 1.13A, this approach is used in either biped robots or four legged running robots. In the insect-based design, the knee joint is located laterally or at a position higher than the hip, Fig. 1.13B, this design to typically used in six-legged insect-based robots. A related application is the use using powered exoskeletons as part of the treatment of patients during rehabilitation following a stroke or spinal cord injury. The exoskeletons are effectively walking robots that are designed to support the patient and aim to enhance the rehabilitation process. The current automotive market is in a state of considerable flux, with rapid advances in autonomous vehicle technology, as well as the goal of reducing our dependence on petrol and diesel" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000433_j.apm.2016.04.030-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000433_j.apm.2016.04.030-Figure2-1.png", "caption": "Fig. 2. Physical geometry for a power-law film slider bearing lubricated with a non-Newtonian Rabinowitsch fluid.", "texts": [ " \u2202 \u2202x [ h 3 \u2202 p \u2202x + 3 20 \u03b1h 5 ( \u2202 p \u2202x )3 ] + \u2202 \u2202y [ h 3 \u2202 p \u2202y + 3 20 \u03b1h 5 ( \u2202 p \u2202y )3 ] = 6 \u03b70 U \u2202h \u2202x + 12 \u03b70 \u2202h \u2202t . (19) This derived nonlinear Reynolds equation can be applied to the general research of bearing performances taking into account the effects of non-Newtonian dilatant and pseudo-plastic fluids, in which the general two-dimensional film profile is h = h (x, y, t) . In order to guide the use of the derived Reynolds equation, a power-law curved film slider bearing lubricated with nonNewtonian dilatant and pseudo-plastic fluids is illustrated in Fig. 2 . Assume the bearing is infinitely wide ( D > > B ) and the bearing is operating under steady state ( \u2202 h/\u2202 t = 0 ), then the film profile can be described by: h ( x ) = ( h A \u2212 h B ) \u00b7 ( 1 \u2212 x/B ) n + h B . (20) This film profile includes the special cases of an inclined plane bearing for the shape index n = 1 and of an parabolic film bearing for n = 2 . Since the side leakage in the y \u2212direction and squeezing effect can be neglected, the general nonNewtonian Reynolds Eq. (19) for the illustrated bearing reduces to: d dx [ h 3 dp dx + 3 20 \u03b1h 5 ( dp dx )3 ] = 6 \u03b70 U dh dx " ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001800_978-3-030-40410-9-Figure4.21-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001800_978-3-030-40410-9-Figure4.21-1.png", "caption": "Fig. 4.21 6DOF revolute joints (PUMA) robot manipulator", "texts": [], "surrounding_texts": [ "Exercises 99\nH = \u23a1 \u23a2\u23a2\u23a3 0.1112 \u22120.8074 0.5795 336.3267 \u22120.9358 0.1112 0.3346 194.1783 \u22120.3346 \u22120.5795 \u22120.7431 332.1964\n0 0 0 1\n\u23a4 \u23a5\u23a5\u23a6\nmanipulator.", "Ghafil HN, J\u00e1rmai K (2019) On performance of ED7500 and lab-volt 5150 5DOF manipulators: calibration and analysis. In: MultiScience\u2014XXXIII. MicroCAD international multidisciplinary scientific conference University of Miskolc, 23\u201324May 2019. ISBN 978-963-358-177-3. https:// doi.org/10.26649/musci.2019.042", "References 101\nGhafil HN, Kov\u00e1cs L, J\u00e1rmai K (2019) Investigating three learning algorithms of a neural networks during inverse kinematics of robots. In: Solutions for sustainable development: proceedings of the 1st international conference on engineering solutions for sustainable development (ICESSD 2019), 3\u20134 Oct 2019. CRC Press, Miskolc, Hungary, 2019, pp 33\u201340 Kuka_manual (2018) KUKA. https://www.google.hu/url?sa=t&rct=j&q=&esrc=s& source=web&cd=1&ved=0ahUKEwjuurbM6-jYAhXSZVAKHUoCC6kQFggoMAA&url= https%3A%2F%2Fwww.kuka.com%2F-%2Fmedia%2Fkuka-downloads%2Fimported% 2F6b77eecacfe542d3b736af377562ecaa%2Fdb_kr_100_2_comp_de.pdf&usg= AOvVaw1uAaR6NX0xxacpztuvEL0n. Accessed 14 Oct 2018" ] }, { "image_filename": "designv11_5_0001800_978-3-030-40410-9-Figure4.7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001800_978-3-030-40410-9-Figure4.7-1.png", "caption": "Fig. 4.7 Two-link planar manipulator", "texts": [ " The first frame in the manipulator is counted from zero, and the designer is free to choose where to assign a coordinate frame rigidly to a specific link. For a prismatic joint, the rotation angle variable should be equal to zero, and the HTM will be a function of the change in link offset (di), which we will denote as (Pi) Hi = \u23a1 \u23a2\u23a2\u23a3 1 0 0 ai 0 S\u03b1i \u2212S\u03b1i 0 0 0 1 Pi 0 0 0 1 \u23a4 \u23a5\u23a5\u23a6. (4.4a) Example 4.1 (Two-link planar robot) Estimate the forward kinematics equations for the two-link planar robot manipulator shown in Fig. 4.7. Assume that the length of link 0 and link 1 is a1 and a2, respectively, where the second link starts from the point of connection with the other link to the centre of the end-effector. Firstly, we have to assign the coordinate frames for this manipulator; the configuration in Fig. 4.7 (in a continuous line) will be chosen to be the rest configuration. The first joint, which is connected to the ground, is a revolute joint; therefore, z-direction will be in the direction of rotation. Assume the positive direction of the z-axis is out of the paper; then by the right-hand rule, x-axis and y-axis of the first coordinate can be determined. For this example, the thumb should be directed out of the paper to the reader while the four fingers directed along with link 1, and by using the explanation of the DH convention and right-hand rule, we can set coordinate frame o0x0y0z0 which is the base frame" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001800_978-3-030-40410-9-Figure5.7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001800_978-3-030-40410-9-Figure5.7-1.png", "caption": "Fig. 5.7 RRR planar manipulator", "texts": [ " functionalDerivative(cc5(3,4),th5)]; J = [Jv;Jw]; Jd = diff(J(t)); % clone JJ = J(t); % clone X = Jd*qd+JJ*q2d function As we have explained before, the Jacobian provides a mapping for the velocities of a point on the manipulator to the joint space. In the previous examples, we considered the end-effector velocity and acceleration. In fact, it is more important to describe the velocity and acceleration of different points on the manipulator rather than on its end-effector, especially when dynamic analysis is required. Consider Fig. 5.7, which shows a 3DOF revolute-joint manipulator. The centre of mass of the link 2 is denoted by oc with linear velocity v and angular velocity w. We can derive the Jacobian of the manipulator at the centre of mass of link 2 directly to represent its velocity and acceleration to the actuators at joint 0 and joint 1. Note that link 3 does