diff --git "a/designv11-83.json" "b/designv11-83.json" new file mode 100644--- /dev/null +++ "b/designv11-83.json" @@ -0,0 +1,9572 @@ +[ + { + "image_filename": "designv11_83_0003526_srin.202100362-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003526_srin.202100362-Figure1-1.png", + "caption": "Figure 1. Summary of the procedures that were used to fabricate the additively manufactured AISI 316L stainless steel tubes: (a) building the tubular preform by WAAM; (b) machining the tubular", + "texts": [ + " The chemical composition of the wire supplied by ESAB is given in Table 1 while the process parameters used in the deposition are provided in Table 2. High-purity (99.99%) argon gas was used as shielding to protect the molten feedstock from oxidization. A cc ep te d A rt ic le This article is protected by copyright. All rights reserved Current (A) Voltage (V) Wire feed speed (m/min) Stick-out length (mm) Gas flow rate (l/min) Layer height (mm) 100 16.5 6 10 10 1.8 The deposition strategy that was utilized for building the tubular preforms (Figure 1) consisted of single-bead circular paths that were repeated for each successively added metal layer (Figure 1a). After deposition, the tubular preforms had 42 mm outer diameter and 5 mm wall thickness and were later machined by turning (Figure 1b) to obtain tubes with a surface quality, a uniform wall thickness and dimensions similar to those of the commercial wrought tubes. The deposited tubes with a length \ud835\udc590 = 70 mm, an outer radius \ud835\udc450 = 20 mm and a wall thickness \ud835\udc610 = 1.5 mm, are shown in Figure 1c. te d A rt ic le This article is protected by copyright. All rights reserved preform to obtain a tube with the required dimensions and surface quality; (c) photograph of a ready-to-use tube. 2.2. Mechanical characterization of the materials The flow curves of the deposited and commercial wrought AISI 316L stainless steel tubes were retrieved from a previous work of the authors.[21] The results are shown in Figure 2 and involved conventional and ring hoop tension tests carried out at room temperature in an Instron 5900R universal testing machine" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002234_s00289-021-03570-8-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002234_s00289-021-03570-8-Figure2-1.png", + "caption": "Fig. 2 Shearing-field and stretching-field generated by velocity-gradient", + "texts": [ + " Moreover, though some researchers have found out the way to make short fiber get circle orientation using special structure die during process of extrusion [24, 25], it is just for production of tubes. How to make short fibers orientated in the radial direction? The related researches on this issue are still in progress. While in this paper, a way to make short fiber get radial orientation in rubber matrix of SFRC has been studied and worked out. During the extrusion process, there are two important factors to make short fiber get radial orientation. One is velocity-gradient which makes the rubber material suffer shearing-field force and stretching-field force in die channel, as shown in Fig.\u00a02. Another is extrusion swelling which is a special character of polymer such as rubber. Therefore, considering both velocity-gradient and extrusion swelling, a type of damexpanding die has been designed, as shown in Fig.\u00a03. According to this structure, the extrusion pressure would be increased by the dam during extrusion process (C area in Fig.\u00a03a). So, the material would get high pressure before the dam (i.e., A area) and get high internal energy for expansion. Then, due to the channel space after dam (i.e., B area) becoming big suddenly, the material would expand after passing the dam even bigger than extrusion swelling. As a result, short fibers in rubber matrix would turn with the expansion process and form an angle with X axis, as shown in Fig.\u00a02. In addition, with the extrusion process going on, the angle would get larger by action of shearing-field force and stretching-field 1 3 force. Eventually, if the angle reached 90\u00b0 ideally, it could be said that short fibers obtained radial orientation. In order to research its radial orientation behavior in rubber matrix during extrusion process, the physical model of short fiber radial orientation has been built as shown in Fig.\u00a04 in which the meanings of X, Y, Z are the same with that in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003265_j.rinma.2021.100211-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003265_j.rinma.2021.100211-Figure2-1.png", + "caption": "Fig. 2. (a) A representation of the interaction of the laser with the sample, showing the difference between the (b) global and (c) local fluence.", + "texts": [ + " A CAD model of the build plate with the printed specimens is presented in Fig. 1(a). The specimens were printed with a 4 mm offset, which allowed them to be removed from the build plate with wire electrical discharge machining. The final dimensions of the specimens were 10 mm \u00d7 10 mm x 10 mm. Fig. 1 (b) shows an individual specimen after removal from the substrate. The markings on the two side faces were added as reference points. The scan strategy selected for the specimens was a bi-directional serpentine pattern (Fig. 2(a)) that globally formed islands. Different combinations of laser powers, hatch spacings, and scan speeds were selected in order to vary the fluence. These combinations are presented in Table 2. Each specimen is identified using the laser power followed by the hatch spacing. Two fluences, the global laser fluence (GLF) and the LLF, were calculated for each specimen, as presented in Fig. 2(b) and (c). The GLF and LLF describe the energy exposure of the entire print surface for one print layer and the energy exposure when the laser travels one spot size distance on a single print layer, respectively. GLF is identical to the VED described in the introduction of this manuscript, but will be referred to as GLF to highlight how both GLF and LLF are measurements of applied energy per unit volume. Mathematically, GLF is defined by the equation presented by Thijs et al. [25], among others, as GLF = P v\u22c5t\u22c5y , (1) where GLF is a function of the laser power (P), scan speed (v), layer thickness (t), and hatch spacing (y)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002801_s00006-021-01119-6-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002801_s00006-021-01119-6-Figure7-1.png", + "caption": "Figure 7. A visual representation of \u03c8{i}, the angle between the joint axis (si) and the link vector (Hi), for i = 2, using an arbitrary 2R planar manipulator as an example", + "texts": [ + " A Type 1b1 transformation would require a secondary actuator that possesses a specialised locking mechanism to fix and vary the orientation of a joint, with respect to its base reference frame. For a Type 1c transformation, point masses or cylindrical masses can be attached to the centre of a link. Type 1a Changing the nodal weight of Hi to increase or decrease the length of the link or change the shape of the link Type 1b1 Changing the nodal weights of si and Hi simultaneously through reorientation, such that the angle between the joint axis and the link \u03c8{i}, as shown in Fig. 7, is preserved Type 1c Changing the mass/mass and moment of inertia distribution of links For every type transformation, we denote P K {i}(Ti) to be the \u2018transfor- mation operation\u2019, where K denotes the transformation type and i denotes the kinematic pair undergoing the Type K transformation. Ti is the \u2018transformation variable\u2019, which is the variable that the Type K transformation embeds onto vector position Xi i+1 (for K = 1a or K = 1b), or the link mass mi. The Type K transformations operations employed in this work are: 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002180_012018-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002180_012018-Figure2-1.png", + "caption": "Figure 2. Construction design of the device for measuring linear displacement of the belt when carrying out perforation: 1\u2013pressure roller, 2 \u2013 roller mounting, 3 \u2013 carriage, 4 \u2013 guide rail, 5 \u2013 weight, 6 \u2013 measuring roller, 7 \u2013 rotation angle sensor, 8 \u2013 sensor mounting, 9 \u2013 arm, 10 \u2013 plate, 11 \u2013 axle", + "texts": [ + " Compensating for positioning error entails calculating the difference between the set displacement and the displacement read by the encoder achieving friction coupling with the belt. The difference in value is sent to the BLDC drive controller operating in closed loop in incremental mode. The motion is carried out until the controller receives information about the drive finishing the motion, at which point the difference is determined once again. If the difference does not exceed the threshold value (0.5 mm), the belt side grippers close and the row of holes is being punched. The construction of the device for measuring linear displacement of the belt is shown on Fig. 2. The device for measuring belt displacement consists of two rollers: upper roller with bearing and lower rotation sensor roller. The cylinder surfaces of both rollers are in contact with, respectively, the top and bottom surface of the belt. The upper roller with bearing is attached to the device body which is connected to the carriage which moves along the vertical guide to facilitate adjustment of its height, enabling to introduce belts of any thickness between both rollers. Moreover, the upper body to which the upper roller with bearing is connected is fastened to an axle, on which weights can be placed to facilitate variable adjustment of the pressure of the upper roller with bearing onto the top surface of the belt, consequently ensuring the belt is indirectly pressed to the lower rotation sensor roller" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000851_imece2007-41027-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000851_imece2007-41027-Figure3-1.png", + "caption": "Fig. 3: Photo of the whole experimental test rig", + "texts": [ + " Testing conditions of gearwheels and its assembly should be similar as actual operational conditions. By reason of test time shortening it is necessary to select larger torque in closedloop circuit than in industrial operation. In our case, we are limited by torque sensor in circuit up to 5000 Nm. In the course of 1500 RPM the circuit is dimensioned for maximal virtual power 785 kW. Testing is mostly running on one load level because of better possibility of result comparing. The whole test-rig with PLC, control panel, converter and hydraulic devices is shown in figure 3. Developed back-to-back circuit is visualized in figure 4, its schema is in figure 5 and descriptions are in following paragraphs. Loading equipment must ensure easy creation of torque in the circuit. Creation of torque is realized by axial movement of the gearwheel (Fig. 5, Pos. 3) with a helical gear in mesh with the pinion in the additional gearbox (Fig. 5, Pos. 14). This system is similar to NASA Glenn Research Center Spiral Bevel Gear Facility [4]. The tensioning screw gives rise to axial force that causes reaction (tangential force) in the gearing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001548_isic.2008.4635965-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001548_isic.2008.4635965-Figure2-1.png", + "caption": "Fig. 2 shows the phase portrait for this system. Proposition 1: Equilibrium [0, 0] attracts all solutions (except the ones starting at other equilibria).", + "texts": [ + " The two lines ca=, i and a, + /il 0 are trajectories, since from (10) and (11), we get dI = = 1 or dOl 1. Thesedal~ & dcal lines connect the unstable equilibrium [-,-v] to the stable equilibria. The other equilibria are saddle points. Based on this topology and a hint of index theory (no lone saddle can be inside a periodic solution) the proposition follows. Since cos cq, + cos /i = 2 at [0,0], equation (12) implies that rj -- 2k1 r. This means that the distance between the two unicycles will converge to zero. U Fig. 3 shows one corresponding \"pursuit graph\" (for point E in Fig. 2). B. 2k1 = k2 case In this case the matrix in (14) is singular, the equilibria are given by 1 = { (c1,, 1) sinc1 = sin l 1}. This set is simply the union of the three lines Fig. 3. Pursuit graph when k1 = 0.1, k2 = 0.4, solid line denotes trajectory of vehicle 1 while dashed line denotes trajectory of vehicle 2, the initial condition for vehicle 1 is x1 (0) = 0; yI (0) = 4; \u00b0 l (0) =-2, the initial condition for vehicle 2 is x2 (0) = 5; Y2 (\u00b0) = 4; 02 (0) = -1, in relative coordinate this corresponds to initial condition a I (0) = 2; 1 (0) = 1 3 2O -2 -3- -3 -2 -1 0 (1 Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000795_jjap.46.6871-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000795_jjap.46.6871-Figure1-1.png", + "caption": "Fig. 1. Schematic diagram of the glucose sensor.", + "texts": [ + " By taking advantage of the semiconductor processing technology, an array of sensors with the same structure and the same size can be fabricated on the same wafer. Since all sensors in the array are fabricated at the same time, each sensor can have the same properties. Therefore, the sensor array with the proposed structure can be used to sense solutions with different glucose concentrations at the same time. The sensor array can also solve the issue of inaccuracy being caused by the inherent pH value of 2\u20133 of the oxide surface as well as the residue glucose solution for repetitive usage.10) Figure 1 shows a schematic diagram of the glucose sensor structure used in this work. It shows that the sensor is based on a typical MOSC with a window opened at the center of the gate area. The metal of the center area was etched away and the gate insulator was left untouched. In this work, the sensors were fabricated on 4 in. p-type (100) silicon wafers. After standard RCA cleaning procedures, a thermal dry oxide with a thickness of 20 nm was grown at 950 C. Then, a 30 nm Si3N4 was deposited onto the SiO2 layer by plasmaenhanced chemical vapor deposition (PECVD)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001520_6.2007-5736-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001520_6.2007-5736-Figure2-1.png", + "caption": "Figure 2: Tooth positions considered for the axial displacement study", + "texts": [], + "surrounding_texts": [ + "American Institute of Aeronautics and Astronautics\n2\nthese seals and the systems they are designed to protect can deteriorate over time, and for this reason, their in-service condition is of great importance and subject of many studies.\nDuring an engine operation, labyrinth seals rarely run according to the nominal conditions for which all experimental and numerical investigations are conducted. As an engine undergoes a maneuver or is transitioned from one power level to another (accel or decel), its stators and rotors deflect radially and axially at different rates due to their respective mechanical and thermal loads and structural/thermal characteristics. This differential movement results in a change in the radial clearance and axial position of the tooth relative to the stator, and often causes the rotor to rub the stator. To prevent potential failure of the seal teeth, abradable honeycomb structures are employed on the seal rack. These structures allow for rub-grooves to form on the honeycomb lands that can subsequently enlarge as the engine undergoes various operational maneuvers.\nStudies on seal leakage of labyrinth seal with and without rubbed honeycombs have been carried out experimentally and numerically1-7. A few studies9-11 have also investigated windage heating and total temperature rise associated with these seals. In all these studies the tooth has been either placed outside the groove, or was considered in its new condition with sharp tooth tips. There are only a few literatures available on the study of tooth tip wear and its axial displacement within the groove. Rhode and Allen13 conducted first measurements and visualization movies for rubbed stepped labyrinth seal with rounded tooth tips. For the cases considered, the leakage resistance decreased due to the presence of worn teeth by 85%, 55% and 70% for the small, medium and large prerub clearances. Only Rhode and Adams8 studied the effect of relative axial displacement on seal leakage for a straight seal with rub-grooves. But they only considered smooth stators and placed the seal tooth outside the rubgroove.\nIn an earlier study12, the authors developed a numerical methodology that is quite general and flexible, allowing for simulations of labyrinth seals in new and in-service conditions. The rubbed and un-rubbed honeycomb seals with rotor seals running inside and outside of rub grooves were simulated, and the flow structure, swirl, and windage heat rise for these configurations were examined in detail. Given the importance of seal deterioration and the impact of its in-service condition on its performance, this paper is entirely devoted to the investigation of the impact of tooth tip wear and its axial migration (placement) within the rub-grooves on the seal leakage and windage heating.\nII. Numerical Modeling Computational flow modeling is carried to quantify the effects of tooth tip wear and its axial displacement within the rub-grooves, on seal leakage and windage heating. The analysis is performed for a straight-through labyrinth seal with negative nominal seal clearance (seal teeth resides inside the rub groove). Our earlier paper12 provides the details of computational modeling and validation of the numerical methodology. In that paper the flow structures, leakage, and windage heating characteristics of seals with and without rub grooves were investigated. Tooth tips were placed inside, outside, and in-line with the rub grooves.", + "American Institute of Aeronautics and Astronautics\n3\nThis paper continues the investigation by studying the influence of tooth tip configuration and wear, as well as axial position of tooth within the rub groove. To obtain a broad understanding, a total of 14 configurations are simulated a summary of which is provided in Table 2. Nine simulations are devoted to study of worn seals covering three tooth tip configurations (sharp-edge, 4 and 12mil corner radii), rubbed and un-rubbed honeycomb lands with tooth tips placed at 0.010 mil radial clearance, and rubbed honeycombs with tooth tips placed in-line with the honeycomb edge (.020\u201d radial clearance). All these simulations are carried out with 1/32\u201dhoneycomb cells (HCs). Figures 1a) through 1f) show schematic of the tooth tips and seal configurations.\nTo find out the impact of tooth tip displacement relative to the rub groove, the sharp tooth tips are placed at five axial positions inside the rub groove as depicted in figures 2a through 2e. Axial position details of the teeth within the rub-grooves are shown in table1. These simulations use 1/16\u201d HCs and 10mil seal clearance. The four rectangular rub grooves (one for each tooth) are each 0.020\u201d deep and approximately 0.125\u201d wide. The rub grooves are the same as the ones used in the previous study12.\nFigure 3 shows the computational domain, the grid, and the boundary conditions. Same boundary conditions are used for all parametric studies. Hexagonal elements are used to generate a structured mesh using ICEM CFD without any GGI (Generalized Grid Interface) between the honeycomb and the seal. Total number of grids is approximately 1.4 million for all configurations. The commercial CFD solver, CFX 10.0 is used for these simulations, and second order high-resolution descritization scheme is employed. More on details of computations and grids can be found reference12. All the cases studies use a pressure ratio of 2.1 across the labyrinth seal with the seal tip Re of 1.21E6." + ] + }, + { + "image_filename": "designv11_83_0000006_iecon.2004.1433381-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000006_iecon.2004.1433381-Figure1-1.png", + "caption": "Fig. 1 Biped robot coordinate frames.", + "texts": [], + "surrounding_texts": [ + "The 30th Annual Conference of the IEEE Industrial Elecironics Society, November 2 - 6,2004, Busan, Korea\nBiped Robot Walk Control via Gravity Compensation Techniques\nOzan AYHAN Faculty of Engineering and Natural Sciences\nSahanci University ozana@su.sabanciuniv.edu\nA b s t r u c c Past three decades witnessed a growing interest in biped walking robots because o f their advantageous use in the human environment. However, their control is challenging because of their many DOFs and nonlinearities in their dynamics. Various trajectory generation and walking control approaches ranging from open Imp walking to systems with many sensors and feedback loops have been reported in the literature. The tuning of the parameters of reference gait is a common complication encountered. Another important problem of the walking control is the ground force interaction of the swinging leg at landing,\nThis paper uses position references for the upper body. Optimization techniques are employed to obtain suitable leg joint torques for the supporting leg to track body reference trajectories. Locomotion is achieved by a swinging leg control scheme which is independent of the body reference position tracking. This method is based on gravity compensation and virtual potential fields appIied to the swinging leg. Soft landing of the leg can easily be handled with this scheme, 3D dynamics and ground interaction simulation techniques are employed for a 12-DOF biped robot to test the proposed method. The simulations indicate the applicability of the method in real implementations.\n1. INTRODUCTION\nThe bipedal structure is one of the most versatik ones for the employment of walking robots in the environment of humans. It has supreme characteristics in obstacle avoidance and compatibility in tasks as human substitutes. However, the dynamics involved is highly non-linear, complex and unstable Il l . Many methods ranging from offline trajectory generation [2] to feedback systems based on multi-sensor-fusion are employed in the control of biped walking robots [3].\nSignificant difficulties arise commonly in defining a suitable walking pattern. Trial and error based tuning of gait parameters are reported. In some works, tuning of the many parameters of the robot gait is carried out by evolutionary control [4]. Energy or joint torque minimization can be used as criteria for the tuning [5]. This paper addresses the reference generation problem in such a way that a small number o f parameters are used in the description of the desired trajectory. The references employed are in the form of desired trajectories for the trunk of the robot. Suitable torques for the supporting leg joints for the tracking of the trunk reference trajectory are obtained by constrained optimization techniques. Locomotion is achieved by a swinging leg control scheme, independent of the body reference position tracking. The control scheme employs gravity compensation and the application of virtual potential fields to the swinging leg. This scheme of locomotion generation can be achieved with\nKemalettin ERBATUR Faculty of Engineering and Natural Sciences\nSabanci University erbatur@sabanciuniv.edu\na smaIl number of walking parameters, since it does not impose any fixed joint trajectories for the legs. The soft landing problem of robot legs is considered by a number of researchers [6-lo]. An additional advantage of the gravity Compensation technique is that soR landing of the swinging leg is achieved by simple tuning of control parameters, The next section describes the biped model used in this paper. Section 3 is a discussion of the gravity compensation technique followed by the application of this system to the biped model with simulation results in Section 4. The conclusion is presented lastly.\n62 1", + "The general biped dynamics are described by the following equation [l 11,\nwhere H , for ( i , j ) E {1,2,3} are sub-matrices of the robot\ninertia matrix. vB is the linear velocity of the robot body coordinate frame center with respect to a fixed world coordinate frame, 0, is the angular velocity of the robot body coordinate frame with respect to a fixed world coordinate frame, and B is the vector of joint displacements of the biped. The vector formed by augmenting 4 , b, , and b3 is termed as the bias vector in this dynamics equation. uEl is the net force effect and U\u20ac, is the net torque effect of the reaction forces on the robot body. u4 stands for the effect of reaction forces on the\nrobot joints. Reactive forces are generated by environmental interaction. T is the generalized joint control vector, typically consisting of joint actuation torques for a robot with revolute joints. H , , , H I 2 , H , , and H,, are 3x3 matrices. For a 12 dof robot with 6 dof at each leg, as described in the previous section; N,, is 3x12, H , is 3x12, H3, is 12x3, H,, is 12x3, and H,, is 12x12.\n111. GRAVITY COMPENSATION TECHNIQUES FOR BIPED LOCOMOTION\nThe expression in (1) reveals the importance of conti-oiling the reactive force to control the body dynamics. It is the reaction force that govems the body dynamics. This is due to the fact that the body is not directly actuated. One can yet perceive reaction forces as the control effort for the body dynamics.\nFrom (l), it can be noted that the body dynamics are given by\n(2) can be expressed in a more compact form as\n- - H++b +rdi, =-U\nwhere\nIn the proposed approach, off-line trajectories for the body coordinate frame center are generated. The reference curve in the xdirection has the shape of a ramp and the reference in the y-direction is a sinusoid. The z coordinate reference is a constant, and the orientation reference for the body is fixed and parallel to the ground. Corresponding reaction forces are generated via inverse dynamics as follows\nwhere K, and K, are parameters of the desired dynamics defined by the control designer. The relation between contact forces and the reactive force on the body is given by\n(7)\nwhere K' is a matrix relating contact forces to body reactive forces. This mahix i s 6x24 in the double support phase and 6x12 in the single support phase; and it is available for the control designer as a part of the NewtonEuler based simulation system. It can also be computed for a physical robot, on-line. f, is the contact force vector, which is 24x1 in the double support phase and 12x1 in the single support phase. It is obtained by augmenting the ground interaction forces acting on the corners of the supporting feet. Rectangular foot shape with four contact points is assumed.\nSimilar to (5), the reference body forces and torques found above can be related to references for the ground interaction forces by the equation\n(8) is solved to obtain the references for the forces at the foot. Note, however, that (8) poses an underdetermined system of equations. Further, not every solution to the system i s physically feasible due to the non-attractive nature of the contact, and no-slip condition to be satisfied for a successful walk.\nf E i = bEh f f i , f,,] itz {l,2 ,...., 8}; with this Let\ndefinition, the constraints can be formulized as follows" + ] + }, + { + "image_filename": "designv11_83_0000141_2005-01-2470-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000141_2005-01-2470-Figure2-1.png", + "caption": "Figure 2. Basic setup of the passenger car example.", + "texts": [], + "surrounding_texts": [ + "To implement the FRF-based inverse substructuring approach, a set of FRFs is measured as shown in Table 1 for just the front coupling set. For more detail information of the symbols and subscripts that are listed in the following table, one may refer to the list of symbols section at the end of this paper. Table1. Measured FRFs corresponding to the front part. Measured FRFs Size [HS]c(a)c(a) 30\u00d730 [HS]c(a)c(b) 30\u00d730 [HS]c(b)c(a) 30\u00d730 [HS]c(b)c(b) 30\u00d730 [HS]o(a)c(a) 5\u00d730 [HS]c(a)c(b) 5\u00d730 [HS]c(a)i(b) 30\u00d76 [HS]o(a)i(b) 5\u00d76 In Table 1, there are 5 response coordinates denoted by o(a) and 6 excitation coordinates denoted by i(b), which are listed in Tables 2 and 3, respectively. Similarly the required FRFs on the rear part of the vehicle are listed in Table 4. There are a total of 5 response coordinates denoted by o(a), which are exactly the same as those of the front part, and 6 excitation coordinates denoted by i(b), which are related to the rear spindle X, Y, Z directions." + ] + }, + { + "image_filename": "designv11_83_0002430_012006-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002430_012006-Figure1-1.png", + "caption": "Figure 1. 3D model of parallel mechanism", + "texts": [ + " [8] analyzed the dynamics of a new type of wall-climbing robot over obstacles based on Kane's method. Based on this, the author designs a spatial 2-DOF parallel rotating mechanism, which has a simple structure and strong load-bearing capacity. On the basis of kinematic analysis, this article used NewtonEuler method to establish the rigid body dynamics model of this parallel mechanism, and obtained its dynamic inverse solution, and laid the foundation for the follow-up research of such mechanisms. The structure of the parallel mechanism is shown in Figure 1. The moving platform is connected to the support shaft through the adapter and the yaw seat. The moving platform and the adapter, and the adapter and the yaw seat are all connected by internal pins. The yaw seat is fixed on the support shaft by screws. The moving platform is connected with the pitch slide through a hook hinge, a pitch link, and a rod end joint ball bearing. The adapter is connected with the yaw slider through a pin shaft and a yaw link. The MEMAT 2021 Journal of Physics: Conference Series 1820 (2021) 012006 IOP Publishing doi:10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000816_cphc.200600745-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000816_cphc.200600745-Figure1-1.png", + "caption": "Figure 1. Schematic depiction of the experimental setup.", + "texts": [ + " The variational electrostatic interaction in this system arises from a gradient change in the density and in the polarity of the accumulating charge on the glass substrate. The motion of the gel under the applied electric field (E) and the factors influencing the crawling velocity and the critical driving electric field were investigated. We use PVA/DMSO gel to fabricate the system because of its good dielectric property. The gel actuator was prepared from a PVA/DMSO gel strip (4.0 mm longG1.5 mm wideG2.8 mm thick) and was driven by the applied electric field of 300 Vmm 1. The experimental setup is shown in Figure 1. Figure 2 shows the self-governing motion of the gel on the gradienty charged glass substrate. The gel displays long-range linear motion with the initial crawling direction opposite to the applied electric field. In response to the change of the charge density and the charge polarity on the glass surface, the gel can actively control its crawling direction and move reversibly in the regions 7.0 0 is an arbitrary function, Q is the flow rate of the material and = \u00b11, the upper sign corresponding to diverging flow, and the lower sign to converging flow. From Eq. (1.1) it is possible to find the non-zero components of the strain rate tensor in the form (1.2) We will adopt a yield condition of the form (1.3) where and 0 are constants of the material, = ij ij/3 is the hydrostatic stress, eq = \u221a 3/2(sijsij)1/2 is the equivalent stress, ij are the components of the stress tensor and sij = ij \u2212 ij ", + "13) The left-hand side of this equation does not change sign on passing through the point = c, while the right-hand side does change sign on passing through the corresponding point = c. Consequently, a solution exists only for one side of the surface = c, and, if the surface = c is situated in the plastic zone, then the condition c = 0 should be satisfied. From relation (1.4) it is possible to obtain (1.14) Then, from Eqs. (1.2) and (1.6) we have the equation (1.15) in which, changing to differentiation with respect to using the first equation of system (1.10), we obtain (1.16) In the case of converging flow, the friction stresses are directed away from the point O (Fig. 1). Therefore, from relations (1.6) it follows that sr > 0 and sin > 0. Furthermore, from the second equation of system (1.2) we establish that < 0, and then, from relations (1.4) and (1.6), it follows that s < 0 and cos < 0. Thus, the angle can vary in the range (2.1) The angle c, which in this case is positive, lies outside the limits of the interval (2.1), and therefore the first equation of system (1.10) and then Eq. (1.11) can be solved numerically in the entire interval 0 \u2264 \u2264 0. One of the boundary conditions for the first equation of system (1", + " It can be seen that, in a very narrow region near the friction surface, high velocity gradients arise. It is interesting to note that this narrow region arises not only on the scale of the characteristic size of the process but also in the zone where asymptotic representation (2.2) is used. In the numerical solution, representation (2.2) was used in the range /2 \u2264 \u2264 /2 + 0.001. Fig. 3 also shows the behaviour of the function u( ) in this range when 0 = 30\u25e6. In the case of diverging flow, the friction stresses are directed towards the point O (Fig. 1), and therefore from relations (1.6) it follows that sr < 0 and sin < 0. Furthermore, from the second equation of (1.2) we establish that > 0, and from relations (1.4) and (1.6) it then follows that s > 0 and cos > 0. Thus, the angle can vary in the range (3.1) The angle c, which in this case is negative, lies within this range, and, as follows from Eq. (1.13), the maximum possible friction stresses arise if = c and = 0 = c. This condition is the maximum friction law in the case under examination" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002430_012006-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002430_012006-Figure2-1.png", + "caption": "Figure 2. Structure diagram and establishment of coordinate system", + "texts": [ + " The MEMAT 2021 Journal of Physics: Conference Series 1820 (2021) 012006 IOP Publishing doi:10.1088/1742-6596/1820/1/012006 guidance of the pitch slide and the yaw slide are respectively realized by guide rails fixed on the support shaft. When the mechanism is working, the yaw and pitch motors drive the respective screws to push the yaw and pitch slides forward or backward. Then the two links drive the adapter and the moving platform to rotate. In order to describe the spatial pose of the moving platform, coordinate systems are established as shown in Figure 2. The base coordinate system oxyz is established at the geometric center of the contact end face of the support shaft and the yaw seat. The x axis is perpendicular to this end face. The y axis is perpendicular to the side of the support shaft where the yaw slide is located, and the z axis is determined by the right-hand rule. The origin 'o of the moving coordinate system ' ' ' 'o x y z is located at the center of mass of the moving platform. And the 'x and 'y axis are parallel to x and y axis respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002180_012018-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002180_012018-Figure1-1.png", + "caption": "Figure 1. Belt edge gripper construction in an automatic device for high precision belt perforation: 1 \u2013 base, 2 \u2013 bottom plate, 3 \u2013 angle piece, 6 \u2013 t-slot nut, 7 \u2013 cartridge actuator, 7a \u2013 actuator piston rod, 8 \u2013 nut, 9 \u2013 pressing disc, 10 \u2013 lock nut , 12 \u2013 connecting angle, 4, 5, 11, 13 \u2013 mounting bolts, BO \u2013 optical gate, MP \u2013 perforation module", + "texts": [ + " It is critical to maintain the base side alignment at the entire length of the belt, as even a small deviation of approx. 1 mm at the distance of several dozen meters causes to exceed the allowable standard tolerances and destroys the belt. To this end, various gates and stops are to be employed to prevent excessive transverse motion of the belt. At the same time, it should be considered that the belt width may vary by at least a few millimetres along its length; therefore, one needs to facilitate automatic adjustment so that the belt is not blocked or wrapped. The critical supports are the belt edge grippers (Fig. 1) present in the perforation module MP, directly in front and after the perforating head. Apart from maintaining the base side alignment of the CAD in Machinery Design: Implementation and Educational Issues (CADMD 2020) IOP Conf. Series: Materials Science and Engineering 1016 (2021) 012018 IOP Publishing doi:10.1088/1757-899X/1016/1/012018 belt, they are to prevent its motion following a transverse movement of the perforating heads. The base of the gripper uses profile 1 with the defined height, with the bottom plate 2 being mounted at its face surface and the plate acting as a sliding surface for the belt" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000775_detc2007-35917-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000775_detc2007-35917-Figure1-1.png", + "caption": "Figure 1. An elastic beam.", + "texts": [ + " Assembling the system heat conduction equations and dynamic equations, Lagrange\u2019s equations of the first kind are derived. The Lagrange multipliers are related to temperature, kinematic and driving constraint equations, respectively. A rotating hubbeam system with simply-supported boundary condition is simulated to show the softening effect due to the temperature increase and the stiffening effect due to the transverse deformation. Finally, thermal bending of flexible beam system applied with heat flux at upper surface is investigated. 