not affect the velocity of the link 2 kinematically, but the reaction from there can affect the link when we consider kinetic analysis, as we will explain in Chap. 7. In this chapter, we have learned how to find the relationship among velocity components of the end-effector and velocities of the joints variables. This relationship is the Jacobian matrix which is driven from forward kinematic equations. Jacobian is very important when we would like to describe the velocities of the centre of masses of the links and end-effector in order to formulate the dynamic equations as we will in Chap. 7. Exercises 121 Exercises manipulator shown in Fig. 5.7 to the joint space. Chapter 6 Path and Trajectory Planning In the previous chapters, we have studied the geometry of the robot and the kinematic solutions associated with that geometry without giving any attention to the constraints imposed by the workspace. We have developed solutions for the forward kinematics and inverse kinematics of the robot manipulator, assuming that we have free space surrounding the robot and its possible configurations. For a workspace with constraints like other robots, equipment, workers, etc" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002210_j.robot.2020.103670-Figure17-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002210_j.robot.2020.103670-Figure17-1.png", "caption": "Fig. 17. Wheelchair and robot model.", "texts": [], "surrounding_texts": [ "w\nr b o\nw\nw p s\nw\nu t s\n4\ni s 1 H\n[Stage 2] <5> The wheelchair and the robot continue to move forward. <6> The rear wheels of the wheelchair come into contact ith the step. <7> The robot continues to push the wheelchair so that the ear wheels of the wheelchair keep in contact with the step and egin climbing. The robot continues to support the wheelchair in rder to prevent it from tipping over backward.\n<8> The wheelchair and the robot move forward, and the rear wheels of the wheelchair climb completely onto the step. The process is complete when the wheelchair\u2019s level sensor determines that the back wheels have finished climbing.\n[Stage 3] <9> The robot folds its front and rear wheel mechanism (Fig. 4) and spreads its manipulators (Fig. 10(a) and (b)). The touch sensors on the forearm links (which are covered with bumpers) detect contact with the wheelchair (Fig. 6), after which the manipulator forearm links are inserted into the stopper (between the front bars and rear bars)(Fig. 9). The wheelchair then stops, and the robot moves forward until the manipulator forearm links come into contact with the front bars of the wheelchair stopper (Fig. 11(a)\u2013(c)).\n<10> If the system detects the step height, it can then determine the angle needed to place the robot\u2019s front wheels on the step (See Section 4). The robot continues to drive forward, pushing against the wheelchair using the forearm links, until the robot\u2019s front wheels are lifted off the surface and its accelerometer system detects the inclination needed to place its front wheels on the step.\n<11> As the robot\u2019s tilt increases, the robot center of mass shifts behind the contact point between the robot middle wheels\nf\nand the ground and the robot begins to tip backward. However, the manipulator forearm links come into contact with the rear bars of wheelchair stopper, which limits the extent of rotation (Fig. 11(d)). As a result, the wheelchair prevents the robot from tipping over.\n<12> The wheelchair and the robot then move forward, and the front wheels of the robot are placed on the upper level of the step.\n[Stage 4] <13> The wheelchair and the robot continue to move forard. <14> Next, the robot\u2019s middle wheels come into contact\nith the step as the wheelchair pulls the robot forward. At this oint, the robot\u2019s rear wheels are lowered and provide additional upport as the robot\u2019s middle wheels begin to climb the step.\n<15> The wheelchair and the robot continue to move forard. The robot\u2019s rear wheels remain lowered. <16> When the robot\u2019s middle wheels have reached the\npper level of the step, the robot\u2019s level sensor detects the end of he step-climbing process (Fig. 14). At this point, both vehicles are topped, and the robot\u2019s rear wheels are folded upward (Fig. 4).\n. Geometrical and static analysis of the step-climbing process\nIn this system, the wheelchair and assistive robot are deployed n a forward-and-aft configuration when the vehicles climb a tep. The system lifts the front wheels of the wheelchair (Stage ) using the velocity differences between the connected vehicles. owever, if the wheelchair is too close to the step in Stage 1, the ront wheels will collide with the step, and a step-climbing failure", "w b t w\nt t r s a i t T a w\nt f\n4\ns\nw ( a r\np v a\nt\nill occur [23]. Similarly, it is necessary to leave a certain distance etween the robot\u2019s front wheels and the step to facilitate lifting he robot\u2019s front wheels. If the distance is too short, the robot\u2019s heels will collide with the step (Fig. 16). Since the mechanism of the robot\u2019s front wheels permits them o be folded (Fig. 4), even if the front wheels collide with the step, hey will be pushed inward, thus causing the wheelbase of the obot to shorten until the wheels are finally able to climb the tep. However, while contacts between the robot\u2019s front wheels nd the step do not always cause a step-climbing failure, there s a limit to the extent of an acceptable wheelbase transformaion. (The minimum value of WBf is 190 mm. See Appendix A, able A.1.) In addition, the robot cannot begin lifting itself if it is significant distance from the step, because it is connected to the heelchair. In this section, we show the results of a theoretical analysis hat clarifies the distance from a step that is necessary to prevent ailure when lifting the robot\u2019s front wheels during step climbing.\n.1. Moving distance of basic coordinate system for the robot inclines\nIn Fig. 18, O is the contact position between the robot\u2019s middle wheels and the road surface, which is also the basic coordinate system for the vehicles before lifting the robot\u2019s front wheels, 0\u03a3B. O moves to O\u2032 as the robot inclines, so the basic coordinate ystem for the vehicles is expressed as \u03a3B. pi (i = 1\u20135) are the joints (p1: axes of the robot\u2019s middle\nheels, p2: shoulder joints, p3: elbow joints, p4: robot hands location where the hands hold the wheelchair push handle), p5: xes of the wheelchair\u2019s rear wheels, and 0pi (i = 1\u20135) when the obot incline is zero.\nThe position vectors for the joints in the system \u03a3B are exressed as Bpi = [Xi Yi]\nT (i = 1\u20135). In particular, the position ectors for the joints when the robot incline is zero are expressed s B\n0pi = [0Xi 0Yi] T (i = 1\u20135).\nFrom Figs. 17 and 18, it can be seen that when \u03a3i is parallel o \u03a3 in the local coordinate system, Bp = [0 R ] T , 1p =\nB 1 B 2\n[lLB hLB] T , 2p3 = [l2 0]T , 3p4 = [l4C 0]T , and 4p5 = [lLA \u2212 hLA] T . Moreover, \u03c6i is the angle of \u03a3i formed by \u03a3i\u22121. In Stage 3, the wheelchair stops on the step and the robot continues driving forward as the system lifts that robot\u2019s front wheels (see Section 3). Since the robot\u2019s angle increases as it moves forward, the position of the basic coordinate system for the vehicles moves from 0\u03a3B to \u03a3B (Fig. 18), and the robot\u2019s center of mass shifts behind the contact point between the robot\u2019s middle wheels and the ground as the angle of the robot increases. During this process, the positions of p1, p2 and p3 are changed, but the positions of p4 and p5 are maintained because the wheelchair remains stopped while the robot\u2019s front wheels are lifted.\nThe homogeneous transformation matrices BT 4 are given below.\nBT 4 : [cos\u03c61234 \u2212 sin\u03c61234 X4 sin\u03c61234 cos\u03c61234 Y4\n0 0 1\n] (1)\nHere, \u03c61234 = \u03a34 i=1\u03c6i is the incline of the wheelchair against the ground. Therefore, \u03c61234 = 0 in Stage 3. The position vector of push handle of the wheelchair before lifting the robot\u2019s front wheels, Bp = [ X Y ] T in \u03a3 , is as\n0 4 0 4 0 4 0 B", "w\nb o\nT h ( r w\n\u2206\nS\n\u2206\nB 0\n\u03c1\nt\nB\nH\ns\ngiven below (Fig. 18).\nB 0p4 :\n[ 0X4\n0Y4\n] = [ cos 0\u03c61 \u2212 sin 0\u03c61 sin 0\u03c61 cos 0\u03c61 ][ lLB hLB ] + l2 [ cos 0\u03c612 sin 0\u03c612 ] + l4C [ cos 0\u03c6123 sin 0\u03c6123 ] + [ 0 RB ] (2)\nhere 0\u03c612 = 0\u03c61 +0 \u03c62, and 0\u03c6123 = 0\u03c61 +0 \u03c62 +0 \u03c63. When the robot center of mass shifts behind the contact point\netween its middle wheels and the ground, the angles of the links f the manipulators, 0\u03c61, 0\u03c612, and 0\u03c6123, shift to \u03c61, \u03c612, and \u03c6123, respectively. Thus, the position vector of hand rim of the wheelchair, Bp4 = [X4 Y4] T in \u03a3B, is as given below.\nBp4 : [ X4 Y4 ] = [ cos\u03c61 \u2212 sin\u03c61 sin\u03c61 cos\u03c61 ][ lLB hLB ] + l2 [ cos\u03c612 sin\u03c612 ] + l4C [ cos\u03c6123 sin\u03c6123 ] + [ 0 RB ] (3)\nhe basic coordinate system 0\u03a3B moves to \u03a3B in Stage 3. \u2206x is the orizontal movement of coordination as the robot\u2019s angle inclines the distance between 0\u03a3B and \u03a3B, Fig. 19). The wheelchair emains stopped until the system finishes lifting the robot\u2019s front heels. p4 does not change. Thus,\nx = \u2212(X4 \u2212 0X4) (4)\nubstituting (2) and (3) for (4),\nx = l4C (cos 0\u03c6123 \u2212 cos\u03c6123) + l2(cos 0\u03c612 \u2212 cos\u03c612) +lLB(1 \u2212 cos\u03c61) + hLB sin\u03c61 (5)\nThe position vector of \u03a3B in 0\u03a3B is described below (Fig. 19).\np\u2206x = [\u2206x 0]T (6)\n4.2. Robot front wheel height needed to climb a step\nThe angle of the robot necessary to place the front wheels on a step is clarified in Stage 3. In Fig. 20, q1 is the position of the robot\u2019s middle wheel axes (Here, q1 = p1, Fig. 18), q2 is the position of the robot\u2019s front wheel axes, and q3 is the tread of the robot\u2019s front wheels. \u03a3I and \u03a3II are the local coordinate system. (Here, \u03a3I = \u03a31, Fig. 18.) \u03c11 and \u03c12 are the angles of \u03a3I and \u03a3II formed by \u03a3B and \u03a3I , respectively.\nFrom Figs. 18 and 20,\n\u03c11 = \u03c61 (7)\nThe position vectors for the joints are expressed as Bqj = [xj yj]T (j = 1 \u2212 3). When \u03a3 and \u03a3 parallel \u03a3 , Bq = [0 R ] T ,\nI II B 1 B\n1q2 = [WBf \u2212 RB + rBf ]T and 2q3 = [0 rBf ]T (Figs. 17 and 20). 12 = \u03c1123 = \u03c11 + \u03c12 because \u03c13 = 0. Thus, the homogeneous ransformation matrices Bt2 and Bt3 are as given below.\nt2 : [cos \u03c112 \u2212 sin \u03c112 x2 sin \u03c112 cos \u03c112 y2\n0 0 1\n] (8)\nBt3 : [cos \u03c112 \u2212 sin \u03c112 x3 sin \u03c112 cos \u03c112 y3\n0 0 1\n] (9)\nHere,\nBq2 : [ x2 y2 ] = [ cos \u03c11 \u2212 sin \u03c11 sin \u03c11 cos \u03c11 ][ WBf \u2212(RB \u2212 rBf ) ] + [ 0 RB ] (10)\nBq3 : [ x3 y3 ] = [ cos \u03c11 \u2212 sin \u03c11 sin \u03c11 cos \u03c11 ][ WBf \u2212(RB \u2212 rBf ) ] +rBf [ cos \u03c112 sin \u03c112 ] + [ 0 RB ] (11)\nWhen the tread of the robot\u2019s front wheels, q3, is at the bottom of the wheel, \u03c112 = \u221290 deg (Fig. 20). In this case, from (7) and (11),\ny3 = WBf sin\u03c61 \u2212 (RB \u2212 rB) cos\u03c61 + RB \u2212 rBf (12)\nere,\nin\u03c61 = \u221a 1 \u2212 cos2 \u03c61 (13)\nFrom (7), (12), and (13),\ncos\u03c61 = e1(e1 \u2212 y3) + WBf\n\u221a 2 \u00b7 e1 \u00b7 y3 + WBf 2 \u2212 y32\ne12 + WBf 2 (14)\nHere, e = R \u2212 r .\n1 B Bf" ] }, { "image_filename": "designv11_5_0002210_j.robot.2020.103670-Figure16-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002210_j.robot.2020.103670-Figure16-1.png", "caption": "Fig. 16. Contact between the front wheels and step.", "texts": [ " The system lifts the front wheels of the wheelchair (Stage ) using the velocity differences between the connected vehicles. owever, if the wheelchair is too close to the step in Stage 1, the ront wheels will collide with the step, and a step-climbing failure w b t w t t r s a i t T a w t f 4 s w ( a r p v a t ill occur [23]. Similarly, it is necessary to leave a certain distance etween the robot\u2019s front wheels and the step to facilitate lifting he robot\u2019s front wheels. If the distance is too short, the robot\u2019s heels will collide with the step (Fig. 16). Since the mechanism of the robot\u2019s front wheels permits them o be folded (Fig. 4), even if the front wheels collide with the step, hey will be pushed inward, thus causing the wheelbase of the obot to shorten until the wheels are finally able to climb the tep. However, while contacts between the robot\u2019s front wheels nd the step do not always cause a step-climbing failure, there s a limit to the extent of an acceptable wheelbase transformaion. (The minimum value of WBf is 190 mm. See Appendix A, able A" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001549_j.ymssp.2020.107328-Figure7-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001549_j.ymssp.2020.107328-Figure7-1.png", "caption": "Fig. 7. Experiment for the slider-crank mechanism.", "texts": [ " Once Slider 2 reaches the maximummovement, it begins to move upward along Rail 2. Meanwhile, Linkage L2 drives Slider 1 to move right along Rail 1. The angular velocity of the motor is 15 RPM. The load applied on the mechanism is 500 N. In the slider-crank