2 DESCRIPTION OF KINEMATICS OF A FLEXIBLE BEAM An elastic beam is shown in figure 1. Two coordinate systems are introduced to describe the system motion: inertial frame 0ev , and body-fixed frame e v of the beam. A material point k0 moves to point k under deformation. The absolute displacement vector of an arbitrary point of the beam is given by urr vvvvvv +=+= 00 , \u03be\u03be\u03be (1) where 0r v represents the displacement vector of the origin of ev , and 0\u03be v , \u03be v are the position vectors of the point with nloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 12/25/2017 respect to ev before and after deformation, respectively, and uv is the deformation vector" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002790_icce50685.2021.9427701-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002790_icce50685.2021.9427701-Figure5-1.png", + "caption": "Fig. 5 shows the design of a prototype model for this project using Google Sketch Up version 7.1.", + "texts": [], + "surrounding_texts": [ + "In this project, the remote control car with a size of 30cm long and 15cm width is chosen as a prototype model with four sensors. Two infrared sensors at the right side of the car are used to define the available space parking. While the infrared sensor at the front and rear of the car are used to detect the obstacle. The simulation results of several rule viewer of Fuzzy based that had been written in rule editor using Matlab is shown as in Table II. Fig. 6 shows the simulation results of condition number 1 and 4, as stated in Table II, respectively. While Fig. 7 shows the complete design of autonomous car parking using fuzzy logic which combination of infrared sensor circuit, microcontroller PIC 16F84A circuit, a motor driver circuit, and pulse width modulation (PWM) circuit, with the output of steering/direction are -0.5 for left/reverse, 0.5 for center/stop, and 1.5 for right/forward, respectively. The development of autonomous car parking using fuzzy logic is successfully designed. This prototype is able to search for a suitable parking space using an infra-red sensor and can reverse the park by itself. PWM is used to control the speed of the DC motor. So, the car can move forward/reverse into the parking space smoothly. On the other hand, this prototype also successfully to turn left and right. Authorized licensed use limited to: Carleton University. Downloaded on May 29,2021 at 14:10:54 UTC from IEEE Xplore. Restrictions apply. IV. CONCLUSION The Artificial Neural Network is an intelligence technique and broadly used in many applications. The proposed method to design an autonomous car parking using fuzzy logic which was capable of finding the empty parking space and then being able to reverse and park itself into the parking space is successfully developed. In this project, the autonomous car parking system works with PIC16F84A and being assisted by infrared sensors. ACKNOWLEDGMENT This research was partially supported by Interface Corporation, Japan, Universiti Tun Hussein Onn Malaysia (UTHM), Malaysia, and Yamaguchi University." + ] + }, + { + "image_filename": "designv11_83_0000161_6.2004-4901-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000161_6.2004-4901-Figure3-1.png", + "caption": "Figure 3. Interception scenario", + "texts": [ + " The initial positions of the target and the pursuer in the Cartesian plane are given by T0 (0, 12) , P0 (0, 0) , respectively. We assume that, initially, the pursuer is launched in the direction of the target. In this case, the pure pursuit has the following kinematics equations vr = 2 cos\u03bb \u2212 3 v\u03bb = \u22122 sin\u03bb (29) The linearization coefficients for \u03bb0 = \u03c0 2 and \u03bbf = 0 are the following a1 = \u22120.4627, a2 = 0.8106, b1 = 1. The trajectory based linearized system is the following vr = \u22120.9253\u03bb\u2212 1 v\u03bb = \u22121.6211\u03bb (30) The motions for the target and the pursuer are depicted in figure 3. The interception is accomplished successfully for both cases. Now, we consider a target with time varying path angle, with vT = 2, vP = 3, \u03bb0 = \u03c0 4 , r0 = 20 \u221a 2. The target path angle is depicted in figure 4. The initial positions for the target and the pursuer are given by T0 (0, 20) , P0 (20, 0) . The linearization coefficients are given by a \u2032 1 = \u22120.2537, a \u2032 2 = 0.9497, b \u2032 1 = 1. The final coefficients a1, a2, b1 are obtained from a \u2032 1, a \u2032 2, b \u2032 1 and the target path angle. Figure 5 shows the motions of the pursuer and the target in the Cartesian plane" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000999_bf02875875-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000999_bf02875875-Figure5-1.png", + "caption": "Figure 5. Padder roller system deflection during operation.", + "texts": [ + " To establish a mathematical model for the padder roller system, the following assumptions are presupposed. 1. Since the external force borne by the roller system is on a symmetrical plane, the force on the upper roller can be turned into an evenly distributed force per unit length, Q, on the lower roller. 2. The bearings supporting the roller are placed at the end points of cylinder pipe and the pressing pressure can be considered as the simply supported ends. 3. Because the length is much larger than the cross-section area, the Euler-Bernoulli beam modal is applied [1]. Figure 5 represents the arbitrary configuration of a roller operational system, where XYZ is the inertia reference frame. Figure 6 is a differential element of the roller corresponding to moving coordinates x, y, z. The x, y, z, are parallel to the inertial reference frame, and G is the center of the mass. When roller is rotating by a angular velocity \u2126 in z direction, the displacement in the Y direction is (1) From Euler\u2019s beam theory, the flexural angle can be expressed as (2) and in the Y direction, there is a angular velocity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000676_978-1-4020-8829-2_9-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000676_978-1-4020-8829-2_9-Figure6-1.png", + "caption": "Fig. 6. Delta robot model", + "texts": [ + " By analogy with the optimization proposed in [2], we suggest that the above-mentioned characteristic length becomes an additional parameter of our optimization problem which initially only deals with purely design parameters. The optimization of this global dexterity index (26) will be first performed on the Delta parallel robot, whose moving platform has only three translational dof. In the second subsection, a more complex model will be considered where the moving platform has six dof (three positions and three orientations): the Hunt platform. The Delta robot (see Fig. 6) is a three-dof parallel manipulator whose moving platform has only translational motions. The model is composed of three legs, each containing a parallelogram to keep the platform in a horizontal plane. The isotropy of the Delta will be evaluated over a 2 cm-sided cube as shown in Fig. 6. The design parameters of this optimization problem are: the lengths of the upper leg lA and the lower leg lB , the characteristic radii of the platform rp and the base rb, and also the distance zc between the base and the center of the workspace volume (i.e. the cube). In the objective definition, the volumetric integral of (26) is discretized into eight points corresponding to the eight vertices of the cube. The performance f to optimize is thus: f (lA, lB , zc, rb, rp) = 1 8 8\u2211 i=1 1 \u03ba ( (Jf )i ) " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001228_12.772684-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001228_12.772684-Figure7-1.png", + "caption": "Fig. 7. A schematic presentation of directed self-assembly of lamellar microdomains of block copolymers. The side wall of the topographic guiding patterns selectively wets one block of the copolymer while the bottom surface is neutral to both blocks. This type of DSA gives sub-division of the guiding pattern with length scales of the lamellar microdomains.", + "texts": [ + " Previous studies show the chemical patterns on substrates are very effective to guide the lamellar microdomains that have periodicity close to the pre-patterns. The chemical guiding pattern removes the defects in block copolymer microdomains, hence provide large area, defect-free order of microdomains. Topographic guiding patterns can be subdivided by lamellar microdomains with carefully controlled bottom and side wall interactions. In this case, the well aligned lamellae simply multiplies the frequency of the guiding patterns hence provides line/space patterns of the smaller dimensions. The schematic in Figure 7 depicts a typical topographic directed self-assembly of block copolymers. The topographic guiding pattern comprises bottom and side wall surfaces with controlled wetting property: the bottom surface is energetically neutral to the two blocks of diblock copolymers thus gives non-selective wetting to both blocks; the side wall surface has selective wetting property to locate one block of the copolymers. Under this wetting condition of the topographic guiding pattern, the lamellar microdomains of block copolymer orient perpendicular to the bottom surface and align parallel to the side wall", + " Minimal defects are expected when the periodicity of microdomains matches with the width of the topographic guiding pattern. By removing Proc. of SPIE Vol. 6921 692129-5 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 06/19/2016 Terms of Use: http://spiedigitallibrary.org/ss/TermsOfUse.aspx one block of the block copolymers within the topographic guiding pattern, higher frequencies of line patterns from lamellae which are sub-dividing the guiding pattern are obtained as shown in Figure 7. Figure 8 shows top-view SEM micrographs of the lamellar microdomains of hybrid guided with topographic patterns prepared by E-beam lithography. The funnel shaped guiding patterns were prepared by depositing silicon which wets selectively the hydrophilic phase (i.e. PEO+OS) of the hybrid. As shown in the electron micrographs, the OS lamellae (approximately 22 nm half-pitch lines) align parallel to the side wall of the guiding patterns. The micrographs also show that the number density of defects depends on the width, length and geometry of the guiding patterns" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002892_s11665-021-05931-w-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002892_s11665-021-05931-w-Figure3-1.png", + "caption": "Fig. 3 Dimensional accuracy of the as-printed AMB2018-01 bridge structures. Cloud-to-mesh (C2M) distances plotted over the 3D scans of the (a) sample B120, and (b) sample B200. (c) Histogram plots of the C2M distances, with dashed vertical lines indicating median C2M distances", + "texts": [ + " Due to negligible contributions of convective and radiative mechanisms to the heat transfer, only heat conduction was considered in this simulation, for which the governing equation is: Q \u00bc q Cp @T @t \u00fe q Cp v rT r k rT\u00f0 \u00de \u00f0Eq 3\u00de where Q is heat generated by laser beam (w/m2); q;Cp; k are the density (kg/m3), specific heat (J/(kg K)) and thermal conductivity (W/(m K)) of SS 316L, respectively; v is laser scan speed (m/s). The values for q;Cp; k were adopted from (Ref 32). Journal of Materials Engineering and Performance First, the bulk geometry of the as-printed bridge samples was evaluated. The total length and width as measured from the 3D scans were 76.08 9 5.25 mm and 76.20 9 5.55 mm for samples B120 and B200, respectively. The C2M distances are plotted in Fig. 3(a) and (b), and the corresponding C2M histograms are shown in Fig. 3(c). The median C2M distance was 53 lm for sample B120 and 167 lm for sample B200. While both samples were found to be larger than the original CAD geometry (75 mm long and 5 mm wide), the deviation from the input file was greater for sample B200, which was fabricated with a higher energy input. A 61% increase in VED correlated with a 215% increase in the median C2M distance. It is therefore evident that the dimensional accuracy of the asprinted parts was significantly affected by the process parameters used during LPBF processing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003039_icit46573.2021.9453629-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003039_icit46573.2021.9453629-Figure7-1.png", + "caption": "Fig. 7. Image of the lumbar fixation mechanism", + "texts": [ + " As the leaf spring used is more flexible than that used in AB-Wear II, its shape can be altered according to the spine posture. Therefore, wearers can perform complex motions that were not possible using AB-Wear II, such as bending and rotation. At the same time, the assistive force applied to the body is decreased because the spring is softer and highly flexible than that used in AB-Wear II. However, this may reduce assistance efficiency. Thus, we introduced the fixation component described in the next section to improve the efficiency of assistance by fixing it firmly to the body. 4) Fixing the device: Fig. 7 shows a basic schematic of the lumbar fixation mechanism of AB-Wear III. The lumbar fixation method of AB-Wear II uses a flexible cloth supporter. Thus, the device provides insufficient assistive power because of slipping. In the newly developed mechanism, the leaf spring is fixed near both hips using a side stopper to form an arc. The rigidity of the device around the vertical axis is low, and a wearer can twist the upper body (Fig.8(a)). On the other hand, the device is quasi-rigid around the transverse axis, and even if a wearer bends the lumbar region, the device does not slip easily on the body (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001940_6.2008-7116-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001940_6.2008-7116-Figure1-1.png", + "caption": "Figure 1. The variable stability projectile.", + "texts": [ + " Variable Stability Projectile Dynamic Model The system consists of two major components, namely, a main projectile body and an internal translating mass. The main projectile body is largely a typical projectile with the exception of an internal cavity that hosts an internal mass. The internal mass is free to translate within the main projectile cavity. An actuator inside the projectile exerts a force on the internal mass as well as the main projectile to move the mass inside the cavity to a desired location. A schematic of the variable stability projectile is shown in Figure 1. Note that the cavity is aligned with the body centerline (axis of symmetry). Two reference frames are used in development of the equations of motion for the system, namely the inertial and projectile reference frames. The two frames are linked by the following orthonormal transformation matrix. [ ] \u03b8 \u03c8 \u03b8 \u03c8 \u03b8 \u03c6 \u03b8 \u03c8 \u03c6 \u03c8 \u03c6 \u03b8 \u03c8 \u03c6 \u03c8 \u03c6 \u03b8 \u03c6 \u03b8 \u03c8 \u03c6 \u03c8 \u03c6 \u03b8 \u03c8 \u03c6 \u03c8 \u03c6 \u03b8 \u2212 = \u2212 + = + \u2212 P I I P I IIP P I I I I Ic c c s s J s s c c s s s s c c s c J T J c s c s s c s s s c c cK K K (1) The cavity containing the internal translating mass is fixed and extends along the projectile\u2019s axis of symmetry, in this case the IP axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001464_acc.2008.4586549-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001464_acc.2008.4586549-Figure1-1.png", + "caption": "Fig. 1. Coordinate systems", + "texts": [ + " In particular we can take uT (x) = \u2212c[x2 \u2212 \u03c6T (x1)] \u2212 \u2206W\u0303T (x) T + \u2206\u03c6T (x) T , (12) VT (x) = WT (x1) + 1 2 \u2016 x2 \u2212 \u03c6T (x1) \u20162 (13) where c > 0 is arbitrary, x = [ xT 1 xT 2 ] T and \u2206\u03c6T (x) = \u03c6T (rT ) \u2212 \u03c6T (x1), \u2206W\u0303T (x) = \u2206W\u0304T (x)[x2\u2212\u03c6T (x1)] \u2016x2\u2212\u03c6T (x1)\u20162 , x2 = \u03c6T (x1), T gT (x1) ( \u2202WT \u2202x1 )T (rT ), x2 = \u03c6T (x1), \u2206W\u0304T (x) = WT (rT ) \u2212 WT (r\u03c6 T ), r\u03c6 T = x1 + T [f(x1) + g(x1)\u03c6T (x1)]. We first introduce the following notations to describe the equation of motion of a ship. Let n, e and \u03c8 be the North and the East positions of a ship and the yaw angle (orientation) of a ship, respectively in the Earth-fixed coordinate system and let \u00b5, v and r be linear velocities in surge, sway and the angular velocity in yaw, i.e., r = \u03c8\u0307, respectively, decomposed in the body-fixed coordinate system (Figure 1). Let \u03b7 = [ n e \u03c8 ] T and \u03bd = [ \u00b5 v r ] T . In the dynamic positioning (DP) problems, the speed of a ship is quite small (\u00b5 \u2243 0, v \u2243 0, r \u2243 0) and we can assume that the damping forces are linear [2]. Hence the equation of motion of a ship can be written as \u03b7\u0307 = R(\u03c8(t))\u03bd, (14) \u03bd\u0307 = A\u03bd + Bu (15) where A = \u2212M\u22121D, B = M\u22121, R(\u03c8) = cos \u03c8 \u2212 sin\u03c8 0 sin \u03c8 cos \u03c8 0 0 0 1 is the rotation matrix in yaw, M is the inertia matrix including hydrodynamic added inertia, D is the damping matrix and the control forces and moment u = [ u1 u2 u3 ] T are provided by thrusters" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000817_1.2712953-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000817_1.2712953-Figure2-1.png", + "caption": "FIG. 2. External views of the magnets before and after viscous deformation, eight-pole radially anisotropic magnet rotor with continuously controlled anisotropy directions, and cross-sectional view of their SPMSM. The intended dimensions of the rotor are 41 mm in outer diameter, 14.5 mm in length, and 1.5 mm in magnet thickness, respectively.", + "texts": [ + "fumitoshi@jp.panasonic.com 0021-8979/2007/101 9 /09K522/3/$23.00 \u00a9 2007 American Institute of Physics101, 09K522-1 [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: 129.49.23.145 On: Thu, 18 Dec 2014 14:23:39 In the final step of the processes, cross-linking reaction of binder was obtained by heating the magnet at 170 \u00b0C for 20 min. The magnets magnetized by a pulsed field of 2.4 MA/m. Figure 2 shows external views of an 8-pole/12-slot SPMSM, a rigid magnet rotor, and the above-mentioned magnets, respectively. The typical properties of morphology, BH max value, anisotropy directions, cogging torque, and back efm were evaluated. Figure 3 shows a scanning electron microscopy SEM micrograph of the fracture surface of a preformed magnet, together with material composition of Nd2Fe14B particles and matrix of the magnet. The isolated structure between Nd2Fe14B particles and matrix including Sm2Fe17N3 fine powder was observed, suggesting that Nd2Fe14B particles were integrated together with Sm2Fe17N3 in a good hybridization" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002185_s00773-020-00792-9-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002185_s00773-020-00792-9-Figure1-1.png", + "caption": "Fig. 1 Synchronization motion diagram of ships during underway replenishment", + "texts": [ + " Herein, \u2016\u22c5\u2016 denotes the 2-norm of a matrix or vector. In this paper, s = [ xs, ys, s ]T and s = [ us, vs, rs ]T represent the position vector and the velocity vector of the supply ship, respectively. To realize the synchronization navigation control of the supply and the main ships, we introduce the concept of the virtual trajectory r , which is a trajectory transformation of m shifting by a distance lm at an angle m relative to the trajectory m of the main ship, with m counterclockwise positive. As is shown in Fig.\u00a01, the coordinate XoOYo represents the NED frame. The origin of the NED frame is chosen as certain point at sea. The axis OXo is directed to the east and the axis OYo is directed to the north. VM represents the main ship and VS represents the supply ship. XMOMYM and XSOSYS are the body-fixed frames of the main ship and the supply ship, respectively. The origins OM and OS are chosen at the centers of gravity of the main ship and the supply ship, respectively. The axes OMXM and OSXS are directed to fore, respectively. The axes OMYM and OSYS are directed to larboard, respectively. From Fig.\u00a01, the trajectory transformation between the virtual trajectory and the trajectory of the main ship is expressed as follows: (1)\u0307 = (\ud835\udf13) , (2)\ud835\udc0c?\u0307? + \ud835\udc02(\ud835\uded6)\ud835\uded6 + \ud835\udc03(\ud835\uded6)\ud835\uded6 = \ud835\uded5 + \ud835\udc1d(t), (3) ( ) = \u23a1\u23a2\u23a2\u23a3 cos ( ) \u2212 sin ( ) 0 sin ( ) cos ( ) 0 0 0 1 \u23a4\u23a5\u23a5\u23a6 , Equations\u00a04\u20136 can be written in the following vector form as: w h e r e r = [ xr, yr, r ]T , m = [ xm, ym, m ]T , = \u23a1\u23a2\u23a2\u23a3 lm cos m lm sin m 0 \u23a4\u23a5\u23a5\u23a6 , and \ufffd m \ufffd = \u23a1\u23a2\u23a2\u23a3 cos \ufffd m \ufffd \u2212 sin \ufffd m \ufffd 0 sin \ufffd m \ufffd cos \ufffd m \ufffd 0 0 0 1 \u23a4\u23a5\u23a5\u23a6 is called as the rotation matrix. Assumption 1 The unknown environmental disturbance di(t)(i=1, 2, 3) acting on the supply ship is bounded and satisfies where p\u2217 i > 0 is an unknown positive constant" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001042_mspec.1964.6500737-Figure16-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001042_mspec.1964.6500737-Figure16-1.png", + "caption": "Fig. 16. High-intensity ion source.", + "texts": [ + " Most ion beam generators consist of a source of ions, an extracting probe or electrode, an elec trostatic focusing lens and an accelerating electric field to impart the desired energy to the beam. A considerable variety of ion sources have been developed for use with high-energy particle accelerators employed in the study of nuclear reactions. One of the most efficient of these is the so-called duoplasmatron developed by M. von Ardenne. In this ion source, electrons from a hot filament are accelerated toward a hollow iron pole tip by a poten tial difference of about 100 volts. As indicated schema tically in Fig. 16, gas admitted to the cathode chamber at a few microns' pressure is ionized by electron impact in the narrow channel at the apex of the pole piece. The 1000-gauss magnetic field in the extraction channel con centrates the discharge and provides a high density of ions. Positive ions are drawn out of this region by the negatively charged hollow probe and focused into a beam by the electrostatic lens action of the accelerating gap at the exit end of the probe. Ion current densities in the ex traction channel may approach 100 amperes per square centimeter and beam currents may be as high as half an ampere" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002886_iemdc47953.2021.9449529-Figure11-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002886_iemdc47953.2021.9449529-Figure11-1.png", + "caption": "Fig. 11. Topology C with winding factors kmws = 0.99, khws = 0.92, khwf = 1: a) Geometry; b) Circuitry with passive diode rectifier; c) main and harmonic machine stator MMF over half a mechanical rotation", + "texts": [ + " Ikc(t) = I\u0302m cos(\u03c9mt+ (k\u2212 1) 2\u03c0 6 ) (10) + I\u0302h cos(\u03c9ht+ (c\u2212 1) \u03c0 2 ) Topology D exhibits a high factor ph pm = 3 which is not desirable for HESMs because it requires very high excitation frequences fexc. Furthermore, since the end-winding length \u223c 1 p is designed for the smaller main pole-pair number pm, the winding exhibits a 300% larger end-winding compared to a normal ph pole-pair winding of the same ampere turns and winding factor. This leads to high harmonic stator copper losses. Another disadvantage is that a complex rotor winding with extra slots has to be housed in the pole shoe of each salient pole (Fig. 11a), leading to geometry modifications which deteriorate the pm performance. The inherent drawbacks can be confirmed in simulation. Topology D cannot reach rated operation for realistic harmonic currents at fexc = 200 Hz. In Section III three different methods of creating additional airgap harmonics for field excitation control have been presented, with three different pole-pair combinations. The following criteria are used to assess the HESM performance. 1) Harmonic coupling factor: In order for the induced voltage to reach a certain value, a harmonic current I\u0302dh is needed which is proportional to gdh = fexc khwsNSk h wfNf p2h " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000524_s0076-6879(08)03415-0-Figure15.1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000524_s0076-6879(08)03415-0-Figure15.1-1.png", + "caption": "Figure 15.1 Principle of total internal reflection (Snell\u2019s law) (left panel) and light confinementwithin optical fiber (right panel).", + "texts": [ + " It will be necessary to consider a few of their optical properties which are most germane for sensing; comprehensive treatments may be found elsewhere (Marcuse, 1991; Snyder and Love, 1983) Classically, one can consider a fiber optic as a solid cylinder of transparent material called the core, surrounded by a layer of transparent material of slightly lower refractive index called the cladding. If light is launched into the core roughly parallel to the axis, it will tend to stay confined within the core because according to Snell\u2019s law it will be totally internally reflected at the core: cladding interface because of the differing refractive indices (Fig. 15.1): \u00f0n21 n22\u00de1=2 \u00bc NA \u00bc sin y; \u00f015:1\u00de where n1 is the refractive index of the core, n2 is the refractive index of the cladding, NA is the numerical aperture of the fiber, and y is the critical angle. Several things are readily apparent: if light is launched into the core at an angle greater than y, the light will not be reflected at the interface and will be lost into the cladding. Similarly, if the fiber optic is bent sharply enough the angle steepens and loss from the core will occur. If light in the core exits the end of the fiber it will not remain collimated, but will spread out to a degree controlled by the refractive index ratio and termed the numerical aperture of the fiber" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000251_ichr.2006.321383-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000251_ichr.2006.321383-Figure2-1.png", + "caption": "Fig. 2. The zmp can be easily related to the torque \u03c4RF . The reaction torque \u03c4X depends on the chosen application point X, but the force fRF is the same for any arbitrary point of the support polygon.", + "texts": [ + " Considering the generalized ground reaction force as applied at a fixed point of the support polygon implies that the foot must remain flat on the ground during the whole motion. This is because if the foot rolls over an edge or corner the point where the reaction force is supposed to act will not be in contact with the ground. The foot not rolling over an edge or corner can be characterized using zmp. The contact with flat foot is characterized by zmp \u2208 S where S is the support polygon. According to Fig. 2, the zmp can be computed using \u03c4RF = \u03c4zmp + rRF,ZMP \u00d7 fRF By definition the torque \u03c4zmp has only a z component, so using eq. (5) we can obtain \u03c4RFx = yRF,zmpfz \u03c4RFy = \u2212xRF,zmpfz Considering a square foot as shown in Fig. 2, the zmp condition can be stated using the reaction moment as follows \u2212yRFBfRFz \u2264 \u03c4RFx \u2264 yRFF fRFz (5a) \u2212xRFBfRFz \u2265 \u03c4RFy \u2265 xRFF fRFz (5b) where the geometry of the foot is shown in Fig. 2. The second condition for keeping the support foot on the ground is that the reaction force must be positive fRFz \u2265 0 \u21d2 z\u0308RF,G \u2265 \u2212g (6) This is because the ground can not pull down the foot. The implication comes from (3) with g = 9.81[m/s2]. The last necessary condition is to avoid slippage and depends on the friction constant between the foot and the ground, denoted by \u03bc: \u2016fRFz\u2016 fRFx \u2264 \u03bc \u21d2 x\u0308RF,G\u03bc \u2264 z\u0308RF,G (7) There are different ways to express the desired motion using the inertial forces. A simple choice is to define a desired trajectory of the center of gravity G to obtain f Ref RF and their integrals" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000368_detc2004-57064-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000368_detc2004-57064-Figure8-1.png", + "caption": "Figure 8. The Cartesian Parallel Manipulator Proposed by Kim and Tsai.", + "texts": [ + " Furthermore, since the first two serial connecting chains 3 \u2211 i 1 f ji 3 dim V m f j dim Am f j j 1 2 while for the third serial connecting chain 4 \u2211 i 1 f3i 4 dim V m f 3 dim Am f 3 3 It follows that Fp \u2211 f ji 3 \u2211 j 1 Vm f j 1 Hence, there is a passive degree of freedom; i.e. there is a selfmotion associated with the third serial connecting chain, and the motion has one degree of freedom. 5.3 Kim and Tsai\u2019s Cartesian Parallel Manipulator. Consider the cartesian parallel manipulator proposed by Kim and Tsai, [5], shown in Figure 8.7 The manipulator is formed by three similar serial connector chains. Each chain is formed by a prismatic pair and three revolute pairs, the direction of the prismatic pair and the axes of the revolute pairs are parallel and generates the Lie subalgebra associated with the Scho\u0308nflies group RP 4 e\u0302 j , moreover the unit vectors e\u03021 e\u03022 e\u03023 are linearly independent; i.e. Vm f j Am f j rp 4 e\u0302 j j 1 2 3 7The same architecture has been also proposed by Carricato and ParentiCastelli, [17; 19] and Kong and Gosselin, [18]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002526_j.matpr.2021.02.507-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002526_j.matpr.2021.02.507-Figure2-1.png", + "caption": "Fig. 2. Block illustration of the hand layout methodology.", + "texts": [ + " The composite sample is developed using a hand layout methodology. There are several methods available for the manufacturing of composite materials. Of all, hand layout techniques are commonly used by researchers. The purpose of choosing a hand layout methodology is attributed to its financial aspect. Indeed the simplicity of which the composite is designed by a hand layout methodology makes this approach common within investigators. The detailed description of the hand layout methodology can be seen in Fig. 2. The mould that is employed in the manufacture of composite specimens is composed of rectangular apparatus of 295 mm in length, 235 mm in width and 7 mm in thickness both from vertical and horizontal apparatus and is covered with a cleaner (Acetone). The purpose of this outer portion of apparatus is to shield the sample in order to prevent particles from getting in to the fortified composites mostly in the course of curing process as well as to impose a constant pressure. As once composite sample is manufactured using a hand layout process, specimens for different types of specimens" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003272_acc50511.2021.9483178-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003272_acc50511.2021.9483178-Figure3-1.png", + "caption": "Fig. 3. Displacement parameters for flywheel inverted pendulum model", + "texts": [ + " The armature of this motor provides sufficient rotational inertia so an additional flywheel is not needed. Neither the motor-shaft position nor its speed are measured. The angle of the pendulum rod is sensed with a Maxon five hundred counts-per-turn optical quadrature encoder. The two base parts and the pendulum rod are 3D printed. They are connected through a pair of 608ZZ ball bearings. The shaft protruding from the base part on the right in Fig. 2 is part of the printed pendulum rod and is what the encoder connects to. The displacement parameters for the FIP are shown in Fig. 3. The angular displacement of the pendulum rod from vertical is denoted with \u03b81. The displacement of the motor shaft from a position aligned with the pendulum rod is denoted with \u03b82. The distance between the pendulum axis and the motor shaft is given as l1 and the distance between the pendulum axis and the center of mass of the rod is lC1. Applying Newton\u2019s second law for rotation to the pendulum rod and combining it with the results of applying Newton\u2019s second law (for translation) to the motor armature yields (m1lC1 +m2l1) (gS1)\u2212b1\u03b8\u03071 +b2\u03b8\u03072\u2212\u03c4m = ( J1 +m2l 2 1 ) \u03b8\u03081 (1) where mi is the mass of the i-th link (in this case the pendulum rod and the motor), g is the gravitational acceleration 2CAD files are available at https://github" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001050_ijtc2008-71009-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001050_ijtc2008-71009-Figure2-1.png", + "caption": "Figure 2 : CAD Model for 30% hollow rolling element.", + "texts": [], + "surrounding_texts": [ + "The contact stresses in hollow members are often calculated by using the same equations and procedures as for solid specimens. This approach seems to be incorrect. The Hertzian theory of contact is based on several assumptions. One is that the profiles of the two bodies are continuous and can be represented to good approximation by a polynomial of second degree. Hertzian theory does not take into account a situation where the cross sections of either or both of the bodies in contact are multiply connected. For hollow rolling element no method is available for the calculation of contact stresses, deformation and all other important terminologies which we have discussed for the solid rolling element. So with the help of FEA we can get the required results. Not only doing the analysis but also determining the optimum value of hollowness (Ratio of inner diameter to the outer diameter of the roller) of the roller for which stiffness will be maximum, is the main theme of this paper. Investigations have been made for hollow rollers in pure normal loading. Different hollowness percentages ranging from 30% to 90% (in the step of 10%) have been analysed in FEA software to find the optimum percentage hollowness which gives longest fatigue life and finally maximum stiffness. Earlier we have done the analysis in the finite element package IDEAS for solid rolling element bearing in different steps. Same steps are carried out for analysis of difference hollowness of hollow rolling element bearing. Following diagram shows the CAD model for hollow rolling element bearing for the 30% hollowness. Copyright \u00a9 2008 by ASME data/conferences/ijtc2008/70344/ on 02/16/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Downloa" + ] + }, + { + "image_filename": "designv11_83_0002325_012023-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002325_012023-Figure1-1.png", + "caption": "Figure 1.Quad tilt rotor with partial tilt wing.", + "texts": [ + " Based on the above research background, in order to solve the problems of QTR time-varying, nonlinear control and ADRC parameter tuning, we use PSO algorithm to self-tune the parameter of ADRC. Firstly, we establish the dynamic model, and then we analyze the ADRC and PSO algorithms. Finally, we simulate and verify the self-tuning algorithm by comparing the ADRC control algorithm before and after optimization based on PSO algorithm in time and frequency domain. The quad tilt rotor with partial tilt wing in this paper is shown in figure 1, including four groups of propellers, front and rear wings, fuselage, elevator, motors, tilting mechanism, undercarriage and flight control system. Both ends of the front and rear wings are designed with a tilt nacelle. The tilt wing is connected to the nacelle and turns with the tilting of the propeller in the nacelle. The propeller is modeled according to the Goldstein vortex theory, and figure 2 shows the velocity and force acting on the blade element. R m is radius, r is the distance from hub center to any point of propeller profile, x is dimensionless value of r , is propeller solidity, /rad s is rotational speed of propeller, /V m s is inflow velocity, is inflow ratio, is blade element angle, T is blade element inflow angle at propeller tip, EV is resultant velocity, ,a t are axial and circumferential induced velocity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002226_012033-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002226_012033-Figure2-1.png", + "caption": "Figure 