mechanism, there are three revolute joints: joint A, joint B and joint C. The initial clearance of the three joints is 0.05 mm. Table 2 lists the properties of the mechanism. A wear experiment of the slider-crank mechanism is conducted, and the test rig is shown in Fig. 7. Since the hardness of steel is much higher than that of brass, wear of the journal is considered negligible. As a result, only wear of the bearings is measured in the experiment. Materials of the joints in the driving system are all steel, and wear of them is also ignored. All the joints are under dry contact. The mechanism operates for 750,000 cycles, and wear of the bearings is measured every 30,000 cycles. The method in [16] is adopted to measure wear of the bearings: The diameter of bearings at 6 different positions uniformly distributed along the circular direction are measured, shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0003139_science.89.2314.419-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0003139_science.89.2314.419-Figure1-1.png", "caption": "FIG. 1. Section of pump.", "texts": [ " SCIENTIFIC APPARATUS AND LABORATORY METHODS A GLASS AND RUBBER LABORATORY PUMP THE pump to be described here is a modification of one designed by Palmer.1 Like Palmer's, it is operated by compressed air, and the fluid being pumped is in contact with glass and rubber only. It has the added advantages of greater capacity-1,500 cc per minute against a pressure of over a meter has been attained-and of operating in any position. 1 J. Palmer, SCIENCE, 80: 229-230, 1934. The construction is shown in Fig. 1. The cylinder is made of a piece of Pyrex tubing with an internal diameter of two inches. Considerable strength is necessary, but for pumps built on a smaller scale Pyrex need not be used. The diaphragm and valves are of rubber dam about 1/40 inch thick. They must not be stretehed too tightly. Ecru dam should be used; the \"buckskin\" rubber dam popular with dentists fails after two or three weeks. The life of these parts is increased by attaching with rubber cement a small piece of rubber (e.g., inner tube) to the diaphragm and to each valve, as shown in the figure", " IAZIER UNIVERSITY OF CALIFORNIA AT Los ANGELES AN INEXPENSIVE WARM STAGE A WARM stage is a very useful addition to the microscope, particularly for such investigations as the demonstration of motile amoebae in stools. Where only an occasional examination is made, the cost ($15 to $20) is likely to be prohibitive. A very serviceable stage can be made at a trivial cost from an electric iron heating element. This comes copper clad and slotted as shown in the drawing. It draws 550 watts at 110 volts, becoming red hot. However, if it is connected to the secondary terminals of a / H 5 ECO'VOAY/ r 7R AA'sFORMER FIG. 1 bell-ringing transformer, it does not take much current and rises to a temperature of 400 C. or less. By connecting a 6-ohm radio rheostat in series, the current can be regulated so as to maintain a temperature of 370 C. The construction is quite simple. With a hack-saw or grindstone, remove most of the side of a piece of A\" pipe about 3V' long, leaving both ends. This furnishes the thermometer carrier. The slot enables the operator to read the temperature. Lay the element on a piece of asbestos and connect the terminals to the 110-volt mains" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000861_012046-Figure8-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000861_012046-Figure8-1.png", "caption": "Figure 8. Configurations of Single-edge-crack-at-hole specimen (radius, r = 5 mm, width of the leaves, w = 25 mm).", "texts": [ " From the calculation results obtained that the stress that occurs in the spring is 590.55 MPa. It can be seen here that the stress that occurs exceeds the yield stress of the material (\u03c3y) as shown in Table 1. Therefore, it can be stated that the spring will fail due to deformed. Furthermore, it will also be examined from the crack analysis by finding out how much the stress intensity factor (KI) occurs in the spring due to the load received. Using a Single-edge-crack-at-hole specimen as shown in Figure 8, the value of KI is can be calculated by means of equation (3) [16]. International Seminar on Metallurgy and Materials IOP Conf. Series: Materials Science and Engineering 541 (2019) 012046 IOP Publishing doi:10.1088/1757-899X/541/1/012046 Where a is the crack length. The crack length is obtained from the visual observation with the value is about 1.5 mm, \u03c3 is the stress that occurs in the spring as it found from equation (2). Fhs is the boundary-correction factor calculated referred to previous research methods [16]" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000396_036103-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000396_036103-Figure3-1.png", "caption": "Figure 3. (a)Bottom/core and (b) topmolds.", "texts": [ " This results from two different effects on the geometry. Since the inductor core is being stretched but the number of turns is remaining constant, the turns per unit length are dropping linearly. However, the cross sectional area is also decliningwith stretch, further reducing the total inductance. In our previous work [7], a two layer channel fabrication process was developed tomake a non-magnetic solenoid. Amodified version (figure 2)was used in this work to create the ferrofluid-based inductors. Polycarbonatemolds (figure 3)were printed using a commercial 3D printer (Stratasys FDMTitan). The top and bottommolds each have raised features to define the conductive traces for the solenoid, while pillars in the bottommold create the vertical vias. A third piece of polycarbonate, a simple bar for defining the core, is placed onto a pedestal on the bottommold. Liquid precursors of Ecoflex 00\u201330 soft silicone (Smooth-on) are then poured into the bottommold (figure 2(a)) and de-gassed under vacuum tominimize air bubbling. Additional Ecoflex is also dropped onto the features of the topmold and degassed, after which the topmold is placed onto the bottommold; simple ring-peg features are used to align the two layers" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002230_ecce44975.2020.9236388-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002230_ecce44975.2020.9236388-Figure1-1.png", "caption": "Fig. 1: Motor under test.", "texts": [ " Sample data obtained by measurements are presented for all considered machine components. Finally, expert-based calibration and optimization of parameters that are difficult to predict are presented in Section V. This is completed by a detailed comparison of modeling results versus additional measurements. Finally, a summary and an outlook about future activities