2. Experimental research complex splined coupling of the Mi-26 helicopter transmission: a) \u2013 coupling assembly at the stand; b) \u2013 barrel splines; c) \u2013 annular thimble splines; 1 \u2013 vibration acceleration sensors; 2 \u2013 fixed fuselage frame", + "texts": [ + " I - splined coupling with flanges in the bearing; 1, 3 - hollow shafts; 2 - bearings of intermediate hearings; 4 - temperature sensors The operational reliability and efficiency of heavily loaded helicopter couplers is determined not only by speed and load parameters, but also by physico-mechanical, physico-chemical, and tribological parameters of frictional processes in frictional subsystems. The basic purpose of the tail shaft is to transfer the rotating moment from the main gearbox to the tail rotor by means of series-connected elastic elements having certain masses and moments of inertia. The splined coupling is designed in such a way that the annular thimble rotates in a bearing mounted on the frame (Figure 2). The analysis of the research has shown that the main faults of the spline couplings of the helicopter transmission are the following: the formation of cracks and delamination of the rubber bearing cage; leakage of lubricant, which stimulates overheating of the coupling and its bearing; deformation and wear products generation of the coupling components; the formation of a sideways clearance in the coupling joints; misalignment of tail shaft bearings; increased outrun of the shaft tube, as well as axle fracture or shaft twisting", + " 1) are unable to inform the pilots in a timely manner about any emerging problems. A more advanced splined joint diagnostics technology is required in order to allow real-time identification of any emergency situations, which would improve the safety of long-distance piloting. Dynamics of Technical Systems (DTS 2020) IOP Conf. Series: Materials Science and Engineering 1029 (2021) 012033 IOP Publishing doi:10.1088/1757-899X/1029/1/012033 The tribological system of the spline joint of the transmission of the Mi-26 helicopter (Fig. 2) refers to systems that are characterized by nonlinear interconnected physico-mechanical, thermophysical, tribochemical, and load-velocity factors, as well as environmental ones. A feature of the coupling tribosystem operation spline joint is in the following. Its operation takes place under the influence of significant vibrations from the main engine and the entire helicopter as a whole. Therefore, modern technical means of measuring physical quantities, digital transmission and signal processing equipment should be used to achieve the goal" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000812_icit.2008.4608539-Figure11-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000812_icit.2008.4608539-Figure11-1.png", + "caption": "Fig. 11. Clearance face and representing lines(Up view)", + "texts": [ + " The three-dimensional trajectory surface of the cutting edge can be obtained by introducing homogeneous transformation matrix. This figure shows the three-dimensional trajectory surface of the cutting edge during the milling process. The tool is moved rightward. The intersection between this three-dimensional trajectory surface and the metal surface corresponds to the geometry of the generated dimple. A. Discussion on experiment 2 Deformation of the dimples mentioned in section 2 can be explained by focusing on the clearance surface of the milling tool. Fig.11 shows the clearance surface of the tool. Lines L1 through L5 are parallel to the cutting edge and equally spaced from neighboring lines. The distance between Line Ln and the cutting edge is 0.1R\u00d7n. Using the numerical model, the three-dimensional trajectory surfaces of the cutting edge and lines L1 through, L5 are calculated in two cases, =15\u00b0 and =30\u00b0. The intersections of the trajectory surfaces and the workpiece surface are obtained as shown in Fig.12. These simulation data show that cutting with a tool of clearance angle =15\u00b0 causes collisions with the workpiece surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001413_j.jsv.2008.02.056-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001413_j.jsv.2008.02.056-Figure4-1.png", + "caption": "Fig. 4. Meridional stress diagram for H 02.", + "texts": [ + " (19) and (20) the expression for finding d1H1 as 2pR1H 0 1d1H1 \u00bc Z v 2\u00f0H1\u00de s \u00f0H 01\u00dedv \u00bc 2pR Z A H1R1 E c Iz0 \u00f0y cos a z sin a\u00de \u00fe 1 A \u00fe n sin a Rt y \u2018 \u00fe 1 2 H 01R1 c Iz0 \u00f0y cos a z sin a\u00de \u00fe 1 A dydz\u00fe Z A H1R1 E sin a Rt y \u2018 \u00fe 1 2 \u00fe nc Iz0 \u00f0y cos a z sin a\u00de \u00fe n A H 01R1 sin a Rt y \u2018 \u00fe 1 2 dydz . (21) Upon performing the integrations in Eq. (21), one finds the expression for the displacement in the direction of coordinate 1 under the H1 loading as d1H1 \u00bc H1RR1 E c2 Iz0 \u00fe 1 A \u00fe nc\u20182 sin a cos a 6RIz0 \u00fe n sin a Rt \u00fe \u2018 sin2 a 3R2t . (22) To obtain our displacement at coordinate 2 due to the H1 loading, we impose the virtual loading H02 at coordinate 2 prior to application of the H1 loading. The meridional stress diagram under the H02 loading is shown in Fig. 4. The virtual stresses sx induced by the H02 loading are found from Eqs. (1), (8), and (10) to be sx\u00f0H 0 2\u00de \u00bc H 02R2d Iz0 \u00f0y cos a z sin a\u00de H 02R2 A . (23a) ARTICLE IN PRESS T.A. Smith / Journal of Sound and Vibration 318 (2008) 428\u2013460 439 It is seen from Fig. 4 and previous discussions relative to Eq. (14) that the stresses sy are given quite closely by sy\u00f0H 0 2\u00de \u00bc H 02R2 sin a Rt y \u2018 1 2 . (23b) By applying the principle of virtual work, we find our expression for determining d2H1 as 2pR2H 0 2d2H1 \u00bc Z v 2 \u00f0H1\u00des\u00f0H 02\u00dedv: (24) By substituting Eqs. (20) and (23) into Eq. (24) and performing the integrations, one finds d2H1 \u00bc H1RR1 E cd Iz0 1 A \u00fe \u2018 sin2 a 6R2t . (25) To find the rotational displacement d3H1 due to the H1 loading, we apply the virtual loading T 03 at and in the direction of coordinate 3 before application of the H1 loading" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000522_bfb0119424-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000522_bfb0119424-Figure4-1.png", + "caption": "Figure 4. The proportional contribution of the linkage joints to friction losses measured at the finger-tip. The first percentages are for the friction opposing the actuators while the numbers in brackets are for the back-driven case.", + "texts": [ + " For the graphs in Figure 3 there was no special tuning of the coefficients of friction to achieve the good fit shown. Thus these figures have been produced without information that is unavailable to the designer or cannot be reasonably estimated. Nevertheless, the results are sensitive to those coefficients which contribute most significantly to the reduction in the loads measured at the finger-tip. For one particular configuration and loading we can breakdown the contribution of the friction in each joint to the overall lowering of performance. Figure 4 shows a diagram of the finger with the approximate proportion of dif- ference between the frictionless and friction model. The bracketed percentages apply to the case where friction loads assist the actuator forces in supporting the finger-tip load. It can be seen that three joints account for about 80% of the performance loss owing to friction. An assessment of the reactions in these joints shows that the moment reactions lead normal forces that are an order of magnitude greater that the direct reactions, and this suggests one way in which the design of the linkage can be improved" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000419_10236210490883111-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000419_10236210490883111-Figure1-1.png", + "caption": "FIGURE 1 Geometry and configuration of two-axial groove bearing.", + "texts": [ + " Under these conditions it is important to determine the deviation of the coefficients along the journal orbit to predict the range of validity of the linearized coefficients along the journal orbit. Hence, there is a need to evaluate the coefficients at the unsteady state position along the journal locus with respect to the static equilibrium position rather than the evaluation of coefficients with respect to the corresponding perturbation amplitudes, as in the case of traditional approach. The two dimensional isoviscous, laminar and incompressible Reynolds lubrication equation under dynamic conditions for the two-axial groove bearing shown in Figure 1 takes a form: \u2202 \u2202\u03b8 ( H 3 12 \u2202P \u2202\u03b8 ) + ( R L )2 \u2202 \u2202 Z ( H 3 12 \u2202 P \u2202 Z ) = 1 2 \u2202 H \u2202\u03b8 + \u2202 H \u2202T [1] Under dynamic conditions, the transient motion of the journal is defined by the journal center position and velocity, such that the film thickness is expressed as: H = 1 + X cos \u03b8 + Y sin \u03b8 [2] The fluid film reaction forces along the locus of the journal trajectory are functions of journal center displacements and velocities. The first order perturbation of pressure and film thickness about the unsteady state journal position under dynamic conditions is: P = P + Px X + Py Y + Px\u0307 X\u0307 + Py\u0307 Y\u0307 [3] H = H + X cos \u03b8 + Y sin \u03b8 [4] Substituting the first order perturbation of pressure (Equation 3) and film thickness (Equation 4) into the Reynolds equation (Equation 1) yields one unsteady pressure (Equation 5) and four unsteady pressure gradients (Equations 6\u20139)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002197_j.matpr.2020.12.125-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002197_j.matpr.2020.12.125-Figure3-1.png", + "caption": "Fig. 3. Gear pair assembly and mesh discretization.", + "texts": [ + " The keyword *DAMPING_GLOBAL is utilized to include the system damping [10], which depends on the parameter Ds = 2xmin (fundamental frequency of the structure xmin = 50). In order to create elements having good quality criteria, hexahedral solid elements were generated using the mapped mesh option of LS-PRESPOST. Many variables should be taken into account before meshing; such as mesh density, meshing quality, mesh distribution, etc. at the region of contact and also at the tooth root. A fine mesh is utilized in order to capture the results more accurately. Fig. 3 highlights the gear pair contact assembly. Fig. 4 highlights the zoomed view of the mesh transition from coarser to finer elements at the involute contact region which extend till gear tooth fillet. Reduced integrated elements formulation was used to reduce the total time taken for solving the complete analysis. Also *CONTROL_HOURGLASS was included to reduce the possibility of occurrence of hourglass modes. Care was taken to make sure that hourglass energy was less than 5% of the total energy of the system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003135_012006-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003135_012006-Figure5-1.png", + "caption": "Figure 5. Coordinate System of Index Finger", + "texts": [ + "1088/1742-6596/1965/1/012006 parallel to the metacarpophalangeal joint, the coordinate system 3 is located at the third hinge of the finger, with the X-axis parallel to the proximal end of the interphalangeal joint, and the coordinate system 4 is located at the end of the finger, with the X-axis parallel to the distal end of the interphalangeal joint. The Z-axis of all the coordinate systems is parallel to the rotational axis of the rotational joint where the coordinate system is located. According to the composition of the index finger, the index finger can be equated to a three-degree-offreedom tandem mechanism. The schematic diagram of the motion of the index finger is shown in Figure 5. From Figure 5, it can be seen that the index finger consists of three rigid links OA, AB, and BC. C1, C2, and C3 are the centroids of rigid links OA, AB, and BC, respectively. The angle between axis X1 and rigid link OA is defined as \u03b81, the angle between axis X2 and rigid link AB is defined as \u03b82, and the angle between axis X3 and rigid link BC is \u03b83. The lengths of rigid links OA, AB, and BC are l1, l2, and l3, respectively. In this paper, the D-H method is used to establish the kinematic model of the index finger", + " ()rand St j i' i i j i X X X X X X (8) iv) Random behavior Random behavior refers to the behavior of choosing a random state in the visual field and then moving in that direction. It is the default behavior of foraging behavior, i.e., the next position Xi' of Xi is r V ' i iX X (9) where r is a random number in the interval [-1,1] and V is the visual distance range. IFEMMT 2021 Journal of Physics: Conference Series 1965 (2021) 012006 IOP Publishing doi:10.1088/1742-6596/1965/1/012006 For the index finger of the robotic hand shown in Figure 5, the given input variables are as follows. l1 = l2 = l3 = 1.5 cm. The end positions are px=2.6, py=3, pz=0, and the movable angle ranges of the three rigid links are \u03b81[0, 90], \u03b82[0, 90], \u03b83[0, 90]. Through solving the optimization model described by equation (5), the output variables \u03b81, \u03b82, \u03b83 are obtained, and the parameters of the AFSA are shown in Table 2. 4.2.Numerical Simulation of Position Inverse Solution In this section, numerical simulation is performed for the inverse kinematics of the bionic hand based on the AFSA" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003438_j.indcrop.2021.113890-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003438_j.indcrop.2021.113890-Figure5-1.png", + "caption": "Fig. 5. Response surface for the effects of a) ultrasound power versus buffer solution/oil mass ratio (B/O), b) ultrasound power versus catalyst/substrate mass ratio (C/S), and c) catalyst/substrate mass ratio (C/S) versus buffer solution/oil mass ratio (B/O) on the yield of enzymatic hydrolysis of crambe oil using oil-free seeds.", + "texts": [ + "55 \u2217 Po \u2217 B/O \u2212 0.20 \u2217 Po \u2217 C/S \u2212 2.75 \u2217 B/O \u2217 C / S (5) Through the use of the software Maple\u00ae, the derivative of the above equation was found with the conditions of 67.8 % ultrasound power, 1.67 mass ratio buffer solution/oil, and 0.06 mass ratio of catalyst/ substrate, yielding 69.50 %. The experimental yield result found under the above conditions was 71.50 \u00b1 2.5 %, confirming that within the investigated conditions, the model can satisfactorily describe the yield of the hydrolysis reaction. In Fig. 5 are presented the response surface plots for enzymatic hydrolysis yield of the crambe oil with oil-free seeds as a catalyst. In Fig. 5 (a), it is observed that the highest yield region is in the highest ultrasound power and buffer solution/oil mass ratio region. When the power effect is evaluated together with the amount of catalyst, Fig. 5 (b), the result is different from the fresh seeds. When a higher power was needed for higher values of catalyst ratio, the two variables had little correlation, and a parabolic effect was observed for the catalyst/substrate ratio effect, with the yield reaching a maximum value at a ratio of approximately 0.05. After that, a small drop occurred, indicating that in this range, the substrate is already saturated with the enzyme, and the increase only hinders the reaction by increasing the viscosity of the medium. In Fig. 5 (c), an interesting effect is observed, the higher the amount of buffer solution added to the reaction, the lower the amount of catalyst required for a good yield. The opposite was also observed, higher amount of catalyst required less amount of solution. With these results, it is possible to perform a cost analysis to determine the best conditions under particular situations. In the optimal experimental conditions obtained through Eq. 5 (68 % power, 1.7 B/O, and 0.06 C/S), a kinetic experiment was performed, and the result is shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001162_fedsm2007-37066-Figure21-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001162_fedsm2007-37066-Figure21-1.png", + "caption": "Figure 21 Side views of instantaneous relative flow patterns around a ski jumper", + "texts": [ + " In this paper, both vertical component Fv and horizontal component Fh of the calculated fluid force acting on the ski jumper including a pair of skis are respectively normalized as Fv /(0.5\u03c1 U2) m2 which is called \u201clift force area\u201d and Fh /(0.5\u03c1 U2) m2 which is called \u201cdrag force area\u201d in this paper. In order to examine the influence of the posture of a ski jumper, the simulation of flows around a ski jumper with different posture styles (V-shape style, parallel style, delta style) as shown in Fig. 20 were performed. 7 Copyright \u00a9 2007 by ASME Terms of Use: http://www.asme.org/about-asme/terms-of-use D Figure 21 shows a couple of side views of instantaneous relative flow patterns around the ski jumper for the V shape style at the time of t = 0.04 s and 0.25 s which are expressed by relative velocity vectors at positions of vortex elements at each moment. It is clearly observed that remarkable separation of the relative flow appears only around the legs at t =0.04 s because both skis are parallel and attack angle of the skis to the relative flow is still very small. On the other hand, at t =0.25 s, the flow separation becomes considerable around the whole configuration including the pair of skis because skis move from parallel phase to the V-shape phase and the attack angle becomes large" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003623_s00170-021-07887-6-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003623_s00170-021-07887-6-Figure4-1.png", + "caption": "Fig. 4 a Overview of test specimens and b schematic diagram of the test setup", + "texts": [ + " Printing below 230 \u00b0C led to underextrusion that manifested as various discontinuities in the deposited layers (e.g., openings, nonuniform layer thickness) (Fig. 3b). Currently there are no mechanical test standards specifically dedicated to 3D printed polymer parts. Hence, the tests in AM community are conducted by using ASTM or ISO standards that are applicable to plastics [6]. Tensile, compressive, and flexural tests under quasistatic loading conditions were conducted by following ISO 527-2, ISO 604, and ISO 178 standards, respectively. Figure 4a presents the geometry and dimensions of the test specimens. CAD models of the specimens were prepared in SolidWorks and then exported as STL files to the slicing software Cura 3.6.0. Five specimens were printed for each test case to derive mean values \u00b1 standard deviation for the strengths and moduli of elasticity. The experiments were performed at room temperature using dual-column universal testing machine Tinius Olsen H25KT (Fig. 4b). Strains were determined by means of video extensometer Tinius Olsen VEM 300 that measured the displacements of gauge marks on the specimens. Tinius Olsen software Horizon was used for machine control, data acquisition, and analysis. Tensile specimens were clamped in manual wedge action grips. Applied force, crosshead displacement, and strains were measured until specimen fracture. Following the conditions specified in ISO 527-2, the dumbbell-shaped type 1B specimens were subjected to uniaxial tensile loading", + " The tensile strength was determined by calculating the ratio between the peak force measured during testing and the initial cross-sectional area of the specimen. Tensile modulus was estimated as a slope of a secant line between the strains of 0.05 and 0.25% recorded at a loading rate of 1 mm/min: Et \u00bc \u03c32\u2212\u03c31 \u03b52\u2212\u03b51 \u00f01\u00de where \u03c31 and \u03c32 are stress values corresponding to strain levels of \u03b51 = 0.0005 and \u03b52 = 0.0025, respectively. Round bar-type specimens of two different lengths were printed for the compression testing in accordance with ISO 604 standard (Fig. 4a). Longer type A specimens were used for the determination of compression modulus, while shorter type B specimens were used to evaluate the strength. For both cases, the loading rate was kept constant at 1 mm/min. The compressive strength and modulus were calculated as in the tensile case. Rectangular bar-type specimens were printed for the threepoint bending testing in accordance with the ISO 178 standard. By following Method A in the standard, a constant loading rate of 2 mm/min was used for the determination of the flexural stress\u2013strain curves" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000184_6.2004-4872-Figure10-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000184_6.2004-4872-Figure10-1.png", + "caption": "Figure 10 Plan view of the ICE vehicle", + "texts": [ + " The control effectors include elevons, pitch flaps, all moving tips, thrust vectoring, spoiler slot deflectors, and outboard leading edge flaps. The conventional control effectors are defined as the elevons, pitch flap, and leading edge flaps. The innovative control effectors are defined as the thrust vectoring, all moving tips and spoiler slot deflectors. Challenges associated with control using the all moving tips and spoiler slot deflectors include zero lower deflection limits, strong multi-axes effects and effector interactions. The state, output and input vectors in these linear models are defined below. Figure 10 shows the vehicle in plan view. ]rpvqwu[x bbb T \u03c6\u03b8= ]aaarpqv[y cgncgycgxssb T \u03c6\u03b2\u03b8\u03b1= ]20DE10DE12DE2DE 19DE9DE15DE5DE4DE13DE3DE[u T = Vehicle models for the two flight conditions follow. The vehicle model possesses one unstable pole at s = 0.3343 for the Mach No. = 0.5, altitude = 15,000 ft flight condition condition, and three unstable poles at s = 0.2055, 1.05 and 0.0186 for the Mach No. = 0.9, altitude = 35,000 ft flight condition. American Institute of Aeronautics and Astronautics Mach No. = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000332_50003-4-Figure3.8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000332_50003-4-Figure3.8-1.png", + "caption": "Fig. 3.8. Equality of the two pure slip characteristics for an isotropic tyre model if plotted against the theoretical slip.", + "texts": [ + "10) for the lateral deflection at pure side slip (try = tana)\" v - ( a - X ) a y (3.22) Comparison of the equations (3.21) and (3.22) shows that the longitudinal deformations u will be equal in magnitude to the lateral deformations v if try = tana equals trx = x / ( l + x ). For equal tread element stiffnesses (Cpx = Cpy ) and friction coefficients (,Ux= ~/./y ) in lateral and longitudinal directions, the slip force characteristics in both directions are identical when tan a and x/(1 +x) are used as abscissa, cf. Fig.3.8. Also, Eq.(3.11) holds for the longitudinal force Fx if the subscripts y are replaced by x and tana by x. Obviously, total sliding will start at trx = x / ( x + 1) = +-l/Ox or in terms of the practical slip at\" 104 THEORY OF STEADY-STATE SLIP FORCE AND MOMENT GENERATION -1 K - ICsl - 1 +_.0 (3.23) X with 2 0 - 2 C x a x 3 / ~ F (3.24) Linearisation for small values of slip x yields a deflection at coordinate x: u - (a - x) x (3.25) and a fore and aft force F - 2 C x a 2 x (3.26) with Cpx the longitudinal tread element stiffness per unit length", + "45) where analogous to expressions (3.6) and (3.24) for the isotropic model parameter 0 reads: 2 c a 0 - 0 = 0 - 2 p (3.46) y x 3 ~ / f From Eq.(3.45) the slip Osl at which total sliding starts can be calculated. We get analogous to (3.9) 1 O ' s l - 0 (3.47) The magnitude of the total force F= IF[ now easily follows in accordance with (3.11): F - p F (1 - 23) - / . t F { 3 0 a - 3(00) 2 + (00\") 3 } for O\" _ O'sl (3.48) F - p F for tr___ %t and obviously follows the same course as those shown in Fig.3.8. The force vector F acts in a direction opposite to V~ or -tr. Hence F _ F ~r -~ (3.49) from which the components Fx and Fy may be obtained. The moment - M z is obtained by multiplication of Fy with the pneumatic trail t. This trail is easily found when we realise that the deflection distribution over the contact length is identical with the case of pure side slip if tanaeq - ~ (cf. Fig. 3.10). Consequently, the formula (3.13) represents the pneumatic trail at combined slip as well if Oy~ry is replaced by 0a" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001838_nems.2008.4484426-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001838_nems.2008.4484426-Figure2-1.png", + "caption": "Fig. 2 shows schematic drawing (a) and photomicrograph (b) of proposed non-enzymatic glucose sensor on silicon substrate. As shown in Figure 2 (a), the proposed nonenzymatic glucose sensors have three electrodes which are comprised of working electrode (WE; nanoporous Pt electrode), counter electrode (CE; plane Pt electrode), and reference electrode (RE; Ag/AgCl electrode or plane Pt electrode).", + "texts": [ + " 1 shows surface morphology of fabricated nanoporous Pt electrode analyzed with AFM (atomic force microscopy, XE-100, Park Systems, Korea). As shown in Fig. 1, RMS (root mean square) values, which are revealed the surface roughness, are approximately 3.041 nm. Figure 1. Surface morphology of fabricated nanoporous Pt electrode (analyzed with AFM) This work was supported by the IT R&D program of MIC/IITA [2005-S093-03, Development of Implantable System Based on Biomedical Signal Processing]. *Contact author: Tel. +82-2-940-5113; Fax. +82-2-942-1502; e-mail. jaepark@kw.ac.kr 704978-1-4244-1908-1/08/$25.00 \u00a92008 IEEE. (a) (b) Figure 2. Schematic drawing (a) and photomicrograph (b) of proposed nonenzymatic glucose sensor on silicon substrate. solution and 0.1 M PBS solution containing 0.1 M glucose solution at the scan rate of 20mV/s. This experiment was performed in order to find an appropriate potential applied for the detection of glucose. This graph shows that the optimum potential is around 0.4V and the sensor with Pt RE has larger response current than the other one. 1.0 0.8 0.6 0.4 0.2 0.0 -100 -50 0 50 100 C ur re nt [ A ] Voltage [V] Ag/AgCl RE in PBS Ag/AgCl RE in Glucose Pt RE in PBS Pt RE in Glucose Figure 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000796_j.mcm.2006.04.010-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000796_j.mcm.2006.04.010-Figure1-1.png", + "caption": "Fig. 1. Top view of the car.", + "texts": [ + "his work deals with the modelling of the motion of a three-wheeled car (see Fig. 1). The front wheel of the car is constrained to roll on a given plane curve.1 In the model derived here, the motion of the car is controlled by two torques: a pedalling torque and a steering torque, both acting on the front wheel. c\u00a9 2006 Elsevier Ltd. All rights reserved. Keywords: Three-wheeled car; Nonholonomic constraints; Steerable front wheel; Plane curve This work deals with the modelling of the motion of a three-wheeled car moving on the horizontal plane, where the front wheel is confined to roll on a given plane curve. The car is composed of three rigid wheels, a rod connecting the point O4 to point 3, and an axle going through the centers of Wheels 1 and 2 (see Fig. 1). The above-mentioned rod and the axle are referred to here as the frame. Wheels 1 and 2 are identical. A pedalling torque \u0393\u03c83 and a steering torque \u0393\u03b4 both act on the front wheel, that is, Wheel 3. It is assumed here that: (i) The motor that rotates Wheel 3 is embedded in the frame at the center of Wheel 3. (ii) The steering mechanism that generates \u0393\u03b4 is embedded in the frame. All wheels are roll without slipping on the above-mentioned plane. This work is, to some extent, a continuation of [2,3]", + " 1 A plane curve in R3 is a curve that lies in a single plane of R3 [1]. In this work, a plane curve is refered to a curve that lies in the horizontal plane. 0895-7177/$ - see front matter c\u00a9 2006 Elsevier Ltd. All rights reserved. doi:10.1016/j.mcm.2006.04.010 Let I, J and K be unit vectors along an inertial coordinate system. Denote by ka ka = cos\u03c6I + sin\u03c6J, (1) a unit vector along the rod (embedded in the frame) that connects the point O3, which is above Wheel 3 (see the definition of r3, Eq. (5) below) to the point O4 (see Fig. 1). Also, let ja be a unit vector along the axle (embedded in the frame) that passes through the centers of Wheels 1 and 2: ja = \u2212 sin\u03c6I + cos\u03c6J. (2) The set B = (ka, ja,K) acts here as a body axis for the car\u2019s frame. That is, the system B is attached to the point rO (that is, the point O; see Fig. 1), which is the center of mass of the frame and is moving with it. The vector rO is given here by rO = xI + yJ + aOK, (3) where aO denotes the radius of Wheels 1 and 2, and bO denotes the radius of Wheel 3, aO = bO + D, D \u2265 0. Denote by ri the center of wheel i , i = 1, 2, 3, respectively. Then, r1 = rO \u2212 L11ka \u2212 L O2 ja, r2 = rO \u2212 L11ka + L O2 ja, (4) r3 = rO + L12 ka \u2212 DK, and r4 = rO \u2212 L11ka, L O2 = L2 2 , L1 = L11 + L12, (5) where r2 \u2212 r1 = L2 ja, r3 + DK \u2212 r4 = L1ka, rO \u2212 r4 = L11ka, r3 + DK \u2212 rO = L12 ka, (6) and r4 represents the location of O4 (see Fig. 1). Then, by using Eqs. (4) and (5), the following equations are obtained: vO = drO dt = x\u0307I + y\u0307J, (7) v1 = dr1 dt = vO \u2212 L11\u03c6\u0307ja + L O2\u03c6\u0307ka, v2 = dr2 dt = vO \u2212 L11\u03c6\u0307ja \u2212 L O2\u03c6\u0307ka, (8) v3 = dr3 dt = vO + L12\u03c6\u0307ja, v4 = dr4 dt = vO \u2212 L11\u03c6\u0307ja . (9) Note that the planes of Wheels 1, 2 and 3 are always vertical to the plane. It is assumed here that the plane curve on which the front wheel is rolling is represented by rW (u) = x(u)I + y(u)J, u \u2208 [u1, u2), u1 < u2 (10) and it is also assumed here that the shape of the curve (given by (10)) and the radius bO of the front wheel are such that, during the wheel\u2019s motion, at each instant there is only one point of contact between the wheel and the curve" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002337_s11665-021-05550-5-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002337_s11665-021-05550-5-Figure2-1.png", + "caption": "Fig. 2 (a) Schematic of the process of LSF; (b) the prepared sample by LSF", + "texts": [ + " The system comprised a 4kw CO2 laser, a five axis four linkage CNC worktable, an inert atmosphere protection chamber, a high-precision adjustable powder feeder and a powder feeding nozzle (Ref 17-19). In the experiment, the alloy powder was sent out by the powder feeder into the laser melting pool through the side feeding nozzle. In order to prevent the oxidation and contamination of titanium alloy during the forming process, high-purity argon was employed in the powder carrier gas and inert atmosphere processing chamber. The working schematic diagram of LSF and the LSF-ed sample were shown in Fig. 2. The technological parameters of LSF used in this experiment were shown in Table 2. These parameters in Table 2 were determined by determined based on some references (Ref 18, 20, 21). By the process parameters, the 709 259 33 mm bulk sample was prepared on the TC17 titanium alloy forging with the size of 1109 309 6 mm, as shown in Fig. 2(b). After complete preparation of the sample, the step sample was cut along the laser scanning direction and perpendicular to the laser scanning direction to process the metallographic sample. After inlaying, polishing and etching, the macromorphology and grain size characteristics of the sample were observed by the optical microscope (OM) and scanning electron microscope (SEM). The measurement of grain size is carried out through the linear intercept technique from OM images. The test method was set as loading 500g and holding time 15s for Vickers hardness" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002953_tia.2021.3089662-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002953_tia.2021.3089662-Figure6-1.png", + "caption": "Fig. 6: Example of an on-load field solution obtained using RFO Analysis. The rotor currents are computed outside the FE Analysis and imposed along the q-axis of the adopted reference frame. Notice that the rotor flux linkage space vector lies along the d\u03bb-axis as required by the RFO condition.", + "texts": [ + " This formulation is described in [21]. However, this approach becomes more complicated when the machine is skewed, where it is necessary to consider several 2D slices and add further constraints to the FE problem equations [22]. The analysis technique adopted in this work is described in [19], [23]. The machine is simulated using the magnetostatic formulation , where both stator and rotor currents are field sources. An example of an on-load field solution obtained using the RFO Analysis technique is shown in Fig. 6. In a given time instant, the stator current is imposed in the chosen reference frame, named d\u03bbq\u03bb in Fig. 6, with the components isd and isq. The rotor currents are computed using the procedures described in [19], [23] and then imposed along the q-axis of the reference frame. The rotor currents are computed so that the q-axis rotor flux linkage results equal to zero. Therefore, the reference frame adopted in the simulations is exactly the d\u03bbq\u03bb, oriented with the rotor flux space vector, and all the vector quantities are referred to this reference frame. From the field solution in Fig. 6, it is possible to derive the stator and rotor flux linkage space vector in d\u03bbq\u03bb, \u03bb\u03bbs and \u03bb\u03bbr . The torque and the rotor angular frequency are derived using the model [24] as: Tdq = 3 2 p (\u03bbsdisq \u2212 \u03bbsqisd) ; \u03c9sl = p PJr Tdq (13) This analysis technique allows to consider the presence of closed rotor slot and the rotor skewing easily. For this purpose, the procedure outlined in [25] has been followed. Various 2D FE slices are considered and each field problem is solved using the static formulation, saving a long computation time" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001395_detc2008-49180-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001395_detc2008-49180-Figure4-1.png", + "caption": "Figure 4. A geared five-bar linkage", + "texts": [ + " Thus, Figure 2 shows that the linkage has three branches represented by segments 1-2, 3-4, and 5-6. Any part of the I/O curve lying outside of the JRS represents the four-bar mobility blocked by the five-bar loop. Figure 3 shows two separate I/O curves representing two branches of the Class I four-bar loop [5, 6], in which one branch is completely blocked by the five-bar loop. The remaining four-bar branch, which lies on both sides of the five-bar JRS sheet, constitutes two branches of the six-bar linkage. 2 Copyright \u00a9 2008 by ASME : http://www.asme.org/about-asme/terms-of-use Downl In Figure 4, the four-bar loop of a Stephenson six-bar linkage is replaced by a gear train and a geared five-bar linkage is formed. The geared train provides a linear I/O curve relating and (Figure 5), which intersects the five-bar JRS boundary at the branch points 1, 2, 3, and 4. Segments 1-2 and 3-4 within the five-bar JRS represent two branches of the five-bar linkage. The branch identification can be carried out in the way similar to that of Stephenson linkages [1, 10]. There should be no dead center position in a linkage branch with full rotatability", + " The resulting dead center positions with 8 and 7 as the input are listed in Tables 3 and 2 respectively and their corresponding points on the 3 vs. 2 curve are shown in Figure 8. One must note the corresponding angular transform relationship between these two inversions of Stephenson linkages, which are indicated in Tables 1, 2, and 3. The superscripts 1 and 6 on the angular displacements represent the corresponding values in Figure 1 and Figure 6 respectively with the same dimensions. A geared five-bar linkage (Figure 4) can also be considered as a multiloop linkage consisting of a five-bar loop and a gear loop. With an imposed gear train or a linear constraint, the five-bar linkage becomes a single degree of freedom mechanism. The following general gearing relationship [6] can be used to express the gear constrain ))(1()( 5054054 n (14) ded From: https://proceedings.asmedigitalcollection.asme.org on 12/15/2018 Terms of Use: where 4 and 5 are the joint variables at joints D and E, 40 and 50 are the joint displacements of the reference gear position, and n is the gear ratio. With the angle definition in Figure 4, the gear ratio is negative if both terminal gears rotate in the same direction with respect to the link DE. Equation (14) can be expressed as 54 n (15) where 40 if 050 . The calculation of dead center singularity of geared five-bar linkage can be carried out with Maple in the way similar to that of Stephenson linkage. In the first category, the input is given through the geared train (joint A or E). The dead center positions are the branch points, which are the intersection points between the linear I/O curve, i.e. Equation (15), and the JRS boundary. This case has been well treated [9] and shown in Figure 5, in which the branch points 1, 2, 3, and 4 are the dead center positions. In the other two categories, the input is given through a joint, such as A (or C) and B (Figure 4), not in the geared train. Since the choice of the output joint does not affect the dead center positions, to offer a unified treatment for having 4 or 5 as the input, consider the relationship between 4 and 5 . The loop closure equation of the five-bar loop ABCDEA of this linkage can be expressed as: )( 2 )( 3 )( 451 1543545 iiii eaeaeaeaa (16) The relationship among all angles can be written as 254321 (17) For the input given through joints A, B, and C, let 5 be considered as the output, which does not affect generality of the discussion on the existence of the dead center positions", + " That is 0)])1(sin( ))1sin((sin[ )])1(cos( ))1cos((cos[ 2 2 2 533 5455 2 533 54551 ana naa na naaa (25) The dead-center positions occur when the discriminate of Equation (18) is equal to zero or 0 )( )( )( 335 335 3 Q P (26) ded From: https://proceedings.asmedigitalcollection.asme.org on 12/15/2018 Terms of Use Thus, the dead center positions can be obtained by solving 0)( 335 P and 0)( 335 Q . It should be noted that, for ease of calculation, Equations (19), (23) and (25) must be rewritten in the half-tangent-angle formula. Example 2: The dimension of a geared five-bar linkage (Figure 4) is given as 345.21 a , 665.02 a , 810.13 a , 235.14 a , 661.15 a , 2/1n , 325.17 . The branch points or dead center positions when the input is given through 5 are listed in Table 4 and shown in Figure 5. The dead center positions with the input given through 3 are listed in Table 5 and shown in Figure 11. There is no dead center position when the input is given 1 or 2 (Figure 10 and Figure 12). The full rotatability of a linkage is input related while the branch is irrelevant to the input. The full rotatability of a linkage should refer to a specific branch", + " The proposed method is illustrated with Stephenson six-bar linkage and geared five-bar linkage. For Stephenson six-bar linkage and geared five-bar linkage, the branches can be 6 Copyright \u00a9 2008 by ASME : http://www.asme.org/about-asme/terms-of-use Downlo identified with the JRS method [1, 10] discussed briefly above. The following discussion on full rotatability takes the effect of using different input joints into consideration. 4.1 Input Given to a Link of the Four-bar Loop or Geared Train: The full rotatability of geared five-bar linkage (Figure 4) with the input given in the geared train has been discussed by Ting [2]. The linear input-output (I/O) constraint must lie within the JRS of the five-bar loop and a unit geared ratio is required [2] unless the host five-bar contains no uncertainty singularity (or dead center positions). For a branch of a Stephenson linkage (Figure 1) to have full rotatability, the following conditions must be satisfied. 1. The four-bar loop is a Class I chain and the input joint must connect the short link of the four-bar loop" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003345_ur52253.2021.9494692-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003345_ur52253.2021.9494692-Figure1-1.png", + "caption": "Figure 1. 5 DOF robot manipulator with 1 DOF prismatic joint.", + "texts": [ + " Khan is PhD student in Department of Mechanical Engineering, Pusan National University, Busandaehak-ro 63beon-gil, Jangjeon 2(i)-dong, Geumjeong-gu, Busan, South Korea, (e-mail: hamzakhan.0496@gmail.com). 978-1-6654-3899-5/21/$31.00 \u00a92021 IEEE 375 20 21 1 8t h In te rn at io na l C on fe re nc e on U bi qu ito us R ob ot s ( U R ) | 9 78 -1 -6 65 4- 38 99 -5 /2 1/ $3 1. 00 \u00a9 20 21 IE EE | D O I: 10 .1 10 9/ U R 52 25 3. 20 21 Authorized licensed use limited to: CALIFORNIA INSTITUTE OF TECHNOLOGY. Downloaded on August 13,2021 at 12:40:07 UTC from IEEE Xplore. Restrictions apply. includes; XY-crane, 1 DOF prismatic joint, 5 DOF robot manipulator (Fig. 1), and the circular RVI environment has been designed in the SolidWorks. In the second step, this CAD file has been converted into MATLAB/Sim Mechanics where the robot theories have been implemented to control this virtual simulator. In this section, the proposed control scheme has been discussed in detail. Nonlinear ESO has briefly explained in the first part, 2nd part describes the ISMC and the 3rd explained the integration of ISMC with ESO. Consider a 2nd order system; ?\u0307?1 = \ud835\udc652 (1) ?\u0307?2 = \ud835\udc62 \u2212 \ud835\udc53(\ud835\udc65, " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003132_978-981-16-0119-4_20-Figure78.5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003132_978-981-16-0119-4_20-Figure78.5-1.png", + "caption": "Fig. 78.5 Conceptual design of propeller with curved cut on both Les", + "texts": [], + "surrounding_texts": [ + "d. Discretization The unstructured grids are employed in this discretization process because the indented numerical technique for this CFD tool is Finite Volume Method (FVM). Computationally, the FVM technique is fit for both structural and unstructural grids, so in this work, purposively unstructural grids are employed, and thereby, centroidbased FVM is implemented in the computation stage. The fine proximity and fine curvature sizing based mesh facilities are implemented in this discretization phase [21, 22,and23].The typicalmeshed structure onbase propeller is revealed inFig. 78.6, in which the quality of mesh is attained as 0.92, and very few elements are formed as aspect ratio of 1.2. In addition to this base process, the grid convergence study has been organized on the BOTH LE Propeller for the given velocity of 10 m/s, and the comprehensive data are revealed in Fig. 78.7." + ] + }, + { + "image_filename": "designv11_83_0000921_1.2833313-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000921_1.2833313-Figure1-1.png", + "caption": "FIG. 1. Mesh of SPM motor with overhanging magnet 1 /8 model . a Homogenization method elements: 143 868 . b Detailed modeling elements: 371 028 .", + "texts": [ + " In order to clarify the effectiveness of the homogenization method, an iron-loss analysis of a SPM motor is carried out. The specification shown in Table I is nearly the same as that of the motor mentioned in Ref. 7 except for the dimensions of the magnet and the rotor. The overhanging magnet generates the perpendicular magnetic flux, which results in the large eddy-current loss at the stator core end. The core is laminated with 80 steel sheets whose thickness is 0.5 mm. An eighth part of the whole model is analyzed because of the symmetry and periodicity. Figure 1 shows the meshes. In the case of the homogenization method, the mesh is not restricted by the laminated structures. On the other hand, in the case of the detailed modeling, one sheet is divided into one layer of hexahedral elements along the laminated direction, and gap elements and double nodes are adopted between steel sheets. To stabilize the convergence characteristic of the nonlinear iterations, the functional NR is applied.8 Figure 2 shows the distribution of eddy-current loss density at 1500 min\u22121" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002798_acs.jpcb.1c01116-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002798_acs.jpcb.1c01116-Figure1-1.png", + "caption": "Figure 1. Three geometries: infinite planar, cylindrical, and spherical bilayers with a given curvature.", + "texts": [ + "24 Numerically, the additional orientation dimensions of wormlike chains make the computation of liquidcrystalline bilayers time-consuming. This computational burden is aggravated when PFM is used to extract the elastic constants because the PFM needs to calculate the free energy of bilayers with many different curvatures. In this study, using the AEM, we analyze the excess free energy of bilayer membranes in three different geometries: an infinite planar bilayer; a cylindrical bilayer, which is extended to infinity in the axial direction; and a spherical bilayer with different curvatures (Figure 1). Specifically, we treat the curvature of cylindrical and spherical bilayers as small parameters and then carry out asymptotic expansions for the modified diffusion equation of propagators and order parameter of polymeric monomers in terms of the curvature, after which the self-consistent field (SCF) equations at each order can be derived. Finally, we can obtain analytical expressions of the free energy at each order, which are related to the elastic moduli and could be computed separately. Notably, a few SCF equations need to be solved to calculate these expanded free energies" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001818_acc.2007.4282978-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001818_acc.2007.4282978-Figure3-1.png", + "caption": "Figure 3 Quanser UFO Figure 4 Earth and body frames of UFO", + "texts": [ + " Careful selection of the time scale and closed loop PD-eigenvalues% ________________________________________________ *When the PD-spectrum is chosen as the product of constant nominal eigenvalues and a time-varying bandwidth , the synthesis formula _ 3 =5 \u00d0>\u00d1 becomes much simpler are crucial to solve the peaking issue in practical control problems as well. More specifically, avoiding too small value of and complex and repeated values of eigenvalues% should attenuate the peaking phenomenon. The UFO is a vertical takeoff and landing (VTOL) aircraft model made by Quanser Consulting, Inc . It has 3 pro-[9] pellers driven by DC motors, which are mounted on a triangular frame, as shown in Figure 3. The UFO has 3 degreesof-freedom (DOF), which are represented by the Euler pitch angle , roll angle , and yaw angle , as shown in Figure 4.) 9 < The modeling and controller design of UFO are presented in [15]. In that paper, pseudo differentiators have been used to estimate body rates of the UFO from Euler angles measured from optical encoders with the kinematic equation of motion. However, pseudo differentiator estimator (PDE) lacks generality and uses information of only part of the dynamics. Most importantly, the estimation errors of PDE are not regulated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002945_978-3-030-69178-3_15-Figure17.4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002945_978-3-030-69178-3_15-Figure17.4-1.png", + "caption": "Fig. 17.4 Assembly of aircraft wing rib", + "texts": [ + " The less relevant and failure-prone tasks are left to the robots. The customers of the new system have more flexibility and independence of labour. Especially in the case of high-volume production, the system can assist the operators properly. The system decreases the number of failures in mixing the products, which occur today very often. Demonstrator #2 was located at Aciturri (ACI) in Spain that manufactures wing ribs made of carbon fibre for commercial aircraft. The selected assembly process is the assembly of wing ribs as shown in Fig. 17.4. An assembly of a wing rib is a substructure that is mainly composed of a Carbon Fibre Reinforced Polymer (CFRP) rib. Several Titanium \u201cT\u201d Clips called TIE are required to be riveted to the rib perimeter by a special tool called CLECOS. The industrial issues encountered are listed as follows: \u2022 Time-consuming manual wing rib assembly process increase costs. Currently, humanworkers completemost of thewing rib assembly operations, such as setting up jigs, positioning and drilling elements accurately on the jig, and checking key quality characteristics" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002881_s42235-021-0043-x-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002881_s42235-021-0043-x-Figure4-1.png", + "caption": "Fig. 4 Components of one wing.", + "texts": [ + " When the slider starts moving from a static state, the wings deflect under the resistance and inertia of the wind, causing the angle of attack to vary. When the limit structures of the connector and sliders are in contact, the wing angle of attack will no longer vary and the wings will move horizontally under the set angle of attack to generate lift. When the direction of the oscillating motion changes, the wing angle of attack will reverse until it is restricted again. The oscillating wings should be sufficiently stiff to sustain dynamic wing loading and as light as possible. Table 1 and Fig. 4 show the design and manufacture of the wings. The design of the wing shape has a wide range of possibilities[38]. The rectangular design of the wing is to simplify the design by reducing parameters in the initial design of the prototype. Like those of most existing MAVs, the wings are composed of lightweight but high-strength materials, namely, carbon fiber rod for the leading edge and wing supporting veins and Journal of Bionic Engineering (2021) Vol.18 No.3 652 polyester film for the membrane" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002995_iciccs51141.2021.9432294-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002995_iciccs51141.2021.9432294-Figure1-1.png", + "caption": "Fig. 1. Testbed model of Quanser AERO.", + "texts": [ + " The rest of the paper is organized as: For the problem formulation given in section II, the Controller is designed in Section III. Section IV discusses simulation results followed by conclusions in section V and references. Quanser AERO consists of two rotors viz. front rotor and a back rotor coupled with two high-efficiency dc motor [4].The front rotor predominately affects the motion about the pitch axis while the back rotor mainly affects the motion about the yaw axis. The Quanser AERO test-bed model is shown in Fig.1 The free-body representation of the Quanser AERO is shown in Fig.2. When an input voltage Vp is applied to the front rotor, the speed of rotation causes a force Fp at a distance rp from the Y -axis. Simillarly, yaw motor voltage, Vy genrates a force Fy at a distance of ry from the Y -axis. Both the rotors of Quanser AERO having the same size and equidistant from each other [4], due to this front rotor affects 978-0-7381-1327-2/21/$31.00 \u00a92021 IEEE 1595 20 21 5 th In te rn at io na l C on fe re nc e on In te lli ge nt C om pu tin g an d Co nt ro l S ys te m s ( IC IC CS ) | 9 78 -1 -6 65 4- 12 72 -8 /2 1/ $3 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002979_iemdc47953.2021.9449609-Figure11-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002979_iemdc47953.2021.9449609-Figure11-1.png", + "caption": "Fig. 11 Cross-sections of optimized 24-slot/20-pole (a) modular and (b) nonmodular machines. In (a), the flux gap width is 0.5mm.", + "texts": [ + " This is mainly because the flux gaps do not only have effect on open-circuit airgap flux density, but also have effect on pitch factor (or winding factor) and flux focusing. These different effects lead to different optimal flux gap widths for maximum airgap flux density and average torque. F. Optimization of 24-slot/20-pole Modular and Non-modular Machine For comparison purpose, the same optimization process proposed in section III. A is implemented for the 24-slot/20pole modular and non-modular machines. The optimized machine parameters are listed in Table VI and the crosssections are shown in Fig. 11. It is worth noting that considering the feasible mesh size in the flux gap region, the minimum flux gap width chosen for the optimization process is 0.5mm, which is exactly the same as the optimized flux gap width for the 24- slot/20-pole modular machine. The results have shown that by introducing the flux gaps, the fitness function will always reduce regardless of the flux gap width, this is the same for the primary objective, i.e. average torque. Therefore, this means that for the 24-slot/20-pole machine, the flux gaps should not be introduced to avoid deteriorating its performance" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002360_012005-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002360_012005-Figure3-1.png", + "caption": "Figure 3. Fishtail closed turn with rectilinear reverse movement, which is not parallel to the field boundary - movement from left to right.", + "texts": [ + " The purpose of the present work is to derive analytical dependences for determining the length of fishtail turns with a rectilinear reverse move that is not parallel to the field boundary, as well as for determining the width of the headland required to perform them in an irregularly shaped field and to analyze the influence of the angle between the direction of movement and the field boundary on the length of the non-working move and the width of the headland at different direction of movement of the unit in the field. A fishtail open turn and a fishtail closed turn with a rectilinear reverse movement in different directions of their performance in a field with an irregular shape are considered. To determine the length of the turn, a geometric method is used, in which the turn is represented by straight lines and arcs of a circle of equal radius. The different types of turns are presented in Figure 1, Figure 2, Figure 3 and Figure 4. The length of the turn is defined as the sum of the lengths of its geometric elements. The width of the headland required to perform the turn is defined as the sum of the segments perpendicular to the field boundary and depending on the elements of the turn. The symbols used in the figures are as follows: \u03b1 is the angle between the direction of working move of the unit and the boundary of the field; p. A \u2013 the beginning of the turn; p. B \u2013 the end of turn; p. O1, p. O2 \u2013 the centers of the first and second curvilinear movement within the turn; \u03b21, \u03b22 \u2013 the central angles of the arcs described in the first and second curvilinear movements; R \u2013 the radius of curvilinear movement (turning radius of the machine-tractor unit); M \u2013 the width of the tractor measured from the outside of the wheels", + " sin \ud835\udefc (23) After entering the headland with width \ud835\udc38\u2032 and before leaving it, the unit makes a rectilinear move with length \ud835\udc59\ud835\udc54 = (0,5\ud835\udc40\u2212\ud835\udc45) cos(\ud835\udefd2\u2212\ud835\udefc) sin \ud835\udefc + (\ud835\udc45\u22120,5\ud835\udc35) tan \ud835\udefc + \ud835\udc59\ud835\udc4e = (0,5\ud835\udc40\u2212\ud835\udc45) cos \ud835\udefd2+\ud835\udc45\u22120,5\ud835\udc35 tan \ud835\udefc + (0,5\ud835\udc40 \u2212 \ud835\udc45) sin \ud835\udefd2 + \ud835\udc59\ud835\udc4e (24) The rectilinear move becomes \ud835\udc59\ud835\udc54 = 0 at \ud835\udefc = tan\u22121 ( (0,5\ud835\udc40\u2212\ud835\udc45) cos \ud835\udefd2+\ud835\udc45\u22120,5\ud835\udc35 (\ud835\udc45\u22120,5\ud835\udc40) sin \ud835\udefd2\u2212\ud835\udc59\ud835\udc4e ) (25) After reaching the boundary of the headland, the unit turns right, moving along a curve reaches the boundary of the field, makes a rectilinear reverse move and again through a curvilinear forward move to the right reaches the boundary of the headland and begins the next working move (Figure 3). When the unit is reversing, the tractor wheels goes outside the headland with width E. Since this part of the field is cultivated and the wheels enter the cultivated field, it is better to determine the width of the headland not according to the capabilities of the unit to perform a turn without a straight move forward in the headland with width E, but by take into account the innermost point in the field reached by the tractor wheels. In this case, the width of the headland is larger and is marked with \ud835\udc38\u2032", + " Since the sum of the central angles is \u03b21 + \u03b21 = \u03c0, the length of the turn is determined by the dependence \ud835\udc59\ud835\udc47 = \ud835\udf0b\ud835\udc45 + \ud835\udc421\ud835\udc422 \u0305\u0305 \u0305\u0305 \u0305\u0305 \u0305 = \ud835\udf0b\ud835\udc45 + 2\ud835\udc45\u2212\ud835\udc35 sin \ud835\udefd1 (26) The central angle \u03b21 is determined by the triangle O1CO2 by the dependence \ud835\udefd1 = tan\u22121 ( 2\ud835\udc45\u2212\ud835\udc35 \ud835\udc35 tan \ud835\udefc +2\ud835\udc59\ud835\udc4e ) (27) Finally for the length of the turn is obtained \ud835\udc59\ud835\udc47 = \ud835\udf0b\ud835\udc45 + [(2\ud835\udc45 \u2212 \ud835\udc35)2 + ( \ud835\udc35 tan \ud835\udefc + 2\ud835\udc59\ud835\udc4e) 2 ] 1 2\u2044 (28) The width of the headland for a small angle between the direction of movement and the field boundary is the sum of segments of the following lengths (Figure 3, (a)): \ud835\udc4e = 0,5\ud835\udc40. cos(\ud835\udefc+\ud835\udefd1) (29) \ud835\udc4f = \ud835\udc3b. sin(\ud835\udefc + \ud835\udefd1) (30) \ud835\udc52 = \ud835\udc59\ud835\udc4e . sin \ud835\udefc + (0,5\ud835\udc35 + \ud835\udc45) cos \ud835\udefc \u2212 \ud835\udc45. cos(\ud835\udefc+\ud835\udefd1) (31) ICTTE 2020 IOP Conf. Series: Materials Science and Engineering 1031 (2021) 012005 IOP Publishing doi:10.1088/1757-899X/1031/1/012005 For the width of the headland, where there is no rectilinear move (lg = 0) is obtained \ud835\udc38 = \ud835\udc4e + \ud835\udc4f + \ud835\udc52 = (0,5\ud835\udc40 \u2212 \ud835\udc45). cos(\ud835\udefc+\ud835\udefd1) + \ud835\udc3b. sin(\ud835\udefc+\ud835\udefd1) + \ud835\udc59\ud835\udc4e sin \ud835\udefc + (0,5\ud835\udc35 + \ud835\udc45) cos \ud835\udefc (32) The width of the headland, taking into account the entry of the tractor into the field during its reverse movement, is determined by the dependences (30), (31) and \ud835\udc50 = (2\ud835\udc45\u2212\ud835\udc35).sin(\ud835\udefc+\ud835\udefd1) sin \ud835\udefd1 (33) \ud835\udc38\u2032 = 2\ud835\udc4e + \ud835\udc4f + \ud835\udc50 = \ud835\udc40. cos(\ud835\udefc+\ud835\udefd1) + \ud835\udc3b. sin(\ud835\udefc+\ud835\udefd1) + (2\ud835\udc45\u2212\ud835\udc35).sin(\ud835\udefc+\ud835\udefd1) sin \ud835\udefd1 (34) ICTTE 2020 IOP Conf. Series: Materials Science and Engineering 1031 (2021) 012005 IOP Publishing doi:10.1088/1757-899X/1031/1/012005 The length of the rectilinear moves of the unit before and after the turn in the headland with width \ud835\udc38\u2032 is \ud835\udc59\ud835\udc54 = (2\ud835\udc45\u2212\ud835\udc35).sin(\ud835\udefc+\ud835\udefd1) sin \ud835\udefc.sin \ud835\udefd1 + (0,5\ud835\udc40+\ud835\udc45).cos(\ud835\udefc+\ud835\udefd1) sin \ud835\udefc \u2212 \ud835\udc59\ud835\udc4e \u2212 (0,5\ud835\udc35 + \ud835\udc45) cot \ud835\udefc (35) From Figure 3(a) it can be seen that when performing a rectilinear reverse, the left rear wheel of the tractor reaches the boundary of the headland. In increasing the angle \u03b1 between the direction of movement and the field boundary, the position of the tractor changes and at some point its rear axle becomes parallel to the boundary of the headland. Then the length of the segment a becomes equal to zero. From dependence (29) it can be seen that this happens at cos(\ud835\udefc+\ud835\udefd1) = 0. Then \ud835\udefd1 = 90 \u2212 \ud835\udefc. After substitution in equation (27) is determines the angle at which the rear axle of the tractor is parallel to the field boundary \ud835\udefc = tan\u22121 ( \ud835\udc59\ud835\udc4e+[\ud835\udc59\ud835\udc4e 2+(2\ud835\udc45\u2212\ud835\udc35)\ud835\udc35] 1 2\u2044 2\ud835\udc45\u2212\ud835\udc35 ) (36) When the angle \u03b1 increases above this value, the headland is limited by the rear right wheel of the tractor, as it is located more inward in the field (Figure 3, (b)). The width of the headland in this case is defined as the sum of segments b (equation (30)), e (equation (31)), as well as the segments: \ud835\udc4e = \u22120,5\ud835\udc40. cos(\ud835\udefc+\ud835\udefd1) (37) \ud835\udc51 = (2\ud835\udc45\u2212\ud835\udc35).sin(\ud835\udefc+\ud835\udefd1) sin \ud835\udefd1 \u2212 0,5\ud835\udc40. cos(\ud835\udefc+\ud835\udefd1) (38) Hence, for the two variants of determining the width of the headland for Figure 3, (b) is obtained: \ud835\udc38 = \ud835\udc4e + \ud835\udc4f + \ud835\udc52 = = \ud835\udc3b. sin(\ud835\udefc+\ud835\udefd1) + \ud835\udc59\ud835\udc4e sin \ud835\udefc + (0,5\ud835\udc35 + \ud835\udc45) cos \ud835\udefc \u2212 (0,5\ud835\udc40 + \ud835\udc45). cos(\ud835\udefc+\ud835\udefd1) (39) \ud835\udc38\u2032 = \ud835\udc4e + \ud835\udc4f + \ud835\udc51 = \ud835\udc3b. sin(\ud835\udefc+\ud835\udefd1) + (2\ud835\udc45\u2212\ud835\udc35).sin(\ud835\udefc+\ud835\udefd1) sin \ud835\udefd1 \u2212 \ud835\udc40. cos(\ud835\udefc+\ud835\udefd1) (40) The two variants of the headland are equalized when e = d, i.e. when there is no rectilinear move before and after the turn in the headland. The length of this move is determined by the dependence \ud835\udc59\ud835\udc54 = (2\ud835\udc45\u2212\ud835\udc35).sin(\ud835\udefc+\ud835\udefd1) sin \ud835\udefc.sin \ud835\udefd1 + (\ud835\udc45\u22120,5\ud835\udc40).cos(\ud835\udefc+\ud835\udefd1) sin \ud835\udefc \u2212 \ud835\udc59\ud835\udc4e \u2212 (0,5\ud835\udc35 + \ud835\udc45) cot \ud835\udefc (41) This dependence can also be represented in the form \ud835\udc59\ud835\udc54 = (2\ud835\udc45\u2212\ud835\udc35) tan \ud835\udefc + (\ud835\udc45\u22120,5\ud835\udc40)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001746_memsys.2007.4433151-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001746_memsys.2007.4433151-Figure4-1.png", + "caption": "Fig. 4. (a) Structure of a SWNT thin film based ACh biosensor. (b) An optical image of a fabricated SWNT biosensor next to a dime.", + "texts": [ + " The sequence of the assembly process is: [PDDA + PSS]2 + [PDDA + SWNT]5 + PDDA + [PSS + AChE]3. The adsorption time for both PDDA and PSS is 10 min. The scanning electron microscope (SEM) images of the assembled SWNT and AChE films are shown in Fig. 3a and 3b, respectively. The SWNTs are interconnected and form a dense network. Both films are relatively uniform due to the presence of the intermediate \u201cannealing\u201d polyelectrolyte layers. The structure of the SWNT thin film ACh sensor is shown in Fig. 4a. The sensor is a resistor-like device with two terminals. The channel between the electrodes is covered by ACh sensing and SWNT conducting layers. Fig. 4b shows an optical image of a fabricated device. The size of the device is 1 cm \u00d7 1 cm, which is smaller than a dime. The electrodes are elongated to the side of the device for probing and detection. The fabrication process of the SWNT thin film sensor is shown in Fig. 5. Cr (100 nm) and Au (200 nm) are evaporated on the Si/SiO2 wafer as adhesion and electrode pad materials, respectively. A layer of positive photoresist (PR1813) is spin coated on the surface and patterned using lithography. The Cr/Au layers are patterned as the electrode pads using wet etching technique" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002783_09544062211012724-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002783_09544062211012724-Figure8-1.png", + "caption": "Figure 8. Singular case where axes of three limbs are parallel.", + "texts": [], + "surrounding_texts": [ + "An example of the 3-RPR PM is used to validate the proposed accuracy analysis method, as shown in Figure 4. A force pair W \u00bc \u00bdF;M > \u00bc \u00bd300N; 100N; 400Nmm > is employed at the center of the manipulator in the global system. The initial coordinates of the connection points at the manipulator(A1;A2;A3) and at the base(B1;B2;B3) are 200cos 3p 2 np 3 200sin 3p 2 np 3 0 BBB@ 1 CCCA and 400cos 4p 3 2np 3 400sin 2p 2np 3 0 BBB@ 1 CCCA respectively, where n\u00bc 1, 2, 3. The global system O \u2013 xy and the body system O1 x1y1 coincide initially. Then the initial pose g0 \u00bc I3 is the 3 3 identity matrix. The rest of the parameters essential to calculation are presented in Table 1. In Table 1, r is magnitude of the nominal clearance, E is elastic modulus of the link, A is section area of the link, k is stiffness coefficient of the locked actuator, a1 and a2 are the weighting factors. In order to verify the accuracy of the numerical computation, an observation Sr indicating the balance of the elastic force and the applied force is used. The Sr is define as Sr \u00bck We \u00feW k (28) where We and W are the elastic force and the applied force, respectively. Obviously, observation Sr approaches 0 when the two forces are balanced. We search for the error pose in the neighborhood of the error-free one, and it is not likely to have more than one neighboring poses that satisfies the force constraint. Therefore, it is safe to say that the close-totarget pose making Sr close to 0 is exact the error pose. The errors of the 3-RPR mechanism with multiple joint clearances under the constant load \u00bd300N; 100N; 400Nmm > are analyzed using the proposed smooth function method and the common step function method, respectively. The equation (18) is solved numerically using the Trust-region method with the initial guess \u00bd0:3; 0; 0 . In a laptop with 2.70GHz CUP and 4.0 GB RAM, the results are obtained, as shown in Table 2. Besides, the error pose of the manipulator is computed as ge \u00bc 1 4:14e 4 0:22 4:14e 4 1 0:16 0 0 1 0 @ 1 A (29) which is close to the target pose g0 \u00bc I4, where I4 is a 4 4 identity matrix. It can be observed from Table 2 that the use of tanh function helps the algorithm to converge and it computes about 10 times faster than using the step function. The target pose of the manipulator g0 is I3. When the load \u00bd300N; 100N; 400Nmm is employed, the elastic force and corresponding deformation of each limb are obtained by equations (25) and (26), respectively, as \u00bd45:5N; 200:1N; 156:9N > and \u00bd 0:0011mm; 0:0043mm; 0:0034mm >. When the clearances are considered, the final strokes of the actuators should be \u00bd0:2011mm; 0:2043mm; 0:2034mm > by equation (27) to fully compensate the errors arisen from clearances and deformation. It can be concluded by comparing the deformation and clearances that the clearances have larger effect on the manipulator accuracy. The employed force can help to eliminate the uncertainty of clearances and therefore benefit the calibration and error compensation. In order to further prove that the proposed method computes faster, 6 more forces are generated and applied to the manipulator. The generated forces and the computation results are presented respectively in Table 3 and Table 4. In Table 4, the first values are the computation results obtained using the proposed tanh representa- tion method and the second values are the results obtained by the step function method. By comparing the results in Table 4, the efficient and accuracy of the proposed method are clearly verified. It is noteworthy that there are many expressions that can substitute the Heaviside function for fast computation. Figure 5 give the comparison of several alternatives. In Figure 5, modified sigmoid function, arctangent function and one of the algebraic functions are respectively f\u00f0x\u00de \u00bc 2 1\u00fe e x 1 (30) f\u00f0x\u00de \u00bc arctan\u00f0x\u00de (31) and f\u00f0x\u00de \u00bc xffiffiffiffiffiffiffiffiffiffiffiffiffi 1\u00fe x2 p (32) It can been observed in Figure 5 that the tanh\u00f0x\u00de function is a shifted and narrowed sigmoid function and tanh\u00f0x\u00de is scaled in [\u20131]. The tanh\u00f0x\u00de is closer to the Heaviside function compared to the others. Singularity analysis with loads A PM loses degree of freedoms at its singular configuration. For the prismatrically actuated 3-RPR mechanism, it has three singular types, as shown in Figures 6 to 9. The singularity depicted in Figure 6 occurs in the limit situation where at least one limb of the 3-RPR PM owns zero limb length. This theoretical situation is easy to avoid in practice. Then the effect of employed load on this situation is not further discussed here. The planar 3-RPR manipulator lose its rotational DoF when three axes of limbs intersect at the common point O1, as shown in Figure 7. Two necessary conditions need to be met to have this kind of singularity when employed loads are considered, that is, 1. Axes of the three limbs can intersect at the common point geometrically within the workspace of the manipulator. 2. No torsional force is applied to the manipulator. It is clear that employed torque can eliminate this singularity. Assuming that the two necessaries conditions are met and a linear force F is employed. The singular configuration of the 3-RPR mechanism can be obtained by the error compensation method discussed in subsection 2.3. Since the directions of the axes \u00bds1; s2; s3 are linear correlated, the H in equation (25) is not invertible. The elastic force \u00bdF01;F02;F03 > should be computed by s1 \u00fe s2 \u00bc w1s3 F01s1 \u00fe F02s2 \u00fe F03s3 \u00bc F (33) where w1 is factor decided by the structure of the mechanism. The limb stroke \u00bdd01; d02; d03 > can be then computed by equations (26) and (27). Then the singularity occurs at the limb length Li \u00bck g0 yi xi k d0i i \u00bc 1; 2; 3 (34) where g0 denotes the singular pose. Figures 8 and 9 illustrate the singular cases where the axes of the limbs are parallel. The manipulator lose the DoF to translate in the direction perpendicular to the axes. These two singular case can only occur on the premise that two necessary conditions are met, that is, 1. Three limbs can move to the parallel positions within the workspace. 2. When linear force is employed, the direction of the force must be parallel to the axes of the limbs. Then, this kind of singularity can be eliminated by employing proper liner force to the manipulator. Given that the two conditions are met and the load is \u00bdF;M >, the elastic force of each limb can be computed by s1 \u00bc w2s2 \u00bc w3s3 F01s1 \u00fe F02s2 \u00fe F03s3 \u00bc F g0 y1 s1 \u00fe g0 y2 s2 \u00fe g0 y3 s3 \u00bc M 8< : (35) where both the factors w2 and w3 are 1 or \u20131, and F is parallel to si. Then the configurations of the 3-RPR mechanism where singularities shown in Figures 7 and 9 may occur can be obtained by equations (26), (27) and (34). To conclude, singularities occur under specific combinations of loads and configurations. When the load is properly exerted to the manipulator, the singularities of the 3-RPR manipulator can be avoided." + ] + }, + { + "image_filename": "designv11_83_0002135_s00202-020-01163-8-Figure11-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002135_s00202-020-01163-8-Figure11-1.png", + "caption": "Fig. 11 The middle rotor of the proposed TSCMG", + "texts": [ + " Therefore, we can infer that the gear structure Fig. 5 The structure of the proposed TSCMG 1 3 given in Fig.\u00a08(b) cannot be assumed as a proper structure for multi-speed gears. According to the analysis conducted in the previous section, a prototype of the proposed TSCMG was implemented to verify the simulation results. The gear consists of inner, middle and outer rotors, where the middle rotor is designed heuristically using a belt shaft so that speed and torque can be extracted from this rotor. Figure\u00a011 shows the output of the middle rotor belt. The important point to notice is the parameters of the middle rotor, which play a key role in the torque of the inner and outer rotors. That is why we can say that the middle rotor is the most critical and effective component of a multispeed magnetic gear, the magnetic field of which affects other rotors at the same time. Furthermore, Fig.