is illustrated in Section VI. In this paper, the thermal characteristics of a 3-phase, brushless SPM machine with outer rotor topology is studied. The motor with disassembled rotor is shown in Fig.1. It has a rated power of 110W at 150mNm and 7100 rpm. Its most important data, as well as its electrical characteristics are summarized in Table I. One big challenge of creating a thermal model for this type of machine is the rotating bell in the air that leads to a complex overall heat flux. Additionally, an air gap between rotor bell and mounting flange allows air- and heat-exchange, which makes modeling even more challenging. The lumped parameter thermal network selected to model the outer rotor SPM machine under test is depicted in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002337_iros45743.2020.9341751-Figure2-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002337_iros45743.2020.9341751-Figure2-1.png", "caption": "Fig. 2: Schematic of a generic robot with m actuators showing the notation used in this paper.", "texts": [ " This makes them unfavorable for moving contacts such as when the contact is rolling or sliding. Velocity-based contact localization methods are less common, but recent work on the HyQ robot [1] suggests they hold merit. The paper follows modified notations from [13]. All vectors are denoted with lowercase boldface fonts (e.g. a or c), all matrices are uppercase boldface font (e.g. J or Ad), all sets are written in uppercase font (e.g. B or C), and all frames are written in roman typestyle (e.g. P or L.) The terms defined in this section are shown in Fig. 2. We consider a robot with m actuators and n rigid links. We assume that the robot motion is constrained to be planar (although we relax this assumption in Section VIB). We assume that the robot has an accurate knowledge of its geometry, and can measure its actuator positions and velocities, which are denoted as q = (q1, . . . , qm)T and q\u0307 = (q\u03071, . . . , \u02d9qm)T . We also assume that the robot has an accurate estimate of its body velocity twist vb S,P, where S is a frame fixed to the world, and P is a frame fixed to the robot\u2019s body" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000840_s0263574719000882-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000840_s0263574719000882-Figure1-1.png", "caption": "Fig. 1. Structure of a 2-DOF robot manipulator.", "texts": [ "25 In this paper, in order to achieve a precise trajectory tracking and robustness against the model uncertainties for the 2-degree of freedom (2-DOF) robot manipulator, the optimized grey wolf optimizer-hybrid fractional order proportional derivative sliding mode controller (GWOHFOPDSMC) and extended grey wolf optimizer-hybrid fractional order proportional derivative sliding mode controller (EGWO-HFOPDSMC) are applied. Furthermore, a novel FOSS is defined. The simulation results of the proposed controllers are compared with FOSMC and hybrid fractional order proportional derivative sliding mode controller (HFOPDSMC), which show the effectiveness of the proposed optimized controllers. A typical 2-DOF robot manipulator is presented in Fig. 1. The dynamic equations of a typical 2-DOF robot manipulator are as follows7 : B(q)q\u0308 + C(q, q\u0307)q\u0307 + G(q) = \u03c4, (1) https://www.cambridge.org/core/terms. https://doi.org/10.1017/S0263574719000882 Downloaded from https://www.cambridge.org/core. Bethel Seminary St. Paul, on 13 Jun 2019 at 18:44:09, subject to the Cambridge Core terms of use, available at where the parameters are defined as follows: q = [ \u03b81 \u03b82 ] , B(q) = [ (M1 + M2)L2 1 + M2L2 2+2M2L1L2cos\u03b82 M2L2 2 + M2L1L2cos\u03b82 M2L2 2 + M2L1L2cos\u03b82 M2L2 2 ] , C(q, q\u0307)q\u0307 = [\u2212M2L1L2sin\u03b82 ( \u03b8\u03071\u03b8\u03072 ) \u2212M2L1L2sin\u03b82 ( \u03b8\u03072 2 ) M2L1L2sin\u03b82 ( \u03b8\u03072 2 ) ] , G(q) = [\u2212(M1 + M2)gL1sin\u03b81 \u2212 M2gL2sin(\u03b81 + \u03b82) \u2212M2gL2sin(\u03b81 + \u03b82) ] , \u03c4 = [ \u03c41 \u03c42 ] , (2) where q, q\u0307, and q\u0308 \u2208 R2 are the position, velocity, and acceleration of the joints, respectively; B(q) \u2208 R2\u22172, C(q, q\u0307) \u2208 R2\u22172, and G(q) \u2208 R2 represent the generalized inertia matrix, the coriolis/centripetal matrix, and the gravitational vector, respectively; and \u03c4 \u2208 R2 presents the joints\u2019 torques" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000549_s11548-016-1475-3-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000549_s11548-016-1475-3-Figure5-1.png", "caption": "Fig. 5 Local coordinate system of the maxillary model", "texts": [ " Table 2 Occlusal characteristics Model Incisor relationship Permanent teeth Upper crowding Lower crowding Buccal relationship left Buccal relationship right Overbite Subject1 I Yes Mild Mild I I Average Subject2 I Yes None Mild I 1/2II Average Subject3 I Yes Mild Mild I I Average Subject4 II div 1 Yes Mild Aligned 1/2II 1/2II Increased Subject5 II div 1 Yes Mild Spaced I I Increased Subject6 I Yes Aligned Mild I I Average Subject7 III Yes Severe Mild 1/2II 1/2II Average Subject8 III Mixed Aligned Aligned I I Reduced Subject9 III Yes Severe Severe 1/2II 1/2II Average Subject10 II div 1 Yes Severe Severe 1/2II 1/2II Increased Subject11 II Yes Aligned Mild I I Average Subject12 II div 1 Yes Mild Moderate I I Increased Subject13 II div 1 Yes Aligned Aligned II II Increased Subject14 III Yes Mild Aligned I I Reduced Subject15 I Yes Aligned Mild I I Reduced I, II div 1, and III are the classification of malocclusion Subsequently, the participants were asked to determine the occlusions of the fifteen test sets of virtual dental models without haptic feedback (non-haptically aligned), again repeating five times each model set. A total of 35 nonhaptically aligned occlusion results for each model set were recorded. A: Validation 1: Translational and angular deviations The local coordinate system of the maxillary model is shown in Fig. 5. O is the origin which indicates the centroid of the dental models. The translational deviation of the centroid and the angular deviation around the x, y, and z directions were used to evaluate the difference of the haptically aligned dental models and the benchmarks. Firstly, the translational deviation (theEuclidean distance) of the centroid in the x, y and z directions between the haptically aligned dental models and that of the benchmarks were computed for each aligned result. It generates 35 sets of the translational deviations for each pair of the models" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002316_rcar49640.2020.9303297-Figure4-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002316_rcar49640.2020.9303297-Figure4-1.png", "caption": "Fig. 4. The parts for finger fabrication in (a), the fabrication scenario in (b), and the fabricated hybrid finger in (c).", "texts": [ " The thickness of the fabricated finger in Fig. 2e is 4.5mm. To solve the above-mentioned problems, we proposed a new structure of the hybrid finger, as shown in Fig. 3. We eliminated the accommodated cavity and designed a incline lock for fixing the soft membrane with the rigid shell. During fabrication, release agent was coated on the interface of the rigid shell. Therefore, an air chamber can be naturally formed by the membrane and the rigid shell. Three parts were 3D printed for fabricating a new hybrid finger, as shown in Fig. 4a. During fabrication, the shell was assembled on the top of the hose insert, and then liquid rubber was pour into the assembly. Finally, by fixing the cover on the top of the assembly, we can complete the fabrication process as shown in Fig. 4b. After curing, the cover and the hose insert were removed and the hybrid finger is finally fabricated as shown in Fig. 4c. This fabrication method eliminated demolding, bonding, and gluing processes, and can significantly reduce the time cost of fabrication. In addition, the finger thickness can be reduced due to the 189 Authorized licensed use limited to: Central Michigan University. Downloaded on May 14,2021 at 03:22:08 UTC from IEEE Xplore. Restrictions apply. elimination of the accommodated cavity. The actual thickness of the fabricated finger in Fig. 4c is 3.0mm. Moreover, the new fabrication method facilitates mass production because of the simplified process, and therefore can lower the cost of the gripper. To package multiple cucumbers simultaneously, we need to minimize the gaps between neighboring cucumbers while packaging to save package space. At the grasping position, the cucumbers are approximately aligned by using a wave conveyor and have enough gaps for the gripper to enter. After grasping, the gripper should be able to shrink the gaps before packaging" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001220_j.matpr.2020.02.550-Figure5-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001220_j.matpr.2020.02.550-Figure5-1.png", "caption": "Fig. 5. Boundary condition for bumper beam.", "texts": [ " The thickness of the bumper beam and brackets is less compared to other dimensions so the shell element is used and the uniform mesh is generated. The meshing is originated by the 2D meshing because the improvements in meshing quality are easy to do in 2D meshing. The element used for the 2D meshing is tria element with the element size of 4 mm. Then after performing quality checks, the 2D tria mesh is converted into tetra mesh of 4 mm size by using a dummy solid element. The meshed model f bumper beam. The boundary conditions as shown in Fig. 5 are used for the analysis of the bumper beam in HyperMesh. Boundary condition contains pressure on the frontal area and the bracket ends are constrained. The stresses induced after performing impact test at low speed i.e. 10, 20 and 30 km/h are shown in Fig. 6. The red color (grids Please cite this article as: S. S. Godara and S. Narayan Nagar, Analysis of frontal b rial, Materials Today: Proceedings, https://doi.org/10.1016/j.matpr.2020.02.550 The stress and displacement induced after performing impact test at low speed i" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002356_978-3-030-50460-1-Figure14.1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002356_978-3-030-50460-1-Figure14.1-1.png", "caption": "Fig. 14.1 Applied deformation: (a) time proportional with varying strain rate, (b) time periodic with different strain amplitude, (c) time periodic with different maximum strain rate, (d) relaxation tests with different preceding strain amplitudes", "texts": [ " Consequently, the constitutive law used for the continuum has to be identified and parameterized accordingly. To this end, we introduce a methodology to characterize the mechanical behavior of a material directly from MD simulations (Ries et al, 2019) based on a classification scheme by Haupt (1993) depicted in Tab. 14.1. Similar ideas of analyzing constitutive behavior via numerical pseudo-experiments have been applied for other classes of materials as well, e.g., Turco et al (2019). We have derived four load cases from this, visualized in Fig. 14.1: time proportional loading with different strain rates to evaluate the rate dependence, time periodic tests either with varying strain amplitude or varying strain rate to quantify the inelastic effects, and relaxation tests with preceding time periodic loading to identify whether a quasi-static hysteresis exists. Up to now, we applied this scheme to investigate atactic polystyrene as a model system under uniaxial deformation and identified purely elastic behavior for small strains, followed by a viscoplastic regime for larger deformations (Ries et al, 2019)" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000745_iet-map.2018.6170-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000745_iet-map.2018.6170-Figure1-1.png", "caption": "Fig. 1 Geometry of the proposed single-polarised band-absorptive FSR (H = 9.5 mm, b = 8 mm, l = 6.75 mm, l1 = 6 mm, w = 0.5 mm, t = 0.508 mm, Resistor = 250 \u03a9, \u025br = 2.2)", "texts": [ " Here, miniaturised band-absorptive FSRs for single polarisation and dual-polarisation are proposed, whose sizes are less than that of the FSR presented in [22] but their absorption bandwidths are wider than that of the FSR presented in [22]. These two FSRs are both realised by cascading a lossy layer and a lossless layer. I-type and Jerusalem-cross-type metal strips are used in lossy layer and lossless layer to achieve miniaturised structures. The geometry of the proposed single-polarised band-absorptive FSR is presented in Fig. 1. Periods of the unit cell along the x- and y-directions are both b and the whole height of the FSR is H. The detailed geometrical parameters of the FSR have been annotated in the figures and listed in the caption. The unit cell of the FSR consists of a lossy layer and a lossless layer. They are both formed by I-type metal strips printed on a substrate. A lumped resistor is welded on the mid of the I-type metal strips of the lossy layer. Equivalent circuit models (ECM) of the proposed singlepolarised band-absorptive FSR is shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0001418_acc45564.2020.9147448-Figure1-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0001418_acc45564.2020.9147448-Figure1-1.png", "caption": "Fig. 1. A CAD render of Midget", "texts": [], "surrounding_texts": [ "I. Introduction\nRaibert\u2019s hopping robots [20] and Boston Dynamic\u2019s BigDog [21] are amongst the most successful examples of legged robots, as they can hop or trot robustly even in the presence of significant unplanned disturbances. Other than these successful examples, a large number of humanoid robots have also been introduced. Honda\u2019s ASIMO [14] and Samsung\u2019s Mahru III [18] are capable of walking, running, dancing and going up and down stairs, and the Yobotics-IHMC [19] biped can recover from pushes.