\u00a012 illustrates the modulators and rotors of the proposed magnetic gear built in this work. An axial cut of the proposed TSCMG is graphed in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001692_sisy.2007.4342648-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001692_sisy.2007.4342648-Figure2-1.png", + "caption": "Fig. 2. The inverted pendulum with its foot and force sensors within", + "texts": [ + " For bipedal walk, the following can be concluded: dynamic balance of the biped mehanism is preserved as long as the ZMP is within the support polygon. If we knew all the forces in the system at any instance, we could determine if the system is in dynamic balance or not. But in systems having many degrees of freedom (DoF) it is quite hard to determine (and would require many expensive sensors to measure, or heavy computations to estimate using a mathematical model). Using the ZMP criterion would provide us with the same result and might be much simpler and less 1-4244-1443-1/07/$25.00 \u00a92007 IEEE 183 expensive to measure. In figure 2, the structure of the inverted pendulum is given. It has three components with their corresponding masses: m1 is the mass ot the flat foot, m2 is the mass of the vertical rod (leg) and finally a mass of m3 at the top represented by a sphere. The pendulum also has two DoFs represented by two cylindrical joints at the base of the leg, that allow rotation around axes ex and ey. Four force sensors are part of the foot. In figure 2, they are marked as fs1 to fs4 (the last one is covered by the leg). They are located near the corners of the rectangular foot as can be seen in figure 3. The inverted pendulum is not attached to the foot at its center, but the joint is slighlty moved towards the back edge of the foot to be more similar to a human ankle. The distances from this joint to the sensors are clearly marked (f ,b,l and r). For the sake of simplicity, the dotted line that connects the four sensors, as it is close to the edge of the foot, will be considered as the borderline of the support polygon", + " For the denominator sum to become zero, some of the measured forces should be negative. For structures fastened to the ground (robot manipulators) it is quite normal, but measuring a negative force in the case of the inverted pendulum with flat foot is a problem that will not be further discussed in this paper. In the following examples, all forces measured by the sensors will be positive. II. INVERTED PENDULUM WITH 4 SENSORS The follwing example anlyzes the motion of the inverted pendulum depicted in figure 2. The details of the foot can be seen in figure 3. Table I contains the mechanical parameters of the analyzed system. The system dynamics is accurately modeled and all the results that follow are results of simulations. [3] The structure of the simulation system is illustrated in figure 5. Its two main building blocks are the inverted pendulum model and the ZMP position reconstruction. The former contains an exact mechanical model of the pendulum. Its inputs are (in this case) the prescribed motion, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000436_intmag.2005.1464145-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000436_intmag.2005.1464145-Figure3-1.png", + "caption": "Fig 3 2D and 3D Torque simulation Fig 4 The nsing tune simulation and measured one", + "texts": [ + " Figure 2 shows the radial direction current component of each coil position, where x is the distance from coil turn left end and r is the radius from coil centers to coil ends oreach end. We will show more detail procedure of deriving the equation in the full paper. From those equations and the air-gap flux density by 2D FEM, we can get torque values at each point of coil. Comparison o f Analysis Result and 3D FEM To verify this method, the result is compared to the torque values from the 3D FEM result at scveral angles. In the fixed current condition, wc h a x calculated torque values from 0\" to 90\". Figure 3 shows values of two simulation results in which we know that the two results reasonably well match. Motion Transient Simulation For the motion transient simulation of the flat-type vibration motor, we have to know the load torque of the motor. In this study, some experiments have been done to obtain the load torque as a function of rotating speed. Interpolation of the experimental values gives the load torque curve expressed as ,where I is the moment of inertia of the rotor (1) 7;(~)=(0.00079673 uz +0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002295_j.robot.2020.103715-Figure9-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002295_j.robot.2020.103715-Figure9-1.png", + "caption": "Fig. 9. Overview of color coding of functional states of joints and segments. Coordinate frames used for the forward kinematics are shown for exemplary segments. Segments and joints were numbered counter-clockwise as depicted. For each segment, a local coordinate system is defined, whose origin coincides with the location of the corresponding joint. The x-axes point to the axes of the next joint in the sequence. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)", + "texts": [ + " Therefore, these segments as well as he connecting joints will be referred to as \u201clocked\u201d segments and oints. On the other hand, the angles of joints between segments ithout ground contact can be changed rather independently within the constraints that will be explained in Section 3.2). herefore, they will be referred to as \u201cfree\u201d segments and \u201cfree\u201d oints. During the rolling motion, segments and joints switch conecutively between these two states. An overview of the different tates of joints and segments is given in Fig. 9. The control of the robot can therefore be reduced to the conrol of the free joints, which consequentially defines the positions f the free segments, and the switch of the rearmost segments nd joints from locked to free state. The speed, by which the robot ocomotes, is controlled via the angular velocity of the foremost oint (relative to the sequence of locked segments). This joint owers the first free segment until it detects ground contact, upon hich joint and segment are switched to locked state and the ext joint starts to lower the following segment", + " If, however, a contact is detected at height that allows to duck down and traverse beneath the bstacle, the position of the contact is recorded to serve as virtual epresentation of the obstacle. After the sensor data acquisition, the planning of movements or the next iteration is initiated. For this purpose, within the virtual representation of the robot, the joint drive that connects the sequence of locked segments with the foremost free segment is explicitly commanded to lower this segment to continue the forward movement (see Fig. 9). The angular speed by which this segment is lowered, defines the overall movement speed of the robot. On the back side of the robot, correspondingly, the hindmost locked segment is tested for whether it can be freed by estimating the stability of the system considering this segment to be missing for support. If the stability can be maintained, the segment is unlocked. The corresponding joint angle can therefore be adjusted by the optimizer in the subsequent posture optimization. This optimization is based on multiple factors that will be described in the following subsection", + " Since the segments are mechanically connected, they cannot be ripped apart by the rather weak servo drives. However, a kinematically imperfect solution would result in mechanical tension and thus in an increase of joint torques and energy consumption. The concept, by which this constraint is maintained, is explained in Section 3.5. Due to the closed chain setup of the robot, segments are identified by cyclic IDs (e.g. Si = Si+Ns = Si+2Ns = \u00b7 \u00b7 \u00b7 with Ns being the number of segments of the robot). As depicted in Fig. 9, joints will be referred to as Ji, indicating connection to segments Si\u22121 and Si (see Fig. 9). The origin of a local segment coordinate system Fi is located on the axis of joint Ji, such that the x-axis points towards the next joint, Ji+1. Since the axis orientations and the length ls of all segments is identical also the transformation matrices are identical. The homogeneous matrix for the transformation from Fi\u22121 to Fi with the joint angle \u03b8i of joint Ji is: i\u22121T i = (cos \u03b8i \u2212 sin \u03b8i ls sin \u03b8i cos \u03b8i 0 0 0 1 ) (4) With this transformation, a point iP (relative to frame Fi) an be described relative to coordinate system Fi\u22121 via i\u22121P = \u22121T i \u00b7 iP ", + " Due to the high number of DoFs, collisions between segments of the robot are possible and must be actively avoided. For this purpose, a collision predictor was implemented that calculates the minimal distances between segments based on their positions a a f f I d c c f 3 l t s t r T T l ( l e s d v i p t l c o t o c 4 i s t o ( s r b s r t w d p p d T h t d t t m v u m t i e s n t a s n A u t s c i s T o c s s r s p i I t s s v 4 t s f u u c s obtained by the kinematics. Within this module, the segments re represented as 2D-capsules (rectangles with two rounded aces, see Fig. 9). However, to reduce the computational load, in a irst step, the distances between enclosing circles are determined. f the distance between any two of these circles is below the efined threshold, the actual distance between the respective apsules is computed. If the minimal distance threshold is underut, a penalty proportional to the overlap is added to the quality actor. .7.2. Robot-obstacle collisions In contrast to the collision avoidance between segments, the ocations of obstacle are not known prior to collisions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003164_s43236-021-00268-y-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003164_s43236-021-00268-y-Figure1-1.png", + "caption": "Fig. 1 Universal motor used in a food mixer", + "texts": [ + " Simulation and temperature measurements are also carried out by operating a universal motor with a DC supply using a bridge rectifier. The obtained results are compared with those of an AC supply, and the improvement in the thermal performance when using the DC supply is presented in this paper. The power consumption improvement, while operating a food mixer with a DC supply is discussed. The energy saving is the most significant improvement. Thus, this paper and its results are beneficial to society, since the life of food mixers can be increased. The universal motor used in the domestic food mixer shown in Fig.\u00a01 is considered for thermal analysis. Food mixers undergo intermittent operation. Food mixers are started with a load. Due to their intermittent operation, the load varies for consecutive cycles. The load is reduced when the food particles get crushed. The initial load torque is decided by the size and texture of the food in the mixer. The texture of the food and the speed of the mixer influence the slope at which the load torque falls with respect to time. Assuming that the losses alone are supplied at the end of the mixer operation, the load torque curves are drawn, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002993_s00170-021-07405-8-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002993_s00170-021-07405-8-Figure7-1.png", + "caption": "Fig. 7 Process diagram of precision sizing", + "texts": [ + " 6, the precision sizing method was designed according to the teeth accuracy requirements of target sun gear, where \u25b3L1 represents the sizing amount of external gear, and \u25b3L2 is the interference value of internal spline. The internal-external teeth will be reshaped by the whole tooth sizing method. This means that the tooth top, tooth flank, and tooth root will be all finished. The principle of precision sizing method is that the material radial flow is promoted by the finishing of external gear, resulting in the internal tooth fitting the splined mandrel more closely. Thus, the tooth accuracy of internal spline can be significantly improved, while the external gear accuracy is enhanced. Figure 7 shows the diagram of designed precision sizing process. The splined punch was first passed through the internal tooth by a press, and then the external gear was reshaped during the cold sizing process. The finite element simulation was conducted using the commercial software DEFORM. The error billet was established by measured tooth thickness andM value. In order to focus on the deformation of a complete tooth and save computer processing time, the FE model of precision sizing operation for cold extruded sun gear was performed using a 1/26 section of billet and tools (punch and die) due to the symmetrical structure, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002020_acemp.2007.4510512-Figure21-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002020_acemp.2007.4510512-Figure21-1.png", + "caption": "Fig. 21 Flux distribution.", + "texts": [ + " The static E&S model has greatly evaluated the loss from actual result, since the effect of the distorted waveform is disregarded. It is found that magnetic field is suppressed by the effect of the eddy current. The static analysis excessively analyses the magnetic field. 4 Magnetic Characteristic Analysis of Permanent Magnet Motor Figure 20 shows the concentrated flux typed surface permanent magnet motor (CSPM motor), which was developed for high density machine by our group and have been obtained by Oita University as the patent. The gap field of this motor has the 1.5 times of he magnetization of permanent magnet. Figure 21 shows the magnetic flux distribution. From this result it was found that the magnetic path becomes short, therefore the inductance increases as this motor. Figs. 22 and 23 show the distribution of the magnetic field strength and the magnetic flux density, respectively. As same as procedure of the vector magnetic characteristic analysis, the distribution of magnetic power loss is obtained as shown in Fig. 24. The magnetic power loss increases in the place that both vectors B and H are large and the axis ratio in rotational magnetic flux is large" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000785_12.822389-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000785_12.822389-Figure1-1.png", + "caption": "Fig. 1. Experimental setup.", + "texts": [ + " The RBCs and the hardened RBCs were suspended in autologous plasma. Furthermore the washed RBCs were suspended in isoosmotic 2.5% solution of Dextran 70 kDa. In all three cases the haematocrit of the samples was adjusted to 5%. The experiment was carried out according to the ethical guidelines laid down by the Bioethical Commission at the Ludwik Rydygier Collegium Medicum in Bydgoszcz of the Nicolaus Copernicus University. The samples were investigated using the experimental setup depicted in Fig. 1. As a light source a halogen lamp is used. Collimated and filtered (570 nm) light is scattered by the sample. The forward scattered light was spatially filtered by an aperture placed in a Fourier plane of the optical system. This aperture reduces the high intensity of unscattered component of the light. A CCD camera (1024 x 768 pixels, 10 bit depth) is placed in the image plane of the optical system. The resolution of the experimental setup was limited by the resolution of CCD camera and was 23" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002179_s0263574720001289-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002179_s0263574720001289-Figure2-1.png", + "caption": "Fig. 2. One oar of the swimmer and its rotational angles (\u03b1n, \u03b2n).", + "texts": [ + "org/10.1017/S0263574720001289 Downloaded from https://www.cambridge.org/core. Clark University, on 03 Jun 2021 at 06:05:32, subject to the Cambridge Core terms of use, available at https://www.cambridge.org/core/terms. or stepwise about the perpendicular axis of the body link. The length of the body links are l + s1 and l + s2, where sn (n = 1, 2) is the expansion and contraction length of the linear actuators. The rotation of each disk about the robot body axes is measured by one angle \u03b2n (see Fig. 2). It should be noted that the angle \u03b1n is fixed and does not change with time. The inertial coordinates are denoted by (X1 X2 X3) and its corresponding basic vectors are (E1 E2 E3) respectively. The position of the swimmer mass center is defined as Xc = (Xc1 Xc2 Xc3) describing the translational motion of the swimmer in the inertial coordinates. The swimmer orientation is determined by Euler angles, (\u03c6 \u03b8 \u03c8). We choose the order of axes transformation as \u03c8 \u2192 \u03b8 \u2192 \u03c6 with axes order of rotation as 3 \u2192 2 \u2192 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003052_icet51757.2021.9451096-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003052_icet51757.2021.9451096-Figure1-1.png", + "caption": "Fig. 1. The DMRW UAV.", + "texts": [ + " The contributions of the paper are described as follows: (1) a high-fidelity six/seven-degree-of-freedom model of the compound UAV is built based on blade element method (BEM); (2) flight characteristics of all three modes and mode transitions are analyzed in detail; (3) a robust adaptive fullmode controller based on derivative-free adaptive neural network and control allocator is designed and verified with numerical simulations. The control scheme could be adopted to design flight control system of rotor/wing compound UAV or e-VTOL. The rotor/wing compound UAV studied in this work is depicted in Fig. 1, which is called double mode rotor and wing UAV, DMRW UAV for short. Tip-jet technology is adopted to drive the rotor in helicopter flight mode, which is mechanically simpler than a shaft-driven rotor by eliminating gearbox and transmission of both main rotor and tail rotor with weight, complexity and life-cycle costs significantly reduced. The typical flight profile and flight stage division of the compound UAV are described in Fig.2. And flight mode are 960 978-1-7281-7673-4/21/$31.00 \u00a92021 IEEE 2 0 2 1 I E E E 4 th I n te rn at io n al C o n fe re n ce o n E le ct ro n ic s T ec h n o lo g y ( IC E T ) | 9 7 8 -1 -7 2 8 1 -7 6 7 3 -4 /2 0 /$ 3 1 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000185_11802372_33-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000185_11802372_33-Figure8-1.png", + "caption": "Fig. 8. Reference frame of child robot", + "texts": [ + " After the mission is fulfilled successfully, all the robots are waiting for a new mission transferred from father robot. Fig.6 shows that when a single child robot4 cannot climb up a slope, other three robots connected with it end to end and help it to overcome this obstacle. In order to help robo4 climb up the slope successfully, the distance between each two child robot should be adjusted during the process of climbing, such as d32 and d43 in Fig.7. Therefore, the inverse kinematics calculation is inescapable. The reference frame of a single robot is shown in Fig.8. The D-H parameters are shown in table.1.Where L1 to L6 are 60cm, 10cm, 60cm, 40cm, 20cm and 5cm respectively. By static analyzing, we found the structure that robot2 and robot1 stay under the bottom of slope can provide a maximum climbing force for robot4. We also obtained an expression, sin 1 cos \u03b1\u03bc \u03b1 \u2265 + . (1) Where \u03bc is a friction coefficient between child robot with the slope, \u03b1 is the maximum gradient of the slope the multirobot team can climb, as shown in Fig.9. Suppose the maximum gradient for a single robot is \u03b2 , arctan( )\u03b2 \u03bc= " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002004_50015-9-Figure12.13-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002004_50015-9-Figure12.13-1.png", + "caption": "Figure 12.13 Transverse section showing multi-component structure of a downhill ski ( from Easterling, 1990, by permission of the Institute of Materials, Minerals and Mining).", + "texts": [ + " The main requirements of the ski are good strength and flexibility along the length with torsional stiffness, so that the skier\u2019s weight is properly distributed while traversing the irregular snow contours. In addition, it is necessary to dampen the ski structure to absorb dynamic impact loading. To meet these demands a multi-layer structure has evolved consisting of a base, usually polyurethane, a shock-absorbing core, usually natural ash or hickory wood, fiberglass and an elastomeric secondary core, steel or high-strength aluminum alloy edges, side walls of glass materials, top layer and reinforcing damping layers. An early design is shown in Figure 12.13, but this has evolved with the exact design and processing commercially guarded in a very competitive industry. In parallel with material development, modern skis have been reduced in length from around 200 cm in 1990 to around 160 cm. The ski is only part of the equipment necessary for skiing. Boots, bindings and ski poles are also needed. Ski boots must provide a firm grip on the skier\u2019s ankles to allow proper control of the ski. External Hytrel\u2013Kevlar components have been used to stiffen the ski boot for this purpose (Hytrel is a thermoplastic polyester elastomer)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002581_icees51510.2021.9383730-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002581_icees51510.2021.9383730-Figure5-1.png", + "caption": "Fig 5. Rotor Geometry with 2 Permanent Magnets", + "texts": [ + " To give most economical balance between the costs of dies, the peripheral length usually between 0.3 m to 0.6 m is chosen for one segment, the left over amount of scrap from the lamination cuttings from steel strips and the assembly cost. In the flux paths of alternating poles, to provide an equal number of turns, the total number of segments is chosen. Fig 4 shows the stator geometry of 36 slots. F. Rotor Geometry In BLDC motor two types of rotor design are there, \u2022 Outer Rotor \u2022 Inner Rotor The windings are located in the motor core for an outer rotor design. Fig 5 shows the stator windings surrounded by the rotor magnets. 2021 7th International Conference on Electrical Energy Systems (ICEES 2021) 119 Authorized licensed use limited to: Carleton University. Downloaded on June 06,2021 at 01:21:49 UTC from IEEE Xplore. Restrictions apply. G. Introduction o f fins The fins are the extended surface that is being used to decrease the thermal resistance at the motor\u2019s solid parts such as stator, magnets and rotor. Thereby increases the rate of heat transfer from surface to adjacent fluid" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001190_09544062jmes539-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001190_09544062jmes539-Figure8-1.png", + "caption": "Fig. 8 Pressures and film thicknesses on the plane of symmetry at the instant \u03c9t = 1.5\u03c0 corresponding to the results in Fig. 7 for A = 8 and 10 \u00b5m, respectively", + "texts": [ + " In detail, for the vibration of A = 8 \u00b5m, the lowest load carrying capacity is Cw = 4.12 \u00d7 10\u22124, and the maximum coefficient of friction is \u00b5 = 0.244, both occur at the instant \u03c9t = 1.322\u03c0 (238\u25e6); and, for the vibration of A = 10 \u00b5m, the lowest load carrying capacity is Cw = 3.55 \u00d7 10\u22125, and the maximum coefficient of friction is \u00b5 = 2.55, both occur at the instant \u03c9t = 1.244\u03c0 (224\u25e6). The pressures and film thicknesses on the plane of symmetry at \u03c9t = 1.5\u03c0 are plotted in Figs 8(a) and 8(b) for these two cases respectively. Note the unit of pressure in Fig. 8 is MPa rather than GPa. It should also be pointed out that a relatively large inlet boundary location was chosen in order to cater for this. For a fixed vibration amplitude A = 8 \u00b5m, a series of time-dependent solutions with various vibration frequencies of f = 250, 500, 1000, and 2000 Hz, together with a quasi-steady-state solution were achieved under thermal and non-Newtonian conditions. For these solutions, the cyclic variations in Cw, \u00b5, pcen, Tcen, hcen, and hmin are plotted in Fig. 9. As expected, in general, the higher the vibration frequency, the stronger the transient effect" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002925_01423312211016188-Figure9-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002925_01423312211016188-Figure9-1.png", + "caption": "Figure 9. Test bearings.", + "texts": [ + "1 3 10\u20134 s, and some peaks also appear in the envelope spectrum near the theoretical fault characteristic frequency of 36.29 Hz and its harmonic frequency. In order to verify model effectiveness, the vibration signal acquisition experiment was carried out on the established bearing fault simulation device, and the inner and outer ring fault bearings were tested. The experimental device is displayed in Figure 7. The arrangement of test points and triaxial vibration acceleration sensors is shown in Figure 8. The test bearings shown in Figure 9 were installed on the right bearing seat in turn. The rotating speed and radial load were set as 4 Hz and 1000 N, respectively. The actual bearing and local fault parameters are the same as simulation settings in Tables 1 and 2. According to the acquisition result, the Z-axis signal was selected for analysis. The sampling frequency is 20 kHz and the experimental results are shown in Figures 10 and 11. The comparison results show that the time domain waveform, shock details, spectrum and envelope spectrum characteristics of simulation signals and experimental signals are similar" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000088_j.jsv.2006.01.044-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000088_j.jsv.2006.01.044-Figure4-1.png", + "caption": "Fig. 4. A crack beam element in general loading.", + "texts": [ + " [17,20] estimated the stiffness of a cracked rotor by introducing a local flexibility matrix. Ref. [23] represented the cracked section by a spring to model a simply supported beam. In this study, the stress intensity factors presented in Ref. [24] are applied and a new beam element with 451 slant surface crack is developed. Consider a rectangular cross-section beam with given stiffness properties, the width and the height of the cross-section are w and h, respectively. The depth of the slant transverse crack is a and the angle towards the cross-section of the beam is 451 (see Fig. 4). The beam is loaded with axial force P1, shear forces P2 and P3, bending moments P4 and P5, and torsional P6. The dimension of the local flexibility matrix depends on the numbers of freedom (each node has six degrees of freedom), here is 6 6. The crack is assumed to affect only the beam stiffness and do not affect the mass distribution (still be same as that of the 2-node Euler\u2013Bernoulli beam element), the stiffness matrix can be calculated as follows. ARTICLE IN PRESS X.Y. Lei et al. / Journal of Sound and Vibration 295 (2006) 890\u2013905 895 Paris\u2019 equation [17] gives the additional displacement mi due to a crack of depth a, in the i direction, as mi \u00bc q qPi Z a 0 J\u00f0a\u00de da , (3) where J(a) is the Strain Energy Density Function (SEDF) and Pi is the corresponding load" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002513_icara51699.2021.9376432-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002513_icara51699.2021.9376432-Figure1-1.png", + "caption": "Fig. 1. ActivityBot Robot", + "texts": [ + " It is an improvement to the CriticalPointBug algorithm. It works in a dynamic environment, assume the space contains bounded stationery and bounded dynamic obstacles. The detected point by the range sensor is called an open point. Depending on the coordinates of the open point, the sub-goal point is calculated. From the sub-goal points, one point is chosen to move to, which has a minimum distance to the goal, which the robot did not visit before. The robot used in this work is the ActivityBot robot, which is shown in Fig. 1a. ActivityBot robot is one of the parallax products [13]. It has a microcontroller called the Propeller microcontroller, which has the feature of multi-core. The Propeller microcontroller has eight cores. This feature allows us to do various tasks simultaneously, as each core can operate separately from others. 27 Authorized licensed use limited to: University of Prince Edward Island. Downloaded on June 02,2021 at 05:11:41 UTC from IEEE Xplore. Restrictions apply. Three sensors had been implemented in the ActivityBot robot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001573_s09-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001573_s09-Figure1-1.png", + "caption": "Figure 1. Schematic diagram of the facility for laser evaporation of nuclear fuel specimens at temperatures above 2000 C: 1\u00d0CO2 laser beam, 2\u00d0 plasma circuit breaker, 3\u00d0shutter, 4\u00d0vacuum balance, 5\u00d0compensation circuit, 6\u00d0ballistic collector with the beam output window, 7\u00d0 turntable for targets, 8\u00d0fixed mirror, 9\u00d0beam window, 10\u00d0lens, 11\u00d0movable mirror, 12\u00d0vacuum chamber, 13\u00d0rapid-response micropyrometer for measuring the temperature in the focal spot of evaporation, 14\u00d0camera to monitor the gas jet.", + "texts": [ + " 28, 050043 Almaty, Republic of Kazakhstan Tel. (727) 309 63 22. E-mail: roza_j@mail.ru Received 25 August 2011, revised 3 October 2011 Uspekhi Fizicheskikh Nauk 182 (6) 645 \u00b1 648 (2012) DOI: 10.3367/UFNr.0182.201206d.0645 Translated by V I Kisin; edited by AM Semikhatov stationary, uniform, driven evaporation is achieved, the overall composition of the gas becomes well defined and identical to the isotopic composition of the fuel, which is significantly different from the actual composition of the laser-heated surface. Figure 1 shows a schematic diagram of the laser-heating facility that is capable of heating specimens of nuclear fuel to temperatures above 2000 C. Carbon dioxide O2 , weakly ionized by glow electric discharge and heated to the plasma state, emits laser beam 1 (see Fig. 1) of about 200 MW in the infrared range at a wavelength of about 5.10 mm. To reduce the high-amplitude pulse fed to the electrodes of plasma injectors of surface discharge and at the same time maintain the density and volume of the plasma produced by them, a plasma circuit breaker is used (2). It includes the source of a current pulse connected to two extended electrodes. The electrodes are separated by an insulator, and together they form a vacuum interelectrode gap and a load in the form of a vacuum or plasma diode connected to the electrodes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002782_j.matpr.2021.04.597-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002782_j.matpr.2021.04.597-Figure3-1.png", + "caption": "Fig. 3. Circular loop structure.", + "texts": [ + " Multiple resonance and wide band applications are the most popular applications for multi-screen FSS. The spacing is the most important factor to consider when designing a multi resonance FSS. This paper describes frequency selective surfaces in the reflection band with three closely spaced resonances. Individual components must be very closely spaced in order to achieve near resonances. The choice of fss is heavily influenced by the transverse electric (TE) and transverse magnetic (TM) fields. The proposed model is designed with circular ring patches as shown in Fig. 3 is printed on 10 mm thick Rogers RT Duroid 6002 substrate having a permittivity value of 2.94. For avoiding grating lobes appearance, unit cell size is decided based on angle of plane wave incidence and highest frequency of operation. Reflection band is Ka band whose transmission is in the range of 35 dB to 45 dB. The S and Ku bands have transmission ranges of 0.15 dB and 0.25 dB, respectively. Up to 40 degrees, the result indicates strong polarization and angle of incidence stability. The proposed FSS took advantage of a large bandwidth in the reflection band, with 61" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001997_978-1-84800-066-7_5-Figure5.4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001997_978-1-84800-066-7_5-Figure5.4-1.png", + "caption": "Fig. 5.4. Ball and beam.", + "texts": [ + " Finally, note that the Jacobian matrix \u2202\u03b2 \u2202\u03be is invertible, in particular \u2202\u03b2 \u2202\u03be = [\u221a m11(y2) 2k2y1\u03be2 \u221a m11(y2) 0 1 ] , and \u03b2(\u03be, y) is invertible with respect to \u03be. The result is summarised in the following statement. Proposition 5.2. Consider the system (5.36), where \u03b21(\u00b7) and \u03b22(\u00b7) are given by (5.42) and (5.43) respectively, and suppose that Assumptions 5.1 and 5.2 hold. Then the system (5.36) has a uniformly globally exponentially stable equilibrium at the origin. Consider the \u201cball and beam\u201d system16 depicted in Figure 5.4, which is described by equations of the form (5.34), where the states y1 and y2 correspond to the position of the ball and the angle of the beam, respectively, F (y, \u03b7) = 0, 15Note that in this case \u03bd may be negative. 16See [67, 72] for details on this problem. M(y) = [ B +my2 2 0 0 A ] , C(y, \u03b7) = [ my2\u03b72 my2\u03b71 \u2212my2\u03b71 0 ] and G(y) = [ mgy2 cos(y1) mg sin(y1) ] , where A = Jb R2 +m, B = J+Jb, m, R and Jb are the mass, radius and moment of inertia of the ball, respectively, J is the moment of inertia of the beam, g is the gravitational acceleration, and \u03c4 is the applied torque" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002309_icccr49711.2021.9349392-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002309_icccr49711.2021.9349392-Figure1-1.png", + "caption": "Figure 1. Robot connecting rod coordinate system.", + "texts": [], + "surrounding_texts": [ + "relationships in the process of active assembly of the circular shaft hole, and derives the relationship between the assembly lateral error and the assembly force and moment, and proposes a method for searching and finding holes and compensating the position error of the shaft hole. Active assembly method. Perform statics analysis on the robot and establish a mechanical model of the robot in a static state. The joint torque obtained by the sampling function of the robot's own joint motor current is used to analyze the force in the assembly process, to obtain the size and direction of the assembly error, and realize the error judgment without external sensors. Afterward, the shaft hole assembly process can be completed through the error adjustment algorithm.\nKeywords-automatic assembly; error compensation;\nJacobian; sensorless\nI. INTRODUCTION\nRobot assembly operation is an important direction of robot application research. At present, the application of robot automatic assembly technology is restricted by assembly accuracy. For the high-precision assembly process, the error of the robot's motion accuracy exceeds the assembly accuracy, which makes it difficult for the robot to achieve precise assembly work [1]. Therefore, it is a challenging task for industrial robots to complete precision assembly operations. In the assembly system, the assembly of shaft holes and similar shaft holes occupies a large proportion, and the assembly process includes hole searching and jacks.\nThe data obtained through the vision system can provide the robot with rich spatial position information, thereby guiding the robot to complete more complex tasks autonomously. The Crb robot developed by Shanghai Jiao Tong University can assemble the 0.6mm aperture. However, the existing visual positioning methods require professional technicians to perform complex calibration and teaching of cameras and robots according to different working\nconditions, which lack flexibility and adaptability in practical applications. Therefore, a new target recognition and positioning method is needed to meet this actual demand.\nHailong Li has developed a flexible axle hole assembly robot, which can locate the target parts through visual sensors and control the entire assembly process according to the detected assembly force information [2]. In terms of structure, ABB series robot is used as the actuator, CCD and six-dimensional force sensor are used as the vision and force perception system of the robot respectively. In the contact adjustment stage, there is contact between the assemblies. The robot recognizes the assembly state through the sensor information of the force sensor, adjusts the posture of the assembly shaft, and makes preparations for the insertion assembly. The insertion of assembly is to move in a straight line along the assembly direction until the assembly is completed. The system has a good compliance control ability, can actively position the workpiece posture, posture deviation adjustment through force feedback, successfully complete the assembly. However, the visual positioning and force perception control of the system work independently, and their cooperative control should be studied in the future to further improve the intelligence of the robot [3].\nResearch on the robot vision positioning method suitable for the automated assembly process of flexible production cells can reduce the manufacturing costs of special fixtures, reduce equipment downtime caused by changes in working conditions and the time consumed by complex calibration processes, and improve the flexibility and adaptability is of great significance in broadening the application fields of industrial robot, improving the intelligent level of industrial robots, and improving production efficiency [4].\nBased on the analysis of the force relationship in the circular shaft hole assembly, this paper proposes a torque feedback based on the current loop. The torque of each joint is obtained through the current of each joint of the robot. The calculation is completed by decoupling the mechanical model of the six-axis robot. So as to determine the size and direction of the deviation of the shaft hole.\n50\n978-1-7281-9035-8/21/$31.00 \u00a92021 IEEE\n20 21\nIn te\nrn at\nio na\nl C on\nfe re\nnc e\non C\nom pu\nte r,\nC on\ntro l a\nnd R\nob ot\nic s (\nIC C\nC R\n) | 9\n78 -1\n-7 28\n1- 90\n35 -8\n/2 0/\n$3 1.\n00 \u00a9\n20 21\nAuthorized licensed use limited to: University of Canberra. Downloaded on May 21,2021 at 20:29:08 UTC from IEEE Xplore. Restrictions apply.", + "II. BASIC ANALYSIS OF ROBOT MODEL\nThe key to modeling the three-dimensional structure of an industrial robot is the determination of the shape and size. The equipment used in our experiment is the St\u00e4ubli TX90 explosion-proof robot, and the SolidWorks threedimensional structure model is obtained according to the robot. Starting from the homogeneous transformation of robot coordinates, a coordinate system is established on each joint, a D-H model is established for the robot, and the mechanical model of the robot is analyzed to find the relative function relationship between the force of each joint and the force of the end\nThe robot arm is composed of a series of connecting arms assembled with each other. In order to describe the relationship between the adjacent arms of the robot, the robot in the figure is modeled by the D-H method, and the reference coordinate system of each joint of the robot is established [5]. The coordinate system of the robot arm linkage is shown in the figure, and the parameters obtained according to the D-H method are shown in the table.\nThe forward kinematics analysis of the manipulator refers to the calculation of the position and posture of the end relative to the reference frame after the parameters and joint angles of the manipulator are known. The essence of positive kinematics transformation is the transformation from joint space to operation space. The transformation relationship\nbetween the adjacent joint coordinate system i-1 and i of the robot is:\n-1 1 1=Rot( , ) ( , ) ( , ) ( , )\ncos sin cos sin sin cos\nsin cos cos cos sin sin\n0 sin cos\n0 0 0 1\ni i i i i i i i i iT z Trans z d Trans x a Rot x a\ni i i i i ai i\ni i i i i ai i\ni i di\n\n \n \n \n\u2212 \u2212\n\u2212 \u2212 = \n(1)\nThe transformation matrix of each adjacent joint can be calculated by above:\n6 6\n6 65\n6\n6\n0 0\n0 0\n0 0 1\n0 0 0 1\nc s\ns c T\nd \u2212 = \n(2)\n5 5\n5 54\n5\n0 0\n0 0\n0 1 0 0\n0 0 0 1\nc s\ns c T \u2212 = \n(3)\n4 4\n4 43\n4\n0 0\n0 0\n0 1 0 0\n0 0 0 1\nc s\ns c T\n \u2212 = \n(4)\n3 3\n3 32\n3\n0 0\n0 0\n0 1 0 0\n0 0 0 1\nc s\ns c T\n \u2212 = \n(5)\n2 2 2 2\n2 2 2 21\n2\n0\n0\n0 0 1 0\n0 0 0 1\ns c a s\nc s a c T\n\u2212 \u2212 \u2212 \u2212 = \n(6)\n1 1 1 1\n1 1 1 10\n1\n1\n0\n0\n0 1 0\n0 0 0 1\nc s a c\ns c a s T\nd\n \u2212 = \n(7)\nIII. MECHANICAL MODEL OF ROBOT\nThe driving device of each joint of the robot provides joint forces and torques, which are transmitted to the endeffector through the connecting rod to overcome external forces and torques. The relationship between the driving force and torque of the joint and the force and torque exerted by the end-effector is the basis of manipulator force control.\nThe Robot Jacobian [7] reveals the mapping relationship between the operating space and joint space, and also\n51\nAuthorized licensed use limited to: University of Canberra. Downloaded on May 21,2021 at 20:29:08 UTC from IEEE Xplore. Restrictions apply.", + "represents the force transfer relationship between the two, which provides a convenient way to determine the static joint torque of the robot.\nThe first three rows of Jacobian of a six-DOF robot operating in three-dimensional space represent the transfer ratio of the linear velocity of the hand to the joint velocity, while the last three rows represent the transfer ratio of the angular velocity of the hand to the joint velocity. And each column of the Jacob matrix represents the transfer ratio of the corresponding joint velocity, the linear velocity of the hand, and the angular velocity. According to the definition of the Jacobian matrix, the differential construction method is used to construct the Jacobian matrix of TX90 robot. The i column of relative Jacob matrix Tj(q) can be obtained by definition:\n( )\n( )\n( )\n( )\ni i\nz x y y x\ni i\nx y y xz\ni i x y y xT z\nzi\nz\nzi\nz\nzi\nz\np n n p n p\np o o p o p\na p a pp aj q n\nn o o a a \u2212 + \u2212 + \u2212 + = = \n \n(8)\nFrom this equation we can calculate the Jacob matrix:\n( )\n( )( )\n( )( )\n( )\n2 2 1 4 23 4 5 6 4 6\n2 2 1 4 23 4 5 6 4 6\n2 2 1 4 23 4 5\n1\n23 4 5 6 4 6 23 5 6\n23 4 5 6 4 6 23 5 6\n23 4 5 23 5\n-\n( )\n( )\nT\na c a d s s c c c s\na c a d s s c c c s\na c a d c s s j q\ns c c c s s c s s\ns c c c s s c s s\ns c c c c\n + \u2212 + + \u2212 \u2212 + \u2212 = \u2212 \u2212 \u2212 \n \u2212 \u2212 + \u2212 \n(9)\n( )\n( )( ) ( ) ( )( ) ( )\n( )( ) ( )\n2 4 3 3 4 5 6 3 4 6 3 5 6 4 3 3 4 5 6 3 4 6 3 5 6\n2 4 3 3 4 5 6 3 4 6 3 5 6 4 3 3 4 5 6 3 4 6 3 5 6\n2 4 3 3 4 5 3 5 4 3 3 4 5 3 4 6 3 5 6\n2\n4 5 6 4 6\n4 5 6 4 6\n4 5\nT\na d s s c c c s s c c s s d c c c c c c s s s s c\na d s s c c s s s c c s s d c c c c s c s c s s s\na d s s c s c s d c c c c c s c s s s j q\ns c c c s\ns c s c c\ns s\n \u2212 \u2212 + \u2212 \u2212 \u2212 \n\u2212 \u2212 \u2212 \u2212 + + \u2212 \u2212 \u2212 \u2212 + + \u2212 = \u2212 \u2212 \n\u2212 \n (10)\n( )\n( )\n( ) 4 4 6 4 5 6 4 5 6 4 4 6 4 5 6 4 5 6\n4 5 5\n3\n4 5 6 4 6\n4 5 6 4 6\n4 5\n+\nT\nd s s c c c c s d\nd s c c c s c c d\nd c s j q\ns c c c s\ns c s c c\ns s\n \u2212 + + = \u2212 \u2212 \n \u2212 \n(11)\n( )\n5 6 6\n5 6 6\n5 5 6\n4\n5 6\n5 6\n5\nT\nc c d\ns s d\nc s d j q\ns c\ns c\nc\n = \u2212 \n(12)\n( )5\n6\n6\n0\n0\n0\n0\nT j q s\nc\n = \u2212 \u2212 \n(13)\n( )6\n0\n0\n0\n0\n0\n1\nT j q\n = \n(14)\nTj(q) is the Jacobian matrix relative to the end coordinate system, that is, given the generalized velocity of each joint relative to the end coordinate system, the generalized velocity of the end relative to the end coordinate system can be solved through the Jacob matrix. Meanwhile, Tj(q) can represent the static force relationship between the generalized force vector of the end and the joint force vector when the equilibrium condition is expressed. The difference between j(q) and Tj(q) is that the reference frame is the base system. The relationship between the two can be expressed as:\n( ) ( ) 0 6\n0 6\n0 j\n0\nTR j q q\nR = \n(15)\nIn order to express the force and moment of the workpiece to be assembled (referred to as generalized force F at the end point for short), fn and Mn can be combined into a six-dimensional vector:\nx\ny\nz\nx\ny\nz\nf\nf\nf F\nM\nM\nM\n = \n(16)\nThe driving force or driving torque of each joint can be written in the form of a six-dimensional vector:\n1\n2\n3\n4\n5\n6\n\n\n \n\n\n\n = \n(17)\nAccording to the virtual work principle:\n52\nAuthorized licensed use limited to: University of Canberra. Downloaded on May 21,2021 at 20:29:08 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_83_0000534_s0065-2458(08)60221-1-Figure48-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000534_s0065-2458(08)60221-1-Figure48-1.png", + "caption": "FIG. 48. Bending of u liquid jet.", + "texts": [ + " GLAETTLI (52) or if (49) and (52) are combined, by (53) Equation ( 3 2 ) mentions a fact of great convenience for model tests: large elements are a good substitute for a high speed camera. Equation (53) is especially useful in estimating the lower limit of the response time when the minimum Reynolds number is known and if there exists a maximum operating pressure that does not depend on the size. This is, in fact, true a t least for the wall interaction type amplifier. It can be explained very easily be considering the pressure difference Ap across a reattached jet such as shown in Fig. 48 when an incompressible fluid is used. This pressure is determined by (54) 1 Ap = - * p * v 2 . r If atmospheric pressure is present on thc outer side of the jet, Ap is limited (pv is the vapor-pressure of the liquid in question). Introducing (54) into (55) and combination with (48) results in by (55) The existence of such an upper pressure limit can easily be verified. Figure 49 shows a graph similar to Fig. 47. The pressure, however, instead of the Reynolds number is plotted on the abscissa. The breakdown observed at p,,,, does not mean that amplification finds an end in principle: it is true that suction control is no longer possible" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001152_006-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001152_006-Figure2-1.png", + "caption": "Figure 2.Maps of the surface density (upper panel), radial velocity (middle panel), and deviations of the angular velocity from the Keplerian profile (bottom panel) according to numerical model [34]. Velocities are in units of the sound speed cT. The thick lines show zero velocity levels.", + "texts": [ + " But because the flux should be conserved along the stream line, the matter density should also change along the stream line and reach amaximum at the apastron. Therefore, the spiral arm formed by the stream line apastrons look likes a density wave, as was shown in [34]. In the observer's reference frame, the precessional wave is almost steady: its shifts by 1 \u00ff3 in the retrograde direction in one orbital period [34]. Hence, the density and velocity distribution in the wave can be considered stationary on time scales of the order of several dozen characteristic disk periods. Figure 2 shows distributions of the surface density and the radial and angular velocity in numerical simulations of the disk in a close binary system [34]. The calculations were 1 The Knudsen number here is the ratio of the particle mean free path to the characteristic scale of the problem, which here can be the disk thickness. For a typical number density in the accretion disk of the order of 1011 cm\u00ff3, the free path length is 10\u00ff10 a.u., while the disk thickness is 910\u00ff2 a.u. carried out for the binary with the following parameters: the accretor mass M1 1M , the donor massM2 0:05M , the binary separation A 0:625R , and the binary orbital period Porb 4830 s", + " This behavior is clear because matter moving along an eccentric orbit can more towards or away from the star at different parts of the trajectory. Nevertheless, it should be noted that in the outer parts of the disk, the radial velocity is positive for all profiles, because the angular momentum of the disk decreases in the outer parts due to the decretion of matter. We stress that the radial velocity difference over the disk is fairly large and along some directions can be as high as several dozen times the speed of sound. As follows from the distributions presented in Fig. 2, the precessional wave is tightly wound. This allows approximately treating perturbations caused by the wave as axially symmetric. This approximation cannot be valid in the central parts of the disk, where the axial symmetry is significantly violated (which can be seen most clearly in the velocity distribution). In what follows, we analyze perturbations excluding the central part of the disk with a radius smaller than 0:08A from calculations. The results of calculations suggest that the outer parts of the disk are subjected to strong gas-dynamic perturbations", + " We write Eqns (7) and (8) ignoring geometrical terms and the tangential velocity: qd qt u0 qd qr qu qr 0 ; 31 qu qt u0 qu qr qu0 qr u c 2T qd qr 0 : 32 For the background velocity variation law u0 / r\u00ff r , in the limit u0 5 cT, it is easy to show (by differentiating the Euler equation with respect to the radius and by eliminating velocity perturbations) that solutions take the form d / exp \u00ffiot r\u00ff r and u / exp \u00ffiot r\u00ff r 2 such that in the immediate vicinity of the point r , the perturbation amplitude decreases and the spatial oscillations become `frozen'. Figure 6 demonstrates that the perturbation amplitude reaches a maximum in the inner part of the disk restricted by the precessional wave, while the turbulence (understood as large values of a) is strongest in the outer part of the disk. This is most clearly seen in the second and third quadrants of the disk, where the location of the amplitude maximum coincides with the radial velocity maximum (see Fig. 2). The perturbation amplitude decreases with decreasing the velocity. In the region bounded by the zero radial velocity, some amplitude growth is observed, especially close to the outer edge of the precessional wave. In the fourth quadrant, where the radial velocity in the inner part of the disk is negative (see also Fig. 3), perturbations and turbulence are almost absent, but there is an amplitude peak at the outer edge of the wave. This behavior of the amplitude can be explained as follows. We use Eqns (7)\u00b1(9) in the approximation of a simulation box whose sizes are much smaller than the distance from the box to the center, as in [6]", + " In this case, it is possible to argue that the rear phase of the perturbation catches up with its front phase in one wave period.Here, the amplitude of perturbationmaxima increases by the mass conservation. In terms of the dispersion relation, this signals the appearance of nonzero values of the perturbation growth increment. We note once again that in our method of instability analysis, we assume axially symmetric perturbations, while the background density and velocity distributions do not have axial symmetry and weakly depend on the angle with the characteristic angular scale 2p (see Fig. 2). In a real disk, perturbations with this or smaller angular scale would shift in the tangential direction due to the background rotation. This could weaken the growth of perturbations due to the radial velocity gradient. However, the necessary condition adopted here for the turbulence to appear, Eqn (25), requires that the characteristic growth time of perturbations be shorter than the Keplerian time. Thus, a perturbation in its growth time cannot leave the region of the precessional wave in the tangential direction and we therefore observe the instability growth", + " Although the expansion of system (11) in the orthogonal set of functions is quite admissible, we cannot be sure that functions (42) and (43) belong to two dual sets in the sense of definitions (39) and (40) for all f from 0 to 2p. Despite this fact, we prefer this method because it offers a more clear physical interpretation of the perturbation growth as a function of the background variable distribution, ensuring the angular coordinate locality. Another assumption is that the proposed approach assumes the use of axially symmetric modes, whereas the background solution depends on the angle. The distributions shown in Fig. 2 demonstrate that the main angular scale of changes is 2p. Therefore, this scale should mainly contribute to perturbations, and in the standard approach, this scale would correspond to the azimuthal number m 1. However, for simplicity, we assume that the perturbation is axially symmetric. Thus, the problem can be formulated independently for each radial direction in the disk. Below, we briefly describe the instability analysis in a disk following the Lynden-Bell\u00b1Ostriker method. A detailed presentation of the method and examples for axially symmetric flows can be found in [49]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000528_2007-01-4140-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000528_2007-01-4140-Figure2-1.png", + "caption": "Figure 2: One Dimensional Pivoted Tilting Pad Bearing", + "texts": [ + " Bartholme and Richard Hergert 2 III: Boundary Lubrication ( <1) In this regime, the entire applied load is carried by the surface asperities. The friction and wear are dependant on the lubrication properties of surface molecules. Friction coefficients in boundary lubrication regimes are typically between 0.2 and 0.4, even rising as high as 1 in some cases. The friction and wear of the material is dry sliding. The result of lubrication regimes during operation with effects on bearing friction is shown in Figure 1. Tilting pad bearings as illustrated in Figure 2 are designed to operate in the hydrodynamic regime. In other words, during normal operation the two bearing surfaces are separated by a fluid film. The factors influencing the formulation of the fluid film will determine the tilting pad bearing limitations. Significant research has been accomplished on tilting pad bearings since Kingsbury\u2019s [3] research several decades ago. Raimondi [4] [5] presented several papers which utilized the use of charts and method-of-solution tables to aid in identifying the desired solution" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000866_detc2007-34052-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000866_detc2007-34052-Figure1-1.png", + "caption": "Figure 1. Outside blade of the grinding wheel (tooth concave side).", + "texts": [ + "14) in [9]), which is basically a compact way to express the well known Euler finite rotation formula [10] (often referred to as Euler\u2013Rodrigues formula [11]), which describes the rotation of a vector about an axis by a given angle. Therefore, transformation matrices are never employed in the invariant approach. The tool employed in the grinding process of spiral bevel pinions has an axisymmetric surface \u03a3e obtained by revolving the blade profile about the cutter axis. A schematic representation of the outside curved blade profile, conjugate to the concave side of the tooth, is shown in Fig. 1. Details on the inside blade profile are omitted for brevity. Part I of the blade profile is a straight line, while parts II and III are portions of two circles with radii \u03c1 f (edge radius) and R1 (spherical radius), respectively. The blade profile has a C1 boundary. Rp is the cutter point radius and \u03b1p is the blade pressure angle. 2 ownloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx? In the linear space R 3 we introduce a single reference system S = (O; x, y, z), with unit vectors (i, j,k), as illustrated in Fig. 1. In this frame, the position vector pe(\u03be, \u03b8) of the generating (tool) surface \u03a3e is represented by its components pe(\u03be, \u03b8), where \u03be is the arc length measured on the blade profile and \u03b8 is the revolution angle about the z-axis. The parametric equations of the topland of the tool, part I, are given by pe(\u03be, \u03b8) = xe(\u03be, \u03b8) ye(\u03be, \u03b8) ze(\u03be, \u03b8) = (X f \u2213 \u03be) cos \u03b8 (X f \u2213 \u03be) sin \u03b8 0 ; \u03be \u2264 0 (1) the blade tip fillet, part II, is defined by pe(\u03be, \u03b8) = [ X f \u2213 \u03c1 f sin \u03c1\u0303(\u03be) ] cos \u03b8 [ X f \u2213 \u03c1 f sin \u03c1\u0303(\u03be) ] sin \u03b8 \u2212\u03c1 f [ 1 \u2212 cos \u03c1\u0303(\u03be) ] ; 0 < \u03be \u2264 L (2) and the same quantity for the tool active flank, part III, is pe(\u03be, \u03b8) = [ Xp \u00b1 R1 sin \u03be\u0303(\u03be) ] cos \u03b8 [ Xp \u00b1 R1 sin \u03be\u0303(\u03be) ] sin \u03b8 Zp \u2212 R1 cos \u03be\u0303(\u03be) ; \u03be > L (3) where the linear functions \u03c1\u0303(\u03be) and \u03be\u0303(\u03be) have the following expressions \u03c1\u0303(\u03be) = \u03be \u03c1 f , \u03be\u0303(\u03be) = \u03bb f \u2212 \u03be \u2212 L R1 (4) With reference to the machine set-up shown in Fig", + " By employing the definition of the rotation operator recalled in (9) and the components given in (12), the parametric equations we(\u03c6) and qe(\u03c6) of vectors in (16) are easily obtained according to (146)\u2013(147) in [14]. The expressions for f and \u03a6 are now pretty straightforward to obtain by employing definitions (14) and the results in (13), (15), and (16). It is worth observing that in the invariant approach all geometrical quantities have a unique representation. Let us define x \u2208 R n to be the vector of the n machinetool settings selected as optimization variables. We do not need to distinguish between parameters related to the geometry of the tool surface like, e.g., Rp and \u03b1p in Fig. 1, and those related to the machine kinematics like, e.g., \u2206XD and 2C in Fig. 2, to mention just a few. The vector of the basic (initial) machine-tool settings is defined as x(0) and its current value at iteration k is x(k). To compact the notation, let us introduce the vector of the Gaussian and motion parameters \u03b6 = (\u03be, \u03b8, \u03c6) and its discrete version \u03b6 i = (\u03bei, \u03b8i, \u03c6i) representing their values at the i\u2212th point Pi of an m-point grid sampled on the tooth surface with settings x. Similarly, we define the i\u2212th triplet on the basic tooth surface as \u03b6(0) i (i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000879_aupec.2007.4548084-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000879_aupec.2007.4548084-Figure2-1.png", + "caption": "Fig. 2: NFO space vectors and frame alignment.", + "texts": [ + " e\u03b1 = v\u03b1 \u2212Rsi\u03b1 e\u03b2 = v\u03b2 \u2212Rsi\u03b2 esq = \u2212e\u03b1 sin \u03b8\u0302ms + e\u03b2 cos \u03b8\u0302ms \u03c9\u0302ms = esq |\u03c8\u2217 s | = esq Lm|i\u2217ms| \u03b8\u0302ms = \u222b \u03c9\u0302ms dt+ \u03b8ms(t0) i\u2217sq = \u03c4\u2217e |\u03c8\u2217 s | = \u03c4\u2217e Lm|i\u2217ms| i\u2217sd = |i\u2217ms| (1) Besides that the flux integrator is excluded and the problems associated with this integrator are eliminated, some interesting properties result from the above simplification. For example, if a disturbance occurs resulting in the assumed position of the stator flux vector moved away from its actual position, then a natural compensation mechanism inherent to the algorithm would try to compensate for the disturbance. This natural selfalignment property (to which the NFO concept owes its name) can be briefly explained with the help of Fig 2a. \u2022 Assume that initially the control frame (d, q) was aligned with the true (x, y) frame. \u2022 Then due to a disturbance the (d, q) frame moves ahead of the (x, y) frame. \u2022 Stator current refence vectors i\u2217sd, i \u2217 sq are applied in the control frame, which results in the stator current vector i\u2217s . \u2022 It projects on the (x, y) axes so that isx < i\u2217sd. Hence the actual flux magnitude is smaller than its reference value, i.e. \u03c8s < \u03c8\u2217s . \u2022 Accordingly, the stator flux voltage projections are such that esq < esy . \u2022 Now the estimated frame angular velocity \u03c9\u0302ms = esq \u03c8\u2217s is smaller than the true frame angular velocity \u03c9ms = esy \u03c8s . \u2022 As a result, (d, q) rotates slower than (x, y) until (d, q) approaches (x, y). Similarly, with the help of Fig. 2b it can be shown that under certain (and quite typical) regeneration conditions, NFO algorithm would show a tendency to frame alignment instability. In other words, in a situation of an initial disturbance, instead of adjusting the assumed orientation of the stator flux, the algorithm would move it further away from the correct position, eventually losing control over the frame orientation. For more detail an interested reader is referred to [3]. An improved version of NFO control was suggested in [4] that helps overcome the above problem in a major part of regeneration region while retaining the algorithm\u2019s inherent simplicity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002543_icm46511.2021.9385623-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002543_icm46511.2021.9385623-Figure1-1.png", + "caption": "Fig. 1. Prototype of multi functional drill.", + "texts": [ + " In section V, advantages of the proposed device and experimental results are discussed. Finally, this paper is concluded in section VI. In orthopedic surgery, surgical drills are utilized to cut the vertebra. Since the spinal cord passes through the inside of the vertebral column, tasks using surgical drills involve the risk of the spinal cord injury. Orthopedists must have skills to achieve delicate tasks. The multi functional drill is developed to support surgeons who perform cutting tasks. In this section, the mechanism and the functions of a prototype shown in Fig. 1 are explained. The prototype of the multi functional drill is driven by a linear motor with two optical encoders and a rotary motor with a rotary encoder. A schematic drawing of the prototype is shown in Fig. 2. The actuators of the prototype are cased in the housing equipped with two linear joints. The coil housing of the linear motor is mounted on the linear joint 1. The rotary motor is mounted on the linear joint 2. Both linear joints move in parallel to the rotary axis of the rotary motor used to cut environment by drill bits" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002943_j.ast.2021.106863-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002943_j.ast.2021.106863-Figure4-1.png", + "caption": "Fig. 4. Trunk Sections (Bottom View). Fig. 5. Hard Surface Clearance.", + "texts": [ + " Expressions of the forces and torques in taxiing phase are given by: (1) Propeller forces and torques The propeller forces F T in frame {B} are given by: F T = [ T L + T R 0 0 ]T (1) Where, T L and T R are thrusts of the left and right propellers. Denote the thrust error as: T = T R \u2212 T L . The propeller torques M T in frame {B} are given by: M T = \u23a1 \u23a3 MT x MT y MT z \u23a4 \u23a6 = \u23a1 \u23a3 0 \u2212 (T L + T R) \u00b7 l yz T \u00b7 lxz \u23a4 \u23a6 (2) Where, lxz and l yz are axial components of the distance between the thrust functionary line and the mass point in the frame {B}, respectively. (2) The ground reaction force The bottom of the trunk is a cross-shaped sectional structure, and the air cushion cavity is divided into four parts (the red numbers) as shown in Fig. 4. Motivated by [38], to facilitate the analysis, the trunk is divided into eight sections (the blue numbers) as shown in Fig. 4. Each section is divided into many segments: M segments per straight section and N segments per curved section (Table 1). Thus the total number of trunk segments are 4 (M + N). In Fig. 5, x, y, z are axial distances between the gravitational centers of the airplane and the cushion in the frame {B}. In Fig. 6, xchi, ychi are axial longitudinal distances between the centers of pressure and cushion in the ith segment. xtki , ytki are axial lateral distances between the centers of trunk contact pressure and cushion in the ith segment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002560_tec.2021.3069096-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002560_tec.2021.3069096-Figure5-1.png", + "caption": "Fig. 5 Flux path of M2 Machine with 18/16 combination. (a) \ud835\udc3c\ud835\udc5e excitation (b) \ud835\udc3c\ud835\udc51 excitation.", + "texts": [ + " However, due to the large rotor pole number and the narrow pole tooth width, the stator flux leaks through the adjacent teeth. The overall flux linkage under d-axis current excitation is almost the same as the situation in Fig. 4(a). Hence the rotor end saliency is not apparently reflected as the stator winding flux linkage. Consequently, the saliency modulation signal strength will be very weak due to the large pole number of the rotor end. 2) M2 Machine The M2 machine has fractional slot number. The magnetic paths of the stator excited flux are unevenly distributed in the rotor as shown in Fig. 5 and Fig. 6. In Fig. 5, the 18 slot distributed winding is configured with 8 magnetic poles and the number of slot per pole per phase \ud835\udc5e is 0.75. It is observed that in Fig. 5(a), the \ud835\udc3c\ud835\udc5e excitation results in the maximum stator inductance. The flux path is not distributed as ideal as the previous case where the slot number is an integer. Some rotor teeth are fully located under the stator teeth carrying flux flow. However, some partially lie under the stator teeth such that the flux flow is influenced. In the case that the stator flux is excited with \ud835\udc3c\ud835\udc51 in Fig. 5(b), although some flux paths have high reluctance since the rotor slots block the flux, the other paths still have low reluctance with rotor teeth located exactly under the stator teeth. As a result, the rotor structure saliency is not apparently reflected as stator winding inductance saliency due to the uneven distribution of the flux loops. Moreover, the pole number of the rotor end is greater than that of the machine PM rotor such that the position signal strength caused by the SMRB is weak. Authorized licensed use limited to: Robert Gordon University" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002838_012021-Figure18-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002838_012021-Figure18-1.png", + "caption": "Figure 18. Total Deformation in E-Glass", + "texts": [ + " Stress distribution in S-Glass is 59.887 MPa and 0.3646 MPa respectively shown in Figure 16. 3.4.3. Strain Distribution The Max. and Min. Strain distribution in S-Glass is 0.00069941 and 0.000005478 respectively shown in Figure 17. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.5. Analysing Testing Result of E-Glass 3.5.1. Total Deformation The Max. and Min. Total Deformation in E-Glass is 0.23872 mm and 0 mm respectively shown in Figure 18. The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. 3.5.2. Stress Distribution The Max. and Min. Stress Distribution in E-Glass is 62.053 MPa and 0.37676 MPa respectively shown in Figure 19. 3.5.3. Strain Distribution The Max. And Min. Strain Distribution in E-Glass is 0.0008937 and 0.0000069443 respectively shown in Figure 20. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002020_acemp.2007.4510512-Figure22-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002020_acemp.2007.4510512-Figure22-1.png", + "caption": "Fig. 22 Distribution of magnetic field strength.", + "texts": [], + "surrounding_texts": [ + "3 Magnetic Characteristic Analysis of Three-Phase Transformer Figure 13 shows the analysis model of three-phase 5- legs construct ure transformer. The used material is the grain-oriented electrical steel sheet. Figure 14 shows the distribution of magnetic flux density. Furthermore, the inclination angle B from rolling direction of the vector B is shown in Fig. 15 in order to evaluate the structure of t he three-phase 5 legs typed transformer.", + "Figure 16 shows the distribution of distortion factor of magnetic flux waveform. The eddy current generates by distorting the magnetic flux waveform. The evaluation of waveform distortion gives large effect for the dynamic analysis. Furthermore, the magnetic power loss under the rotational flux is larger than under alternating flux. Figure 17 and 18 show the distribution of magnetic power loss by the dynamic E&S model and by the static E&S model, respectively. By using E&S model, magnetic power loss can be analysed directly. The comparison with the result by the conventional method (static E&S model) was examined in order to investigate features of the result by the dynamic analysis with dynamic E&S model, as shown in Fig. 19. The static E&S model has greatly evaluated the loss from actual result, since the effect of the distorted waveform is disregarded. It is found that magnetic field is suppressed by the effect of the eddy current. The static analysis excessively analyses the magnetic field.\n4 Magnetic Characteristic Analysis of Permanent Magnet Motor Figure 20 shows the concentrated flux typed surface permanent magnet motor (CSPM motor), which was developed for high density machine by our group and have been obtained by Oita University as the patent. The gap field of this motor has the 1.5 times of he magnetization of permanent magnet.\nFigure 21 shows the magnetic flux distribution. From this result it was found that the magnetic path becomes short, therefore the inductance increases as this motor.", + "Figs. 22 and 23 show the distribution of the magnetic field strength and the magnetic flux density, respectively. As same as procedure of the vector magnetic characteristic analysis, the distribution of magnetic power loss is obtained as shown in Fig. 24. The magnetic power loss increases in the place that both vectors B and H are large and the axis ratio in rotational magnetic flux is large. It is found that the major magnetic power loss is being generated at the tooth division of the stator core.\n5 Conclusions\nIn this paper we proposed the dynamic integrated typed E&S model for analyzing the eddy current effect under distorted magnetic flux conditions. And furthermore it was clarified how the effect of eddy current influence the vector magnetic characteristic.\nThe usefulness of dynamic vector magnetic characteristic analysis by using dynamic E&S model is shown in the following. (1) This method can analyse the effect of the eddy current which arises by the distorted flux waveform. (2) This method can consider the dynamic vector magnetic property and can analyse it. (3) The analysis of magnetic power loss (iron loss) can be directly carried out the magnetic characteristic analysis of the electrical machines.\nREFERENCES [1] H. Shimoji, M. Enokizono,\u201dE&S2 Model for Vector Magnetic Hysteresis Property\u201d, Journal of Magnetism and Magnetic Materials, 254, 290-292, 2003 [2] M. Enokizono, H. Shimoji, T. Horibe, \u201dLoss Evaluation of Induction Motor by using Hysteresis E&S2 Model\u201d, IEEE Transactions on Magnetics, Vol.38, No.5, pp.23792381,(2002) [3] S. Urata, H. Shimoji, T. Todaka, M. Enokizono \u201cMeasurement of two-dimensional vector magnetic properties on frequency dependence of electrical steel sheet\u201d, International Journal of Applied Electromagnetics and Machanics, 20, pp.155-162, 2004. [4] M. Enokizono, S. Urata, \u201cDynamic Vector MagnetoHysteretic E&S Model and Magnetic Characteristic Analysis\u201d, COMPUMAG 07, Aachen in Germany, 2007" + ] + }, + { + "image_filename": "designv11_83_0000438_haptic.2006.1627118-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000438_haptic.2006.1627118-Figure8-1.png", + "caption": "Figure 8: A cross-section of a contact region between a rigid object and a face of an elastic object", + "texts": [ + "4, the deformed S inside the local neighborhood of contact is convex when viewed from A. Thus, as the two objects cannot penetrate each other, the boundary features of A will hinder the deformation of S, and the neighboring points of the original contact point on S will begin to contact A. As the result, the original contact point grows to be a contact region along the boundary features of A under heavier deformation. In other words, the contact region of S has the shape formed by the contacting portions of the boundary features of A. An example is shown in Fig. 8. Theorem 3.8 When a rigid convex polyhedral object A contacts a face of an elastic convex object B that is originally undeformed, only one contact region between A and B can be formed. Proof: From Corollary 3.1, only one contact point can be formed between A and a face FB of B before B deforms. From Theorem 3.4, the local deformation of FB is convex. From Theorem 3.6, the portion of FB outside the local neighborhood of contact remains convex. Thus, no other contact region is possible between A and B based on Corollary 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002434_978-3-030-59864-8_13-Figure8.1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002434_978-3-030-59864-8_13-Figure8.1-1.png", + "caption": "Fig. 8.1 Cross section of graphite resistance furnace", + "texts": [ + " Massey 49 The resistance through the graphite heating element induced a radiative heating mechanism. The heating element was contained by multi-layer shielding directing the radiative heating circumferentially inward. Graphite heating elements that are manufactured from a high purity carbon composite providing excellent temperature uniformity and repeatability were employed. This heater design used rounded edges and well-developed gap spacing to maximize uniformity throughout the length of the \u201chot zone\u201d. A cross section of the furnace is shown in Fig.