\nDespite these accomplishments, all of these systems are prone to falling over. Even humans, known for natural and dynamic gaits, whose performance easily outperform that of today\u2019s bipedal robot cannot recover from severe pushes or slippage on icy surfaces. Our goal is to enhance the robustness of these systems through a distributed array of thrusters.\nHere, in this paper, we report our efforts in designing feedback for the thruster-assisted walking of a bipedal robot, called Midget, currently being developed at Northeastern University. The biped is equipped with a total of six actuators, and two pairs of coaxial thrusters fixed to its torso as shown in figure I. Each leg is equipped with three actuated joints, the actuators located at the hip allow the legs to move sideways and actuation in the lower portion of the legs is realized through a parallelogram mechanism.\nPlatforms like Midget that combine aerial and legged modality in a single platform can provide rich and chal-\n1Pravin Dangol and Alireza Ramezai are with the Department of Electrical and Computer Engineering, Northeastern University, Boston, MA 02115 dangol.p@husky.neu.edu, a.ramezani@northeastern.edu\n2Nader Jalili is with the Department of Mechanical Engineering, University of Alabama, Tuscaloosa, AL 35487 njalili@eng.ua.edu\nlenging dynamics and control problems. The thrusters add to the array of control inputs in the system (i.e., adds to redundancy and leads to overactuation) which can be beneficial from a practical standpoint and challenging from a feedback design standpoint. Overactuation demands an efficient allocation of control inputs and, on the other hand, can safeguard robustness by providing more resources.\nThe challenge of simultaneously providing asymptotic stability and constraint satisfaction in legged system has been extensively addressed [25]. The method of hybrid zero dynamics (HZD) has provided a rigorous model-based approach to assign attributes such as efficiency of locomotion in an off-line fashion. Other attempts entail optimization-based approaches to secure safety and performance of legged locomotion [7], [5], [6].\nThe objective being pursued here is key to overcome a number of limitations in our platform and involves fast performance constraint and impact invariance satisfaction. To put it differently, smaller robots have faster dynamics and possess limited actuation power and these prohibitive limitations motivate us to look for motion control solutions that can guarantee asymptotic stability and satisfy performance with minimum computation costs.\nInstead of investing on costly optimization-based scheme in single support (SS) phase, we will assume for well-tuned supervisory controllers as found in [23],\n978-1-5386-8266-1/$31.00 \u00a92020 AACC 3217\nAuthorized licensed use limited to: Carleton University. Downloaded on August 01,2020 at 10:07:07 UTC from IEEE Xplore. Restrictions apply.", "[17], [2] and will instead focus on fine-tuning the desired joint trajectories by implementing an intermediary filter based on the emerging idea of reference governors [10], [1],[11], [9] in order to satisfy the performance being sought. Since these modifications and impulsive impact events between gaits lead to deviations from the desired periodic orbits, we will enforce invariance to impact in a robust fashion by applying predictive schemes withing a very short time envelope during the gait cycle, i.e. double support (DS) phase. To achieve hybrid invariance, we will leverage the unique features in our robot, i.e., the thruster.\nThis work is organized as follows. In section II the multi-phase dynamics for a planar walking gait is developed. The SS phase is modeled, and a two point impact map followed by a non-instantaneous DS phase are introduced. In SS phase gaits are first designed based on HZD method, constraints are imposed on an equivalent variable length inverted pendulum (VLIP) model through an explicit reference governor (ERG), the equivalent control action are then mapped back to the full dynamics. During DS phase a nonlinear model predictive control (NMPC) scheme is introduced to ensure performance satisfaction and steer states back to zero dynamics manifold ensuring hybrid invariance. Results are shown in section III, and we conclude the paper in section IV.\nII. Thruster assisted model with extended\ndouble-support phase\nA full cycle of our model involves consecutive switching between 1) SS phase where only one feet is on the ground, 2) an instantaneous impact map that occur at the end of the SS phase and 3) DS phase where both feet stay in contact with the ground. This model is slightly different form previous works on under-actuated planar bipedal locomotion [12], [26], [3] [4], [13] which assume the double support phase is instantaneous. The extended double support phase will provide a time envelope before the onset of the swing phase for post-impact corrections.\nDuring SS phase the biped has 5 degrees of freedom (DOF), with 4 degrees of actuation (DOA), shown in Fig. 2a. Following modeling assumptions widely practiced, it is assumed that the stance leg is fixed to the ground with no slippage, and the point of contact between the leg and ground acts as an ideal pivot. The kinetic K(q, q\u0307) and potential V(q) energies are derived to formulate the Lagrangian, L(q, q\u0307) = K(q, q\u0307)\u2212 V(q), and form the equation of motion [25]:\nDs(qs)q\u0308s + Hs(qs, q\u0307s) = Bs(qs)u (1)\nwhere Ds is the inertial matrix independent of the under-actuated coordinate, Hs matrix contains the Coriolis and gravity terms, and Bs maps the input torques\nto the generalized coordinates. The configuration variables are as follows: qT is the absolute torso angle; q1R, q1L are the angles of the \u201dfemur\u201d relative to torso; and q2R, q2L are the angles of virtual \u201dtibia\u201d relative to \u201dfemur\u201d as shown in Fig. I and 2a. The configuration variable vector is denoted by qs = [qT , q1R, q1L, q2R, q2L]\nT \u2208 Qs.