\u00a08.1. The thermocouple housing design varied based on the test objective. In the early stages of this effort, there were many holes for type C thermocouples to maximize the data collected per run. A large thermocouple housing mass induced a significant strain on the system, which outweighed the amount of data collected. As this effort matured, the housing size was decreased to improve efficiency and standardize the calibration process. The standardization will be discussed in the next section. The final thermocouple housing consisted of five holes: three holes for the calibration of the type C thermocouple batch, one hole for the type K reference thermocouple, and one hole for the pyrometer blackbody" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003407_j.optlastec.2021.107425-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003407_j.optlastec.2021.107425-Figure8-1.png", + "caption": "Fig. 8. Temperature field change cloud charts of surface sintering area of WSPC powder (a) t = 0 \u03bcs (b) t = 20 \u03bcs (c) t = 1200 \u03bcs (d) t = 1810 \u03bcs (e) t = 2410 \u03bcs (f) t = 3620 \u03bcs.", + "texts": [ + " 7 (b), because the WSPC single-layer parts are formed after laser multichannel sintering, heat accumulation occurs in the depth of the sintering and formed heat-affected zone material adheres to the bottom of the sintering pool, the thickness (sintering depth) of WSPC single-layer parts is greater than the simulated value, but the two values are basically consistent, with a deviation within 5%. Therefore, the structure size of sintering pool is reasonable in numerical simulation. Y. Yu et al. Optics and Laser Technology 144 (2021) 107425 Y. Yu et al. Optics and Laser Technology 144 (2021) 107425 Fig. 8 shows the cloud charts of temperature field change in surface sintering area of WSPC powder under different laser radiation time when the laser power is 10 W, the preheating temperature is 80\u2103, the scanning speed is 2 m/s, and the scan spacing is 0.2 mm. When time t is 0 \u00b5s, the laser heat source does not act on the powder bed, thus the temperature of the powder bed is equal to the preheating temperature of 80\u2103. When time t is 20 \u00b5s, the laser heat source acts on the powder bed and begins to move along the sintering direction, the surface temperature of the powder bed continues to rise with the increase of laser radiation time, and the sintering pool translates along the direction of laser sintering" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003453_s10846-021-01454-7-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003453_s10846-021-01454-7-Figure5-1.png", + "caption": "Fig. 5 Differential Drive kinematics of the swimming robot", + "texts": [ + " To achieve steering, a differential drive mechanism of twowheel mobile robot, has been adopted for this purpose [32]. The pectoral fins are grouped on a common axis, and each fin can separately be driven either forward or backward. The velocity of each fin can be tuned in an independent manner. To perform the turning motion, the differential drive principle of two-wheel mobile robot is adopted. The robot rotates about a specific point. This point is located along its common left and right fin axis and it is defined as the Instantaneous Center of Curvature (ICC) as shown in Fig. 5. By varying the velocities of the two fins, the trajectories of the swimming robot will be changing. Due to the velocity rate of rotation wt about the ICC must be the same for both fins, so left and right velocities can be obtained as given: Vr \u00bc wt R\u00fe l 2 \u00f02\u00de Vl \u00bc wt R\u2212l 2 \u00f03\u00de Where l can be defined as the distance that lies between the centers of the two fins, Vr, Vl are the right and left fin velocities. Whereas R is the distance from the ICC calculated point to the midpoint between the fins" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001316_j.etap.2007.05.005-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001316_j.etap.2007.05.005-Figure1-1.png", + "caption": "Fig. 1. Metabolic activation of tetryl.", + "texts": [ + "edu (S.R. Myers). t t 1 N c N q d S 382-6689/$ \u2013 see front matter \u00a9 2007 Elsevier B.V. All rights reserved. oi:10.1016/j.etap.2007.05.005 oh, 1984). A mixture of nitric and sulfuric acids is then used o nitrate the compound to form 2,4-dinitromethylaniline and rude tetryl. Crude tetryl is then filtered, washed with water, and issolved in acetone, which is evaporated and filtered to yield ure tetryl. .2. Metabolism Recent metabolic experiments in vitro demonstrate that etryl undergoes N-denitration (Fig. 1) to form N-methyl-2,4,6rinitroaniline (N-methylpicramide, NMPA) (Anusevicius et al., 998). Miskiniene et al. (1998) found that the flavoenzyme, ADPH: thioredoxin reductase (TR) of Arabidopsis thaliana, aused redox cycling of tetryl and parallel nitrite elimination and MPA formation under aerobic conditions. Rat liver NAD(P)H: uinone oxidoreductase was also found to catalyze the reductive enitration of tetryl to form NMPA (Anusevicius et al., 1998; hah and Spain, 1996). The reaction was believed to proceed S" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002167_jae-201576-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002167_jae-201576-Figure2-1.png", + "caption": "Fig. 2. Magnetic gear with 2.667 gear ratio.", + "texts": [ + " Some MGs even introduced the armature coil in their design to produce a magnetic gear effect\u00a0[31,32]. K. Attalah\u00a0[7] configuration will be used in this paper to study the gear efficiency and modeling. Although surface-mount PM rotor is not robust in a high-speed application, the structure produces the highest torque density and magnetic field utilization among other MG structure, which is essential in eddy current loss study. This study assumes that the PM is firmly attached to the rotor surface. Figure\u00a02 illustrates the magnetic gear having 6 high-speed inner rotor pole pairs, pi, 16 lowspeed outer rotor pole pairs, np and 22 pole pieces placed between the rotor, po. Seven magnetic gears with different gear ratios were designed according to the application note from the JMAG software\u00a0[33]. The magnet used in this simulation is Neomax 35AH, while the pole piece, inner rotor yoke and outer rotor yoke use NSSMC 35H210 soft magnetic material. The overall size of the structure and the number of inner pole pair were fixed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000368_detc2004-57064-Figure9-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000368_detc2004-57064-Figure9-1.png", + "caption": "Figure 9. Kim and Tsai\u2019s Modified cartesian Manipulator.", + "texts": [ + "org/about-asme/terms-of-use Do mobility of the parallel manipulator is given by equation (8), F r \u2211 i 1 fi k \u2211 j 1 dim Vm f j dim Am f a 12 3 \u2211 j 1 4 3 3 Furthermore, there is no passive degrees of freedom since Fp \u2211 fi k \u2211 j 1 dim Vm f j 12 3 \u2211 j 1 4 0 It should be noted that Kim and Tsai, [5], after applying the mobility criterion based on the Kutzbach-Gru\u0308bler criterion, equation (1), found that this parallel manipulator has zero degrees of freedom. 5.4 Kim and Tsai\u2019s Modified Cartesian Manipulador. Consider the cartesian parallel manipulator, similar to that proposed by Kim y Tsai [5], shown in Figure 9. The manipulator is composed of three serial connecting chains. The first two serial connecting chains are formed by a prismatic pair and three revolute pairs, the direction of the prismatic pair and the axes of the three revolute pairs are parallel and the kinematic pairs generate the subalgebra associated with the Scho\u0308nflies group RP 4 e\u0302 j , the third chain is formed by a prismatic pair and only two revolute pairs, the direction of the prismatic pair and the axes of the revolute pairs are parallel" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003002_09544070211024109-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003002_09544070211024109-Figure7-1.png", + "caption": "Figure 7. Detailed drawing of the moving and the stationary parts of the friction tester.", + "texts": [ + " Depending on the test conditions, lubricant or moisture was applied to the surface and then the specimens were pressed against each other with the defined normal force (2, 5, and 10N). Then the driving unit (moving part) slides at the test speed (50, 100, and 200mm/s) to produce squeak noise from the pair of specimens in contact. To perform tests using a numerous number of test specimens designed as above, it is necessary to attach and detach the specimens quickly and efficiently. To this end, a method was devised to attach the specimen easily on the friction tester. As depicted in Figure 7, the specimen holder applied to the moving part is designed to be detachable. The specimen holder can be removed from the friction tester to replace the test specimens and can easily be put back together with the tester by turning the knob on the side. The surface of the specimen holder was made to have the curvature of 200mm in radius. It is to prevent the test specimens from falling off during friction test. To the stationary part, an acrylic plate of 3mm thick, 70mm wide, and 150mm long was made to attach the test specimen, and a frame to fix the plate was installed on the tester stage" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002838_012021-Figure28-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002838_012021-Figure28-1.png", + "caption": "Figure 28. Stress Distribution in HSLA Steel", + "texts": [ + " Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.8. Analysing Testing Result of HSLA Steel 3.8.1. Total Deformation The Max. And Min. Total Deformation in HSLA Steel is 0.19371 mm and 0 mm respectively shown in Figure 27. The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. 3.8.2. Stress Distribution The Max. And Min. Stress Distribution in HSLA Steel is 181.96 MPa and 1.2159 MPa respectively shown in Figure 28. 3.8.3. Strain Distribution The Max. And Min. Strain Distribution in HSLA Steel is 0.0011085 and 0.00000815 respectively shown in Figure 29. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 The Analysis report of maximum total deformation, maximum equivalent stress and maximum equivalent strain is presented below in bar graphs: 4.1. Maximum Total Deformation: Maximum Total Deformation is shown in carbon fiber i.e" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000847_s10409-008-0179-5-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000847_s10409-008-0179-5-Figure8-1.png", + "caption": "Fig. 8 The configuration of the internal force in the Kagome cell", + "texts": [ + " The cell wall deformation behavior is also stretching dominated under either uniaxial or shear loading. 3.1 Mechanical analysis of the triangular cell For Kagome lattice under different kinds of in-plane loadings, the analysis of the mechanical properties is almost same with the full triangular lattice. The Kagome cell under general in-plane loading is shown in Fig. 7. We also consider the case of biaxial stresses \u03c3x and \u03c3y . The symmetry of force in the y directions requires: M1 = \u2212M \u2032 1, Q3 = 0, M \u2032 3 = M3D. (13) Equivalence of forces in the lattice strut requires (shown in Fig. 8): Q3l + 2Q1l + M3D + M1 = M \u2032 1 + M \u2032 3. (14) Equivalence of external forces with the stress \u03c3x and \u03c3y in the lattice cell requires: \u221a 3N1 + Q1 = 2\u03c3ybl, N3 + N1 \u2212 \u221a 3Q1 2 = \u221a 3\u03c3x bl. (15) Noting that the relative level deformation \u03b4Dh and the relative rotation \u03b8D in the middle point \u201cD\u201d of the level strut are equal to zero, the additional condition can be acquired, \u03b4Dh = 0, \u03b8D = 0. From Eqs. (13) to (15) and \u03b4Dh = 0, \u03b8D = 0, the internal forces of each strut can be obtained, i.e. N1 = \u221a 3 3 \u03c12\u03c3x + (12 + \u03c12)\u03c3y 6 + \u03c12 bl, N3 = \u221a 3 3 (18 + \u03c12)\u03c3x \u2212 (6 \u2212 \u03c12)\u03c3y 6 + \u03c12 bl, Q1 = (\u2212\u03c3x + \u03c3y)\u03c1 2 6 + \u03c12 bl, M1 = \u2212M \u2032 1 = 2 3 (\u03c3x \u2212 \u03c3y)\u03c1 2 6 + \u03c12 bl2, M3D = \u2212M \u2032 3 = 1 3 (\u03c3x \u2212 \u03c3y)\u03c1 2 6 + \u03c12 bl2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003270_s41403-021-00254-7-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003270_s41403-021-00254-7-Figure4-1.png", + "caption": "Fig. 4 Schematic illustration of Ti6Al4V-ELI 3D printed\u00a0test coupon layout (Note:\u00a0Notches were machined later, i.e. after stress-relieving heat treatment)", + "texts": [ + " The mechanical properties of 3D printed components are dependent on the surface roughness and porosity which are significantly affected by the process parameters of the LPBF process. A manufacturing plan that is likely to ensure that parts and specimens printed with the same process parameters will result in repeatable mechanical properties. Test specimens printed along with the parts are oriented in such a way to ensure that tensile properties in any arbitrary orientation in the part are captured and reported. The orientations of the test coupons 3D printed along with the components are schematically shown in Fig.\u00a04. The hardness of specimens\u2019 representative of each direction is shown in Table\u00a02. The mechanical properties of specimens in different orientations are shown in Table\u00a03. Mechanical properties have been evaluated on three samples representative of each orientation. The representative engineering stress\u2013engineering strain curves are shown in Fig.\u00a05. It was found that the hardness values were fairly consistent and the values were slightly higher in Z and 45\u00b0 directions as observed in Table\u00a02. The tensile properties (0", + " The standard deviation in 0.2% proof stress values across all four orientations is 65\u00a0MPa. 0.2% proof stress and UTS were found to be comparatively higher in vertical and inclined directions than the horizontal build direction. It is also observed that the increased strength has not affected the % elongation and %RA of these specimens. It may be noted that in case of impact test specimens in X, Y orientations, the notch is made perpendicular to the cross section of layers one above the other as shown in Fig.\u00a04, the crack can easily propagate between any two layers leading to a layer separation during loading. The energy required for this case would hence be less than that in case of a specimens in which the notch is made cutting across many printed layers as shown in Fig.\u00a04. Thus, the crack has to propagate across many printed layers below it as in case of 45\u00b0 orientations. For Z direction, since the notch is made parallel to the 3D printed layers, it can lead to easy crack propagation between any 2 layers and thus would expect to exhibit lower impact strength compare to any other orientations shown in this study. For fracture toughness testing, the notch has been created in such a way that the loading will happen in the orientation in which the specimen has been printed. The notch and specimen orientation are shown with respect to the printed layers in Fig.\u00a04. Results of Charpy impact test and fracture toughness evaluation using KIc tests show that the material possesses good toughness. Similar to the trend seen in strength, vertical and inclined specimens show better toughness than horizontal specimens. It may be noted that in case of fracture toughness test specimens in X, Y orientations, the notch is made perpendicular to the cross section of layers one above the other as shown in Fig.\u00a04, the crack cannot easily propagate across layers. The same may be compared to the crack propagation in Z direction Fig.\u00a04 where the crack can propagate in between any two layers. The crack path in case of specimens in 45\u00b0 orientation as shown in Fig.\u00a04 would be more tortuous compared to that in X, Y and Z directions. Karolewska et\u00a0al. (2020) studied the mechanical properties of Ti6Al4V manufactured by LPBF and conventional drawn bars. They showed that the strength for the LPBF samples is about 11% higher than that for drawn bars. Further, the %elongation of LPBF samples is approximately 48% lower than that exhibited by drawn bars. This is due to higher hardness and strength values in LPBF samples compared to the hardness of samples of drawn bar in annealed condition" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001511_speedham.2008.4581320-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001511_speedham.2008.4581320-Figure1-1.png", + "caption": "Fig. 1. Definition of converter input and output electrical quantities.", + "texts": [ + " INTRODUCTION Electrical engineers are used to describe a transformer by its voltage ratio, which can be deduced from the number of turns of the different windings. This represents a strong means in modeling the ideal transformer in a given circuit. If we try to use the same philosophy in describing a static converter behavior, we can introduce a time dependant \u2018transformation\u2019 ratio that has only two possible values \u20180\u2019 or \u20181\u2019 [1,4], linking input and output voltages of the converter, but also the output and input currents of the same static device. Fig. 1 gives an example of a static converter with its defined electrical quantities. The transformation ratio has been called \u2018tsd\u2019, which is an abbreviation of \u2018transparency and separation distribution\u2019 or TRANSEP-distribution [1]. The electrical quantities of Fig.1 are then expressed as in (1) and (2), reminding so the ideal transformer basic expressions, but depending on time. ( ) ( ) ( )a et t tsd tu u (1) This work is a part of a project that was supported by the strategic fund of the University of Applied Sciences of Western Switzerland, as well as by Bombardier Transportation Ltd, Z\u00fcrich, Switzerland. ( ) ( ) ( )e at i t tsd ti (2) Through variation of time and some other parameters, equation (3) generates a programmable switching signal (which is in fact a distribution) able to control any type of static converter", + "( , )sin( ) cos( ( )) n n m tS n k t n (5) The different parameters used above are defined as follows: ML : Number of switching levels per converter branch : Pulsating frequency of the fundamental signal : Impedance angle of the load supplied : Displacement angle between converter phases : Displacement angle of the carrier k : Multiplication factor between the carrier and fundamental frequency n : Summation index m0 : Modulation depth of the output signal DC component m1 : Modulation depth of the output signal fundamental In the example of Fig.1 the number of switching levels per branch \u2018ML\u2019 is equal to 2. The application of (3) with two different values of the displacement angle (0 and respectively) produces the signals illustrated in Fig. 2. Actually, the rectangular signal is the TRANSEPdistribution corresponding to the converter, while the Analytical Model Application on a Converter for Traction Drives J. El Hayek*, G. Skarpetowski**/*** * University of Applied Sciences of Western Switzerland, Fribourg, (Switzerland) ** Cracow university of Technology, Cracow, (Poland) *** Bombardier Transportation Ltd, Z\u00fcrich, (Switzerland) SPEEDAM 2008 International Symposium on Power Electronics, Electrical Drives, Automation and Motion 1205 978-1-4244-1664-6/08/$25" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001095_s10015-008-0539-z-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001095_s10015-008-0539-z-Figure1-1.png", + "caption": "Fig. 1. A schematic diagram of the omni-directional robot", + "texts": [ + " When constructing the time-scaling function, the whole time interval of the reference trajectory is divided into many segments, in which the time-scaling function is interpolated as a cubic spline curve according to the chosen control points. However, it should be noted that there are no ways to determine these control points simultaneously in the previous research. Therefore, a PSO method is adopted to overcome this diffi culty. In this section, it is assumed that the omni-directional mobile robot consists of the orthogonal-wheel assembly mechanism proposed by Pin and Killough.6 A schematic diagram to illustrate the motion of the omni-directional robot is given in Fig. 1, and a world-frame [xw, yw]T and a Key words Omni-directional robot \u00b7 Time-scaling \u00b7 Particle swarm optimization When applying time-scaling techniques to determine the time-optimal movement of a plant along a specifi ed path, the corresponding scenario is usually as follows. In the beginning, a reference trajectory is synthesized to meet geo- moving-frame [xm, ym]T are defi ned as shown in Fig. 2. The former denotes a frame that everything discussed can be referenced and the latter is a frame attached to the center of the gravity of the robot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002772_978-3-030-67750-3_14-Figure30.1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002772_978-3-030-67750-3_14-Figure30.1-1.png", + "caption": "Fig. 30.1 Wheelbarrow applying second class lever", + "texts": [ + "my \u00a9 The Author(s), under exclusive license to Springer Nature Switzerland AG 2021 M. H. Abu Bakar et al. (eds.), Progress in Engineering Technology III, Advanced Structured Materials 148, https://doi.org/10.1007/978-3-030-67750-3_30 357 A wheelbarrow is small transporting hardware and usually with one wheel at the front, two handles for users to hold and lift it, and the load will be in between of the handle and front wheel. The wheelbarrow eases the user by having the load in the middle of the fulcrum and the effort, applying the second-class lever as in Fig. 30.1. Nowadays, there have been numerous designs of wheelbarrow developed, aiming to ease the users and to increase productivity. There is a difference between wheelbarrow used at nursery or being used by gardeners if compared to the one that is used at the construction site. Thewheelbarrow used by the gardener has a smaller and shallower tray bucket to lodge the load while the wheelbarrow used at the construction site has a deeper tray bucket. Meanwhile, wheelbarrows at the construction sites are also varying depending on their application, such aswheelbarrowwith onewheel and two wheels" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000629_9780470612286.ch1-Figure1.5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000629_9780470612286.ch1-Figure1.5-1.png", + "caption": "Figure 1.5. Singular positions of the St\u00e4ubli RX-90 robot", + "texts": [ + " After performing all the calculations, we obtain the following solutions: 1 y x 1 1 atan2(P , P ) ' \u03b8 =\u23a7\u23aa \u23a8 \u03b8 = \u03b8 + \u03c0\u23aa\u23a9 \u03b82 = atan2(S2, C2) with: 2 2 2 2 2 2 2 2 2 2 YZ X X Y ZC2 X Y XZ Y X Y ZS2 X Y \u23a7 \u2212 \u03b5 + \u2212\u23aa = \u23aa + \u23a8 \u23aa \u2212 \u03b5 + \u2212=\u23aa \u23a9 + with \u03b5 = \u00b1 1 X = \u2013 2Pz D3 Y = \u2013 2 B1D3 Z = (RL4)2 \u2013 (D3)2 \u2013 (Pz)2 \u2013 (B1)2 B1 = Px C1 + Py S1 z z 3 P S2 B1C2 D3 B1S2 P C2 atan2 , RL4 RL4 \u2212 \u2212 + \u2212 +\u239b \u239e\u03b8 = \u239c \u239f \u239d \u23a0 4 x y x y z ' 4 4 atan2[S1 a C1 a , C23(C1 a S1 a ) S23 a ]\u03b8 = \u2212 \u2212 + \u2212\u23a7\u23aa \u23a8 \u03b8 = \u03b8 + \u03c0\u23aa\u23a9 \u03b85 = atan2(S5, C5) with: S5 = \u2013C4 [C23 (C1 ax + S1 ay) + S23az] + S4 (S1 ax \u2013 C1 ay) C5 = \u2013S23 (C1 ax + S1 ay) + C23 az \u03b86 = atan2(S6, C6) with: S6 = \u2013 C4 (S1 sx \u2013 C1 sy) \u2013 S4 [C23 (C1 sx + S1 sy) + S23 sz] C6 = \u2013 C4 (S1 nx \u2013 C1 ny) \u2013 S4 [C23 (C1 nx + S1 ny) + S23 nz] NOTES. 1) Singular positions: i) when Px = Py = 0, which corresponds to S23RL4 \u2013 C2D3 = 0, the point O4 is on the axis z0 (Figure 1.5a). The two arguments used for calculating \u03b81 are zero and consequently \u03b81 is not determined. We can give any value to \u03b81, generally the value of the current position, or, according to optimization criteria, such as maximizing the distance from the mechanical limits of the joints. This means that we can always find a solution, but a small change of the desired position might call for a significant variation of \u03b81, which may be impossible to carry out considering the velocity and acceleration limits of the actuators, ii) when C23(C1ax + S1ay) + S23az = 0 and S1ax \u2013 C1ay = 0, the two arguments of the atan2 function used for the calculation of \u03b84 are zero and hence the function is not determined. This configuration happens when axes 4 and 6 are aligned (C\u03b85 = \u00b11) and it is the sum (\u03b84 \u00b1 \u03b86) which can be obtained (see Figure 1.5b). We can give to \u03b84 its current value, then we calculate \u03b86 according to this value. We can also calculate the values of \u03b84 and \u03b86, which move joints 4 and 6 away from their limits, iii) a third singular position occurring when C3 = 0 will be highlighted along with the kinematic model. This singularity does not pose any problem for the inverse geometric model (see Figure 1.5c). 2) Number of solutions: apart from singularities, the St\u00e4ubli RX-90 robot has eight theoretical configurations for the IGM (product of the number of possible solutions on each axis). Some of these configurations may not be accessible due to their joint limits. z2 z1 O4\u2261O6 z3 O2 O3 a) Singularity of the shoulder (Px = Py = 0 and S23RL4 \u2013 C2D3 = 0) 1.3.1. Direct kinematic model The direct kinematic model of a robot gives the velocities of the operational coordinates X in terms of the joint velocities q " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001400_6.2008-7171-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001400_6.2008-7171-Figure7-1.png", + "caption": "Figure 7 X-4 Flyer frame of reference", + "texts": [ + " The simulation results demonstrate that the desired \u201ctriangular\u201d formation (Figure 5) accurately follows the reference path after initialization and stays on the path thereafter. In this research, the developed formation strategies for three air vehicles are extended to the tight formation flight of X4-Flyers unmanned rotorcrafts. The simulation differs from the previous example by considering the actual dynamical behavior of the rotorcraft in a 3D environment adding an altitude separation command to the formation schemes (Figure 6). American Institute of Aeronautics and Astronautics IV. X4-Flyer Mathematical Model Let us define the frames of reference as shown in figure 7: a fixed-to-earth frame G(xG, yG, zG) and a local moving frame attached to the center of mass O of the vehicle, A(xA, yA, zA). G is assumed to be Galilean. Orientation of frame A relative to frame G is expressed with the rotation matrix [RA/G], equation (1). The airframe orientation in space is given by the Euler angles: roll (\u03c6 ), pitch (\u03b8) and yaw (\u03c8) with angular velocities respectively (p, q, r). (1) \u23a5 \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a2 \u23a3 \u23a1 \u2212 \u2212+ +\u2212 = )cos()cos()cos()cos()sin( )cos()sin()sin()sin()cos()sin()cos()sin()sin()sin()cos()sin( )sin()sin()cos()sin()cos()sin()cos()cos()sin()sin()cos()cos( / \u03b8\u03c6\u03b8\u03c6\u03b8 \u03c8\u03c6\u03c8\u03b8\u03c6\u03c8\u03c6\u03c8\u03b8\u03c6\u03b8\u03c8 \u03c8\u03c6\u03c8\u03b8\u03c6\u03c8\u03c6\u03c8\u03b8\u03c6\u03b8\u03c8 GAR The following model represents an ideal case, where all variables are known or measured exactly with no measurement noise and no external disturbances" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001350_imece2008-67917-FigureA.1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001350_imece2008-67917-FigureA.1-1.png", + "caption": "Figure A.1: Fixture for compression test", + "texts": [ + " Additional components may include antioxidants, adhesion agents, flame retardant agents and special process-enhancing chemical additives. Every ingredient of a rubber recipe may affect physical properties, independently or dependently on each other. The mixing and curing process is also critical in determining these properties. Therefore it is necessary to experimentally identify material characteristics of the particular rubber used in this research. Compression tests on the rubber are conducted using the United Test Systems universal testing machine. 12.7 mm diameter specimen was used for compression test as shown in the Figure A.1. True stress \u2013 true strain curve is shown in Figure A.2 and A.3. As expected, rubber experiences different phases during the test. Results show that Young\u2019s modulus of elasticity at the force induced by the tightening torque is 1.70 MPa for 3 mm thick rubber and 2.60 MPa for the 1.5 mm thick rubber respectively. Experimental results show that the material behaves linearly in the neighborhood of these values. Poisson ratio of 0.49 is used in both cases. Copyright \u00a9 2008 by ASME rl=/data/conferences/imece2008/70919/ on 06/30/2017 Terms of Use: http://www" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000903_j.aca.2007.05.014-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000903_j.aca.2007.05.014-Figure3-1.png", + "caption": "Fig. 3. Selection of mode in solvent exchange using hollow fiber. Mode 1: end of fiber open; mode 2: end of fiber closed during injection of sample solution a b 3", + "texts": [ + " Mode selection Three experiments were performed: 100 L analyte solution n carbon tetrachloride (20 g mL\u22121) was injected into (1) a ollow fiber with open end, (2) a hollow fiber with closed end nd finally (3) a hollow fiber with open end which was closed y heat before washing of analytes. In all experiments 100 L ethanol was used to wash the precipitated analytes. In all cases ow rate of sample solution was manually adjusted so that it did ot flow out the fiber. The obtained methanol solutions were njected to HPLC-DAD system and the obtained recoveries are lotted versus mode used in Fig. 3. As can be seen in the case f experiment (1) recoveries are very low whereas in cases of xperiments (2) and (3) almost similar recoveries were obtained. t was noted that in the case of experiment (3), evaporation of ample solvent from hollow fiber was very fast in comparison ith experiment (2). It seems that most of analyte molecules re precipitated in the pores of the hollow fiber, which were ot washed when the end of fiber was open in the washing step ith methanol (experiment 1). But an open end fiber accelerates he exit of sample solvent and its evaporation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002464_s38313-021-0623-5-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002464_s38313-021-0623-5-Figure6-1.png", + "caption": "FIGURE 6 Rotor of the EESM with rotating secondary winding and stationary ferrite core (\u00a9 Mahle)", + "texts": [ + " In the transmitter developed by Mahle, the two halves of the core are not separated, and one stationary core is used instead. It is fixed to the stationary part of the electrical machine, and only the secondary winding can ro - tate freely within the stationary core of the transformer. This design avoids the limitations due to the disadvantageous properties of the sintered powder material (ferrite). The secondary winding\u00a0is designed as a planar winding on a printed circuit board, thus allowing optimal utilization of the installation space, FIGURE 6. A power electronics circuit is used to control the excitation current for both the conventional conductive transmitter and the inductive transmittor. When the conductive system is replaced by an in - ductive transmitter system, the power electronics need to be adapted, as the transformer requires AC voltage. As it is shown schematically in FIGURE 7, a fullbridge push\u2013pull converter (Single Ac - tive Bridge, SAB) is used. On the input side, the SAB has an inverter stage consisting of four power semiconductors" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002639_j.matpr.2021.03.457-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002639_j.matpr.2021.03.457-Figure5-1.png", + "caption": "Fig. 5. (a) Schematic Diagram of the four-bar linkage with counterweight[15]. (b) Model of Optimized linkage using ADAM [15].", + "texts": [ + " Mishra et al.[15] emphasized the work based on the static and dynamic structural analysis. DE algorithm was utilized to get optimum link lengths which could be minimized the gap between the desired and actual path. For analysis, a point load was taken on the coupler. Different types of load and shaky moments are designed and investigated. The counterweight was provided to avoid the deflection. The combined effect optimization and counterweight were analyzed for path generation and is as shown in Fig. 5(a and b) and Fig. 6(a and b) [15]. Tari et al. [16] designed a slider-crank four-bar linkage for achieving the predefined target points and analyzed the data by considering 558 slider-crank four bars in similar pairs over any eight specified target points. To reduce the formulation of a multi-degree polynomials system, classical elimination methods are used. A forced homotopy method works to remove corrupt results, planning them to solve at infiniteness, which circumvents tedious post-processing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000428_j.ppnp.2004.12.001-Figure18-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000428_j.ppnp.2004.12.001-Figure18-1.png", + "caption": "Fig. 18. Definition of the angles for the \u03c7c angular distributions.", + "texts": [ + " \u2013 The fractional electric octupole amplitude, a3 E3/E1, contributes only to the radiative decay of the \u03c7c2 and is predicted to vanish in the single quark radiation model [77]. If the photon is emitted by a single quark, the octupole amplitude vanishes as a consequence of angular momentum conservation, if the J/\u03c8 is pure S-wave. A small D-wave component would make a3 non-zero, but still negligibly small [76,78]. \u2013 Finally B2 0 (\u03c7c2) is the helicity 0 amplitude in the \u03c7c2 formation process. The angular distributions of reaction (39) can be expressed as a function of three angles (see Fig. 18): the polar angle \u03b8 of the J/\u03c8 with respect to the direction of the antiproton, in the p p center-of-mass system; the polar angle \u03b8 \u2032 of the positron in the J/\u03c8 rest frame with respect to the direction of the J/\u03c8 in the \u03c7c center-of-mass system; the azimuthal angle \u03c6\u2032 between the J/\u03c8 decay plane and the \u03c7c decay plane. Using the helicity formalism the angular distributions of reaction (39) can be expressed as functions of the measured angles \u03b8 , \u03b8 \u2032 and \u03c6\u2032, with coefficients which depend on the helicity amplitudes B and the multipole transition amplitudes a [79]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002682_mias.2021.3063090-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002682_mias.2021.3063090-Figure2-1.png", + "caption": "FIGURE 2. Two versions of the DCVT: (a) a version for a voltage of 6 (10) kV and (b) a version for a voltage of 110 kV.", + "texts": [], + "surrounding_texts": [ + "The improved accuracy of the invented methods of RFL is primarily due to the use of innovative digital current transformer (DCT) and digital voltage transformer (DVT) measurements as information signals [19]\u2013[23], having the following advantages: a high measurement accuracy (error of not more than 0.1% on current and voltage); a large frequency measurement range, including dc measurement; and no saturation from aperiodic component of current and SC currents. Nontraditional current and voltage sensors [small-sized current transformer (CT), dc sensor, Rogovsky coil, and resistive voltage divider; see Figures 1 and 2] are used in the DCTs and DVTs produced by ISPU together with Research and Production Association (RPA), LLC, digital measuring transformers, which have greater accuracy in a wide frequency range compared to traditional current and voltage transformers (Figures 3\u20136). Currently, DCTs and DVTs have been developed for voltage classes 6 (10), 35, and 110 kV. The cost of a threephase set of digital transformers (DTs) (for the specified voltage classes) is approximately equal to the cost of a three-phase set of electromagnetic measuring transformers with a device for converting analog signals (merging unit). The use of DCTs and DVTs for RFL tasks is promising due to the high accuracy of converting and digitizing signals of primary quantities. The application of DCTs and DVTs for RFL purposes on power lines with single-end supply was considered in [17]. The concept of an information system installed on the anchor supports of power lines (transmission lines) Authorized licensed use limited to: University of Exeter. Downloaded on May 27,2021 at 19:33:21 UTC from IEEE Xplore. Restrictions apply. IEEE Industry Applications Magazine JULY/AUGUST 20214 without additional mounting structures is presented in Figure 7. In [18], the results of studies of Rogowsky coils are presented showing that the usage of these coils for measurment of the derivative of current does not introduce errors in the RFL methods and, in some cases, significantly increases the accuracy of measurement Authorized licensed use limited to: University of Exeter. Downloaded on May 27,2021 at 19:33:21 UTC from IEEE Xplore. Restrictions apply. JULY/AUGUST 2021 IEEE Industry Applications Magazine 5 due to the absence of possible methodological errors in the mathematical calculation of the derivative. At the moment, the information system device is in trial operation. In [21]\u2013[23], the RFL method is offered based on synchronized two-side EMEV measurement with the use of DCTs and DVTs as a part of an MFS. An MFS is a set of devices for a digital substation consisting of \u25cf digital measuring transformers (located on an open switchgear) \u25cf digital circuit boards of DCTs and DVTs containing, in addition to measurements, diagnostic and switch control functions (located on an open switchgear) \u25cf modular relay protection and automation device (located at the operational control point; see Figure 8). The prototype device of the MFS is manufactured by ISPU in collaboration with industrial partners\u2014RPA LLC digital measuring transformers and joint-stock company Ivelectronaladka (Figure 8). Because in this article we consider only the RFL function as part of the MFS, we use the term \u201cprototype RFL device.