\nAs pointed out earlier we assume that an impulsive effect similar to what described in [15] occurs between two continuous modes of SS and DS. We follow the steps from [15] to model the impact event and solve for the reaction forces Fext. The unconstrained version of (1) are considered by augmenting qs and including the hip position, qe = [qs, pH ]\nT . The Lagrangian is reformulated and the impulsive force is added on\nDe(qe)q\u0308e + He(qe, q\u0307e) = Be(qe)u + \u03b4Fext (2)\nwhere the external force Fext acting on each feet end p = [p1, p2] T can be expressed as following\nFext = JT\u03bb = [ \u2202p1/\u2202qe \u2202p2/\u2202qe ]T \u03bb\nwhere \u03bb = [\u03bb1, \u03bb2] T , shown in Fig.2b, is the Lagrange multiplier and assumes that both legs stay on the ground upon impact and Jacobian matrix J is given by J = \u2202p\n\u2202qe . After assuming that the impact is inelastic,\nangular momentum is conserved and two legs stay in contact with the walking surface, the impact map is resolved[\nDe(q\u2212e ) \u2212J(q\u2212e )T\nJ(q\u2212e ) 04\u00d74\n] [ q\u0307+e \u03bb ] = [ De(q\u2212e )q\u0307\u2212e\n04\u00d71\n] (3)\nwhere the superscript + denotes post-impact and \u2212 denotes pre-impact. It is straightforward to show that like [25], the Jacobian matrix J has full row rank and the inertial matrix is always positive definite, the matrix on\n3218\nAuthorized licensed use limited to: Carleton University. Downloaded on August 01,2020 at 10:07:07 UTC from IEEE Xplore. Restrictions apply." ] }, { "image_filename": "designv11_5_0002540_j.triboint.2021.106983-Figure3-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002540_j.triboint.2021.106983-Figure3-1.png", "caption": "Fig. 3. Assembly drawing journal bearing test rig.", "texts": [ " The bulk oil temperature in oil tank was maintained at 30 \u25e6C for all tests. The data were observed and recorded for friction and vertical displacement with the help of winducom lab view software attached with test rig. Each test was repeated till similar results were observed. The lubricants for testing are selected N1 to N20 having PIB as VI improver from Table 1. Similarly, the experimental conditions were designed such that in conformal journal bearing we get hydrodynamic lubrication regimes. Fig. 3 shows apparatus used for measuring coefficient of friction and film thickness of journal bearing. Bearing assembly have highly polished journal size of \u03d5 50 \u00d7 1000 mm. Journal is made of case hardened steel having 60 HRC and supported on either ends on bearings. The bearings are housed inside spindle housing and in turn in housing. This housing is firmly fixed on base plate and has relief at middle to accommodate test bearing and sump for lubricant. Journal is driven by AC motor through 1:1 pulley and teeth belt with maximum speed of 3000 rpm" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0000557_aim.2016.7576989-Figure17-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0000557_aim.2016.7576989-Figure17-1.png", "caption": "Figure 17. Experimental setup for conveyance measurements", "texts": [ " 16 shows that the highest estimated conveyance amount per unit of time is obtained for the 0.5 s motion interval. In other words, the maximum conveyance amount at the time of the applied pressure of 40 kPa is assumed to correspond to the motion interval of 0.5 s. In this section, we examine whether the carrier is conveyed from the supply side by the peristaltic movement. We checked whether the supplied carrier is conveyance to the opposite side, and investigate the conveyance amount per time. Fig. 17 shows the experimental setup. We place the elbow pipe and the straight pipe made of acrylic on one end of the conveyor. At first, the conveyor is placed in an empty state in order to supply the carrier from the top of the acrylic pipe, and fill up the acrylic pipe. Furthermore, the carrier has sufficient amount that is always filled by the acrylic pipe. The applied pressure was 40 kPa because this is the pressure required for Conveyor unit ON / OFF valve Air compressor Regulator Fluctuation volume full extrusion in the conduit", " Increasing the amount of the carrier in the conduit leads to improvement in the conveyance amount. We will investigate the construction of a system for supplying a sufficient amount of powder in future work. As described in the previous section, it was found that an applied pressure of 40 kPa and motion interval of 0.5 s gave the highest conveyance amount. We therefore used these parameters to examine conveyance of the developer used in the actual printer by peristaltic conveyor in order to calculate the conveyance amount per unit time. We used the experimental equipment for Fig. 17 described in the previous section and the cyan toner (manufactured by RICOH) (density 2.0 g/cm3) as the developer. Fig. 22 presents the images obtained during the developer conveyance, showing that the developer was pushed in the arbitrary direction by the peristaltic movement. The obtained experimental results are shown in Figs. 23 and 24. The conveyance amount of the developer is lower than that for the carrier conveyance and the conveyance rate for the developer is 2.0 g/s, lower than the 5.6 g/s rate for the carrier conveyance" ], "surrounding_texts": [] }, { "image_filename": "designv11_5_0002166_s10846-020-01259-0-Figure9-1.png", "original_path": "designv11-5/openalex_figure/designv11_5_0002166_s10846-020-01259-0-Figure9-1.png", "caption": "Fig. 9 The three-axis platform with the designed attitude measurement unit", "texts": [ " A\u0302 (k + 1) = diff ( Af (k + 1) , x\u0302 (k + 1) ) (36) After calibrating the accelerometer, magnetometer and gyroscope using our designed calibration methods, and solving the attitude by fusing the data using the designed attitude estimator using C# to fuse the sensors\u2019 data, we conducted relevant experiments with data collected from the IMU\u2019s inertial sensors of the attitude measurement unit to verify the validity of the above studies. The designed attitude measurement unit shown in Fig. 1 was deployed on a three-axis platform, which consisted of three attitude measurement units, three motion units driven by three DC motors (Portescap 16G88214E servo motor), a microcomputer and a camera mounted on the end of the three-axis platform. And the three-axis platform was mounted on a quadrotor as shown in Fig. 9, which was used to test the three-axis simultaneous motion control based on improved orientation vector SLERP (spherical linear interpolation) method, with the details shown in our previous work [30]. One of the experimental result is as shown in Fig. 10, where, the horizontal ordinate indicates the serial number of the data point, and the vertical ordinate indicates the Euler angles. It can be seen that the fluctuation range of the roll angle and the pitch angle are both within \u00b10.05\u25e6, while the yaw angle fluctuation is up to \u00b10" ], "surrounding_texts": [] } ]