\u201d The RFL method involves the use of synchronized measurements (emergency and preemergency) on both sides of a high-voltage overhead line to calculate the fault locations and updated line parameters (for example, [16]). In this article, the study of the RFL algorithm based on two-sided measurement includes the following: \u25cf the presentation of the RFL algorithm itself and its application features in conjunction with digital current and voltage transformers \u25cf the results of a preliminary study of the RFL algorithm on simplified models in Simulink \u25cf a general methodology for studying a mockup prototype of an RFL on RTDS and criteria for evaluating the effectiveness of the algorithm \u25cf the results of the study of a mockup prototype of an RFL device on a real network model in RTDS." + ] + }, + { + "image_filename": "designv11_83_0002545_tcst.2021.3064801-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002545_tcst.2021.3064801-Figure8-1.png", + "caption": "Fig. 8. Skill weighted action integration in the target subgoal vicinity. (a) Robot eventually reaches the subgoal. (b) Robot goes through the subgoal vicinity without visiting the subgoal.", + "texts": [ + " We define a method for interpreting and encoding such operator adjustment actions via an SWAI defined by \u03c1 = \u03b2 \u222b t f t0 a(t)1( \u2229 |X = G) dt (16) where t0 and t f , respectively, denote times at which the operator enters and exits the AEHR, \u03b2 is a positive design scaling parameter (controls how much the designer wants to scale the integration results based on the knowledge of the skill level of the operator), \u03c1 \u2208 R m is the vector denoting the adjustment for the target subgoal, a(t) \u2208 R m is the action vector, and the indicator function 1( \u2229 |X = G) is given by 1( \u2229 |X = G)= { 1, X \u2208 ( \u2229 ) given X = G 0, otherwise. (17) The indicator function provides a mechanism for SWAI to encode subgoal adjustment information only when the robot is within the intersection of and , and it has not reached the target subgoal. Fig. 8 shows an illustration of such conditions. Fig. 8(a) illustrates this concept by showing that, in , the robot trajectory \u03c4 is perturbed by operator\u2019s subgoal adjustment actions a(t), a(t + t\u03b5). The closed-loop subgoal (G) tracking actions are denoted by aCL (t), aCL (t + t\u03b5), and aCL (t + 2t\u03b5), and the positions of the robot are denoted by X (t), X (t + t\u03b5), and X (t +2t\u03b5). Notice that, at the time t +2t\u03b5, the operator may be satisfied with the adjustment and might not take any additional action; thus, a(t + 2t\u03b5) = 0 is not shown in the figure. Fig. 8(b) illustrates another scenario where the SWAI finishes encoding adjustment information since the robot has left the AEHR without reaching the target subgoal. Once the robot either reaches the subgoal or leaves the AEHR, as shown in Fig. 8, we update the last visited subgoal L to the previous prediction result G, which is the subgoal \u03bbk in the subgoal set \u03bb. Then, adjustment to this subgoal is made in the m-dimensional robot joint space by \u03bb+ k = \u03bb\u2212 k + \u03c1 (18) where \u03bb+ k denotes the adjusted position and \u03bb\u2212 k denotes its previous/nominal position. Authorized licensed use limited to: East Carolina University. Downloaded on June 29,2021 at 06:02:49 UTC from IEEE Xplore. Restrictions apply. Algorithm 2 BSC With Subgoal Adjustment A real-time implementation for BSC with subgoal adjustment, as discussed in this section, is provided in Algorithm 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001385_ijeee.45.2.8-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001385_ijeee.45.2.8-Figure5-1.png", + "caption": "Fig. 5 EMKellner with two rotatable arms.", + "texts": [], + "surrounding_texts": [ + "Some results of earlier PEMs are now presented as examples of what can be achieved. EMKellner: Development of a device for the electromechanical pouring of wheat beer This project, run during the winter course in 2005/2006, had a enormous response in terms of publicity generated on radio and on TV in Germany. Five groups of students had to develop a device for the electromechanical pouring of wheat beer. Wheat beer is a German speciality which foams extremely during pouring. The requirements placed on the device were as follows: the use of a standard commercial bottle and glass with a volume of 500 ml; power consumption of less than 48 W; and a maximum footprint of a DIN A4 page. Additionally, it was not allowed to put any item into the bottleneck. Finally, the beer had to be fi t for consumption and completely fi ll the glass. Four out of the fi ve groups elected to imitate the movement of pouring with a tipping of the bottle relative to the glass. This resulted from many preliminary tests. The realisations were individual, however. On 14 February 2006 the students presented the results of their development. The simplest solution realised the pouring with a fi xed angle between glass and bottle. The realisation of the second group was more complex. Three motors were used to imitate the corresponding human movement. With the help of a microcontroller it was also possible to shake out the rest of the dregs from the bottom of the bottle. Another device realised a parallel at UNIVERSITE LAVAL on July 2, 2015ije.sagepub.comDownloaded from International Journal of Electrical Engineering Education 45/2 at UNIVERSITE LAVAL on July 2, 2015ije.sagepub.comDownloaded from International Journal of Electrical Engineering Education 45/2 at UNIVERSITE LAVAL on July 2, 2015ije.sagepub.comDownloaded from International Journal of Electrical Engineering Education 45/2 In terms of technical development, the focus of this seminar lies in the teaching methodology. Creating the operating instructions was also part of this seminar. The experiment is divided into three parts. First, a theoretical overview of pressure measurement is given. This includes an introduction to common measuring principles, the evaluation of measurement failures and the fi eld of application of different sensor types. The second part contains the static pressure measurement. The sensor characteristics introduced in the fi rst part have to be verifi ed. To this end the measuring results of different sensor types are compared with the results of a reference sensor. A self-developed heating sensor also enables the investigation of temperature as a disturbance variable. The last part of the experiment is the dynamic pressure measurement. The students have to obtain the amplitude response of a differential pressure sensor. An electrodynamic speaker with a known transfer function generates a dynamic pressure in the air using a sinus excitation of the speaker\u2019s membrane. The sensor measures the resulting variation of pressure in front of the speaker. The desired pressure range can be analysed step by step and consequently the dynamic behaviour of the sensor can be determined. Fig. 8 Simulation of three kinds of oscillators. at UNIVERSITE LAVAL on July 2, 2015ije.sagepub.comDownloaded from International Journal of Electrical Engineering Education 45/2 Figure 9 shows the successful result of the development. The fi rst execution of the experimental environment in the subsequent summer course showed that the participating students achieved their didactic aim. After participating in this laboratory the students have a comprehensive overview of static and dynamic pressure measurement. Behind the scenes: general conditions To fulfi l the didactic demands on the PEM some aspects of the generation and execution of a PEM-seminar have to be taken into account. Thus, the themes should have some criteria to show: 1 Intensive coaching and monitoring of the scientifi c assistant; 2 Preferably a wide solution space; and 3 Preferably an interdisciplinary/electromechanical topic. The main focus of the PEM-seminar lies in the conception phase. Thus, the task should also allow treatment that is of a didactically appropriate breadth. Our experience of the past few years has shown that a topic can not be too easy from the didactic point of view. In particular, a manageable theme will encompass both creativity and refl ect the chosen way to reach the solution. This enlarges the solution Fig. 9 Photo of the developed experimental environment. at UNIVERSITE LAVAL on July 2, 2015ije.sagepub.comDownloaded from International Journal of Electrical Engineering Education 45/2 space, improving the grounding of the decisions made by the students. Simultaneously the learning effect increases and so does the value of the experience for the students. In summary it may be said that themes are applicable for a PEM-seminar if a wide solution space and consequently a range of innovative results are desired. The strict executing of defi ned work packages would miss the didactic aim of independent organisation during goal-oriented working. The scientifi c assistant plays the role of a heedful looker and a technical consultant. He scrutinises the methodical proceeding, reminds about forgotten partial problems and demands reworking if necessary. This procedure has the aim to inspire an independent refl ecting decisionmaking process. This mentoring results in high amount of work for the scientifi c assistants. But regarding the student\u2019s advancement, it\u2019s worth it. Another aspect should not be underestimated: students, especially during the fi rst PEM seminars, need assistance in methodical working. Understanding a theoretical content does not mean that you can apply it to your special topic. Defi ning, varying and further detailing the function structure brings about most problems. This required non-intuitive working can only be acquired in its whole achievement with much practice." + ] + }, + { + "image_filename": "designv11_83_0001001_aim.2008.4601625-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001001_aim.2008.4601625-Figure4-1.png", + "caption": "Fig. 4 First Type X-crawl walking.", + "texts": [ + " The basic principle of quadruped walking is that, the ground touching tip points should form a triangular support area which embraces the CoG so the stability of the body is already maintained and then joint actuators make the body shift. Therefore, walking can be considered as an inverse kinematics problem. The swinging leg has no contribution to the progress of the body. Legs swing in an appropriate sequence for the locomotion of the body and forming the new support plane in the next step. X-crawl in [2] is briefly explained in this section. As can be noted from Fig. 4, suppose the legs\u2019 reachable areas are indicated as rectangles and legs will move over the midline of the rectangles. In every step the body progresses 1/4th of the length of the midline and the swing leg takes a step equal to 3/4 of the length of the midline. As shown in Fig. 4(a) the support triangle is formed by legs 1, 2, 4. The third leg starts to swing. After the third leg swings (takes a step), the body and the legs takes the position shown in Fig. 4(b). In Fig. 4(b), the body is shifted 1/4th of the length of the midline from the position in Fig. 4(a). It seems that the tip points of supporting legs (black dots) moved backwards, but actually they did not moved, the actuators of the supporting legs make the body shift and new reachable areas for the new body position are formed. In other words, supporting tip points are fixed, but the actuators of the supporting legs are energized to shift the body in +x direction. Only the third leg\u2019s tip point progresses in the air to take a step. In Fig. 4(b) the support triangle is formed between the legs 2, 3, 4 and the next leg to swing will be the leg1. In Fig. 4(c) and Fig. 4(d), the same process is repeated. In Fig. 4(c), the support triangle is formed between the legs 1, 2, 3 and the next leg to swing will be the leg4. As in Fig. 4(d), the support triangle is formed between the legs 1, 3, 4 and the next leg to swing will be the leg 2. If one more step is wanted to be taken from the position in Fig. 4(d), it is obviously seen that the next position will be the position same as Fig. 4(a). As can be seen in Fig. 4 (a) and (b), the CoG is in the border of support triangle which means no static margin for stability. Remember that the reachable rectangles in Fig. 4(a, b, c, d) are not same, after every step these reachable areas formed again due to the new position of the body. In Fig 4, the swinging sequence is 3 - 1 - 4 - 2. It can also be 4 - 2 - 3 - 1. X-crawl in [1] is briefly explained in this section. As mentioned in introduction, this technique has an important advantage that after every gait cycle the robot takes back its initial position. In this manner a gait cycle is divided into four sections. After every section the progress of the body is S/4, where S denotes the step size. All of the legs take the same step size. In Fig 5(b), the third leg swung, remaining legs supported the body and made the body progress S/4, the support triangle is formed between the legs 1, 2, 4", + " For the first and second type of X-crawl, swing time for one leg is taken as 1.5s. For the third type of Xcrawl (Trotting) swing time is chosen as 0.7s. For a leg to complete its swing motion, first of all basalar joint of the corresponding leg lifts the leg up, then its thoracic joint makes the leg to swing and finally again basalar joint forces the leg to land. Meanwhile the supporting legs shift the body in the +x direction. The thoracic joint positions of leg1 and leg3 are shown in Figure 7. The interval between 0-1.5s is for taking the initial position in Fig. 4(a). Leg3 starts to swing at t=1.5s. The constant positions of the actuators mean the legs are either lifting or landing. As can be seen there is a phase difference between leg1 and leg3 which can also be seen from the thoracic joint positions from Fig4. Leg2 and leg4 have also a similar phase difference which is not shown in this figure. In Figure 8 the thoracic actuator positions in two gait cycles of the second type of X-crawl are shown. In this gait type the simulation is done with maximum step size calculated from the equation (4). The first swing leg is leg 3. Thoracic angle of leg3 is zero for 0.3s which means the leg is lifting. When t=1.3s the thoracic actuator of the third takes its position and then the actuator is constant for landing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002732_j.engfailanal.2021.105451-Figure17-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002732_j.engfailanal.2021.105451-Figure17-1.png", + "caption": "Fig. 17. Positions of strain gauges T1-T3, radial plates and cylinder of undercarriage.", + "texts": [ + " This was expected based on the results of the calculation, because due to the limitations of the experiment, a reduced vertical load was adopted, which is not of sufficient intensity to revive all the gauges. Strain gauge T5 is positioned on upper horizontal plate of undercarriage, where vertical plate connects the horizontal plate. Strain gauge T8 is positioned on the upper horizontal plate of slewing platform where vertical plate connects the horizontal plate. Rosette (T1-T3) is positioned on the vertical plate of undercarriage, near the outer cylinder, in the zone of stress concentration. Fig. 17 shows the position of the rosette. The results obtained using strain gauges T1-T3 are compared with results obtained using system for optical non-contact measurement, and this group of the results is named MM 1\u20133. The position of the recording equipment during measurement conducted in spot MM 1\u20133 is shown in Fig. 18. A line in stress field obtained using DIC that imitates a strain gauge is shown in Fig. 19. Namely, the software provides the ability to determine the distance between any two points at any time" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002004_50015-9-Figure12.24-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002004_50015-9-Figure12.24-1.png", + "caption": "Figure 12.24 (a) Schematic illustrating a defected photonic crystal being used as a waveguide bend. (b) A proposed 3-D photonic crystal (Povinelli and co-workers, 2001, by permission of American Physical Society).", + "texts": [ + " Just as electrons are scattered by crystals and exhibit phenomena such as Bragg reflection governed by dispersion surfaces, similar properties can be exploited for monochromatic light, but then the \u2018crystal\u2019concerned must have a periodicity comparable to the wavelength of light. A photonic crystal is therefore a tailor-made 2-D or 3-D periodic structure of dielectric material, with a lattice periodicity comparable to that of the optical wavelength. A simple application is a waveguide bend, shown in Figure 12.24a, which is a 2-D lattice of particles of a dielectric material (silica, alumina, etc.), but with missing particles along a bent path. The spacing of the particles is such that strong Bragg reflection is experienced by the incident light, so that the latter is prevented from entering the lattice. The light is thus forced to travel along the curved path of missing particles. More complicated 3-D structures have also been proposed (Figure 12.24b) which exhibit interesting and useful dispersion properties. 12.4.6 Mechanical properties of small material volumes The strength of submicron-sized metals and ceramics is considerably higher than their bulk forms. In Chapter 10, we have seen that the fracture strength of brittle glasses and ceramics is strongly size dependent, due to the decreasing chance for the material to contain large flaws as its size reduces. Thus, glass and carbon fibers are orders of magnitude stronger than their bulk counterparts, and they are highly flexible so they are effective strengtheners and stiffeners in plastic-based composites (Section 11" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000650_1.2747505-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000650_1.2747505-Figure1-1.png", + "caption": "FIGURE 1. Small-diameter and high-pressure ECT sensor with 8 electrodes", + "texts": [], + "surrounding_texts": [ + "To a certain sensor pipeline, the capacitance values between electrodes are mainly determined by electrode area. For a given electrode number and angle, the capacitance between electrodes is proportional to the axial length of electrode [3], so to make capacitance measurement more easily, long electrode is preferable. However, too long electrode can result in spatial filter effect, and make the sensor loss the response ability to high frequency signal of spatial structure. Therefore, it is necessary to calculate the sensor capacitances so that both of the requirements are satisfied. The result of ECT sensor capacitances simulated by ANSYS 9.0 is shown in Fig.2. For the 8-electrode system, sensor electrode 1 is selected as excitation electrode, and the others as detect electrodes, so 7 electrode pair capacitances can be calculated. The result includes the empty sensor capacitances when the sensor is filled with air and the full sensor capacitances when it's filled with coal powder. The relative dielectric constant of coal powder, air, pipeline and filling layer is 3.5, 1, 8 and 2.5 respectively. From the simulation result, the standing capacitance of adjacent electrodes is about 60 pF per meter, and they are 10 to 100 times larger than the other electrodes. However, the capacitance change of adjacent electrodes is about 0.53 pF/m, smaller than the other ones (1.3 to 1.75pF/m). Further more, their changes is negative because the pipeline wall is too thick compared with the sensor inner diameter so that the high sensitivity area is occupied by sensor pipeline wall [4]. The above capacitance simulation gives a powerful reference of ascertaining the axial length of electrode. For the purpose of enabling the measurement circuit to measure the capacitance changes and keeping adequate frequency response ability, the adopted electrode length is 10 cm [5]. Thus, the capacitance changes between empty and full sensor range from 0.05 pF to 0.18 pF." + ] + }, + { + "image_filename": "designv11_83_0000748_s0890060408000036-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000748_s0890060408000036-Figure1-1.png", + "caption": "Fig. 1. The arc placed tangential to the involute.", + "texts": [ + " A multilayer feedforward NN based on a BP algorithm is used to train the network, and the predicted values of the fillet stress are verified using the results obtained with FEM. In the first step, an involute profile is obtained from the first principle as the locus of a point on a tangent line rolling over the base circle without slipping, and the spur gear tooth flank is obtained by rotation of this profile through a half-tooth angle. With the gear center as the origin and the centerline of the tooth as a reference (Fig. 1), polar angle ui of a point on the flank at radius Ri is given by ui \u00bc u0 \u00fe inva0 invai, (1) where u0 \u00bc p/2z and ai \u00bc cos21(Rb/Ri), z is the number of teeth on the gear, a0 is the design pressure angle, and Rb represents the base circle radius. The radius Ri is varied from the tip circle to the base circle in an incremental manner (Dr) to construct the complete involute profile. For cases where the root circle is inside the base circle, the profile is further extended by a radial line up to the root circle. In the next step, a circular arc is placed such that it is tangential to the involute curve as well as the root circle (Subba Rao et al., 1993). This is done using a computer-aided approach as described below. The arc with a radius (rf ) is divided using the same incremental radius (Dr) as shown in Figure 1, and angles subtended fi are calculated using cosfi \u00bc (Rr \u00fe rf )2 \u00fe (Rr \u00fe rf si)2 r2 f 2(Rr \u00fe rf )(Rr \u00fe rf si) : (2) In the next step, the angle ui of the involute profile and corresponding fi of arc are added. The maximum angle umax is obtained with respect to the tooth centerline as umax \u00bc max (ui \u00fe fi): (3) The radius Rmax corresponding to umax yields a point at which the arc is tangential to the involute. To establish the circular arc below the tangential point corresponding to Ri, the angle ci is calculated as follows ci \u00bc umax fi: (4) To compare the proposed fillet with the trochoidal fillet, the tooth flank along with the trochoidal fillet is generated using a basic rack with a tip radius Rc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000032_iecon.2005.1569121-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000032_iecon.2005.1569121-Figure2-1.png", + "caption": "Fig. 2: Schematic view of two in-wheel motors equipped with the fault-tolerant inverter configuration considered. Dashed lines indicate additional cabling to achieve fault tolerance and circles indicate additional connection and isolation devices.", + "texts": [ + " However, if adopted in a multidrive system, using a common dc-link and equipped with only one additional inverter leg to handle faults in several PMSMs, the connection of the additional inverter leg can be made simpler (less expensive), since it is connected to the neutral point of the faulted PMSM rather than to the specific faulted phase. An example of a multi-drive system that can benefit from the configuration considered is an electric vehicle equipped with two (or more) in-wheel motors, as exemplified in Fig. 2. The well known three-phase model of a salient PMSM (see, e.g., [10]) is applicable with the additional constraint that the current in the isolated phase is identically zero in fault mode. Assuming that the faulted inverter leg (corresponding to phase a) has been disconnected, the phase-toneutral voltages, vb,n and vc,n (see Fig. 1), are governed by vbc,n = Rsibc + \u03c9r \u2202Lbc,n \u2202\u03b8 ibc + Lbc,n dibc dt + \u03c9r\u03c8bc (1) where Rs is the stator resistance, vbc,n =[vb,n vc,n]T , ibc = [ib ic]T and Lbc,n = 1 3 [ L11 L12 L21 L22 ] (2) \u03c8bc = \u2212\u03c8m [ sin(\u03b8 \u2212 2\u03c0/3) sin(\u03b8 + 2\u03c0/3) ] " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002498_iwed52055.2021.9376353-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002498_iwed52055.2021.9376353-Figure5-1.png", + "caption": "Fig. 5. Rotor and stator configuration of the interior permanent magnet wheelhub motor.", + "texts": [ + " The d- and qaxes voltage references are converted into the \u03b1\u03b2 stationary reference frames and then into three-phase voltage references. Each of these references is applied to its openend winding. The current feedback is read from all three phase current sensors due to the fact that their sum is not equal to zero having additional zero sequence component. These three currents are transformed into the \u03b1\u03b2 reference frames and then to the dq reference frames in order to provide the feedback to the current controllers. III. SIMULATION RESULTS The sketch of the traction motor is depicted in Fig. 5. Parameters of the machine are listed in Table I. It is interior permanent magnet synchronous machine with the continuous power of 100 kW and the maximum peak power of 200 kW. The continuous torque density of the motor taking into account only active materials is 13.4 Nm/kg, which is much bigger than for the competitors [15] for continuous operating mode. Such a high torque density was granted by specific parameters of the machine, where the ohmic losses are predominant at the low speeds and the iron losses are predominant at the high speeds" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003120_j.promfg.2021.06.034-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003120_j.promfg.2021.06.034-Figure5-1.png", + "caption": "Figure 5: A) Next closest pathing where islands are greedily chosen. B) Global minimum where the true shortest path is calculated", + "texts": [ + " Both general investigations were carried out on a Dell workstation with an Intel Xeon W-2133 CPU processor, 32GB of RAM, and an Nvidia QUADRO GPU with 5GB of VRAM. A practical example of a V-shaped object was shown in Fig. 2. Naturally, numerous variables contribute to overall build time including physical properties of the system such as velocities and accelerations, interruption by table motion, ratio of travel to build motions, as well as the geometry itself including sparsity and pathing density inside the islands. This geometry was meant to provide a basic example of a typical large-scale object. In Fig. 5A and 5B, two different travel patterns are shown. The first is the commonly used next closest island algorithm while the second is the global minimum calculation. The distance traveled by the machine was excluded for both cases as the object was in the same location in the build volume when it was sliced. For the first travel algorithm, the total travel distance was found to be 89.1 inches while the second travel algorithm was found to be 80.8 inches. This corresponds to a travel distance that is approximately 9% shorter" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001607_cca.2008.4629655-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001607_cca.2008.4629655-Figure3-1.png", + "caption": "Fig. 3. Description of formation", + "texts": [ + " Measurement of the Formation Deviation Whenever mobile systems on a coordinated lawnmower have the task to measure data that shall be merged together afterwards, it is important to keep the desired formation. The quality of the formation needs to be calculated in every time step and added over the mission time to judge the employed procedure. Let FD be the general value of the Formation Deviation, fd(k) the value of the kth of l steps with a step size of T. Then FD is defined as ( ) 1 l k TFD fd k l = = \u2211 . (4) Fig. 2. Basic definitions for track following For the definition of the Formation Deviation term in a single time step, we use a formation description like depicted in Fig. 3. The orientation of axes follows the NED coordinate system [7]. The adjustment of the vehicles is described by the angle \u03b1m for vehicle m of p. This is the angle between the vectors from vehicle m to m-1 and to m+1 respectively. It is postulated that the vehicle numeration is done that way that there are no intersections between these lines to the proximate neighbors (formation pattern as npolygon). Each n-polygon is described by n-1 angles, so one of the vehicles need not to be checked. The link between every vehicle and its follower (m,m+1) is described by the distance dm,m+1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002255_tmag.2021.3057124-Figure13-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002255_tmag.2021.3057124-Figure13-1.png", + "caption": "Fig. 13. Eddy current density distributions in rotor structures after sudden short circuit. (a) After one cycle. (b) After five cycles. (c) After ten cycles. (d) After 20 cycles.", + "texts": [ + " From these results, both dc and ac curves obtained by the proposed method well agree with the measured curves, while those obtained by the conventional 2-D analysis are not in good agreement. Especially, the initial transient profile in the 2-D analysis differs from the measured one. This would be because the effects from conductive structures in the end region are not considered. In contrast, by using the proposed method, such end-region effects can effectively be considered. The eddy current density distributions after the sudden short circuit are shown in Fig. 13, which makes the electromagnetic distributions in the turbine generator clear. For example, we can see that eddy current concentrates on the cross-slot and the wedges in the end region. Based on these results, the designers can discuss magnetically weak points of the generators in case of transient conditions such as power system faults. Using the proposed method, it takes about from 1.5 to 2 months when the number of cores in sharedmemory-type parallel computation is 4. According to the result shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003132_978-981-16-0119-4_20-Figure71.2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003132_978-981-16-0119-4_20-Figure71.2-1.png", + "caption": "Fig. 71.2 Motorcycle steering assembly [7]", + "texts": [ + "4) where Vp is volume of the part, Vb is the volume of bounding box, Ap is the surface area of the part, As is the surface area of the sphere which is considered to be a less complex shape, and Nh is the number of holes. These geometry-related parameters are measured from the mass properties of the 3D CAD model. If the MCF value is more than 44 falls under the high complexity part category, these designs are more suitable for AM according to [6]. out. Section 71.3.1 discusses an existing motorcycle steering assembly, which is already discussed in [9], and Sect. 17.3.2 discusses a throttle pedal assembly discussed in [8]. Initially, there were seven components in the design, and Fig. 71.2 shows the initial design and components in the design. The total number of parts in the assembly is reduced to four by using the complex product network measure. The network from the relationship matrix is drawn as shown in Fig. 71.3a, and the centrality score is calculated for each node. Based on the centrality score, the candidate for part consolidation is identified and part count reduced to four from seven. The product network after the consolidation is shown in Fig. 71.3b, and the centrality score of each part is graphically represented in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000368_detc2004-57064-Figure12-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000368_detc2004-57064-Figure12-1.png", + "caption": "Figure 12. Parallel Manipulator Gough-Stewart.", + "texts": [ + "org/about-asme/terms-of-use D dimensional algebra of spatial translations Am f a 3 j 1 Am f j t 3 Thus, this parallel manipulator satisfies the assumtions of proposition 2, and the mobility of the parallel manipulator is given by equation (8), F r \u2211 i 1 fi k \u2211 j 1 dim Am f j dim Am f a 12 3 \u2211 j 1 4 3 3 Furthermore, for the passive degrees of freedom, see equation (30), Fp r \u2211 i 1 fi k \u2211 j 1 dim Am f j 12 3 \u2211 j 1 4 0 5.7 Gough-Stewart Parallel Manipulator. Consider a Gough-Stewart parallel manipulator, shown in Figure 12. The manipulator is formed by six serial connecting chains. Each serial connecting chain is formed by a pair of spherical pairs located at the ends of the serial connecting chains and a prismatic pair. The kinematic pairs of each serial connecting chain generate the Lie algebra, e 3 ; i.e. Vm f j Am f j e 3 j 1 2 6 It should be noted that the Lie algebra has dimension six and 7 \u2211 i 1 f ji 7 dim Vm f j dim Am f j 6 j 1 2 3 4 5 6 Furthermore Am f a 6 j 1 Am f j e 3 Thus, the parallel manipulator is a trivial one, this class of platforms was analyzed in Proposition 3 in Section 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003433_978-3-030-40667-7_6-Figure2.1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003433_978-3-030-40667-7_6-Figure2.1-1.png", + "caption": "Fig. 2.1 Illustration of the hexapod with the reference frame of the Base (x y, z) (B-frame), that is fixed, the Platform reference frame (x\u2032, y\u2032, z\u2032) (T-frame), mobile relatively to the base. The frame rotated compared to the Base frame, the M-frame (x\u2033, y\u2033, z\u2033), will be used in an example at the end of the chapter", + "texts": [ + "1007/978-3-030-40667-7_2 Mathematics of\u00a0the\u00a0Hexapod Giancarlo\u00a0Ruocco and\u00a0Alfonso\u00a0Alessandro\u00a0Tanga Nomenclature B-frame orthogonal reference system in the base bi the vector defining the coordinates of the lower anchor point of the leg ci number of constraints that a joint imposes D interval of positional variables DoF degrees of freedom fi degrees of freedom of the joint i number of the leg (1, \u2026, 6) j number of joints in a system li the vector representing the ith leg mb the translation vector in M-Bar M mobility M-BAR the segment that connects the two platforms\u2019 center (O and O\u2032 of the M and M\u2032 frame) M\u2032-frame orthogonal reference system in the platform M-frame orthogonal reference system in the base n number of connected rigid bodies in a system qi the vector that connects the origin on the base to the end of the leg Pi end of the leg Pi the vector that represents the position of the end of the leg on the platform in the T-frame PRB rotational matrix T-frame orthogonal reference system in the Platform The hexapod can be described as two non-regular hexagons (Fig.\u00a0 2.1), whose vertices are connected by six segments. The geometry of the hexapod can be schematized as a structure composed of three different parts. First, one fixed plate, referred to as the Base. Second, a moving plate, named the Platform, which can reach a great variety of positions and orientations, depending on the specific location of the vertices. Last, the six segments of variable length, denominated the Legs of the hexapod. From this point forward, we will consider the center of the Base as the origin of the reference framework, the B-frame with orthogonal axes x, y, and z", + " Obviously in the case of the hexapod this is not real, because of physics limitations, like the length of the legs that connect the two plates. Another limitation is that the two plates cannot be superimposed one on the other. The legs of the hexapod can be mathematically thought as vectors. The vectors are mathematical object that describe a quantity that can be well characterized by an intensity, a direction, and a verse. In this case, the intensity is the length the of the leg, the direction is given by the three rotation angles, and the verse is always positive, because it is always in the z\u00a0 >\u00a0 0 half-space. In Fig.\u00a02.1, the vector that describes the i-th leg that connects the origin and the point Pi is qi. To obtain this vector, the position of the platform in the B-frame must be known. This information is inside the translation vector M-Bar (mb) that gives the positional linear displacement of the origin of the platform frame with respect to the Base reference framework. Another information needed is the position of the end of the leg on the platform in the T-frame, and this is the vector (pi). It must be multiplied by the rotational matrix PRB (described in the next paragraph) to obtain its coordinates in the B-frame", + " Since the existence of 40 configurations of the general Stewart-Gough platform had been first demonstrated numerically by Raghavan [1], many researchers have applied elimination method to find all the solutions of the problem. An example of a possible numerical solution through a minimization will be shown. In this example [2], we define a set of six variables and define their correlation with the length of the six legs (l1, \u2026, l6). The set of new variables [1] (\u03b11, \u03b12, \u03b13, \u03b14, \u03b15, mb) is used as positional variables to express the position and orientation of the mobile platform relative to the base platform. mb is the already defined length of the M-bar (shown in Fig.\u00a0 2.1); (\u03b11, \u03b12) denote the transformation angles (\u03b8, \u03c8) from B-frame (x; y; z) to M-frame (x\u2033; y\u2033; z\u2033); (\u03b13, \u03b14, \u03b15) denote the transformation angles (\u03b8, \u03c8, \u03c6) from M\u2032-frame to T-frame, where M\u2032-frame is translated from M-frame by moving the origin of M-frame to the origin of T-frame. The positional variables of the mobile platform may be calculated by solving the following equation: R w R t b l iT i i ibm mt , ,+( ) - = = \u00bc( ) 2 2 1 6 where w\u00a0=\u00a0(0, 0, mb), Rbm\u00a0=\u00a0R(\u03b11, \u03b12, 0), Rmt\u00a0=\u00a0R(\u03b13, \u03b14, \u03b15), bi\u00a0 =\u00a0 (bi1, bi2, bi3)T, and ti\u00a0 =\u00a0 (ti1, ti2, ti3)T denote coordinates of the joints of the ith leg on the base and on the mobile platform, respectively, (i\u00a0 =\u00a0 1, 2, \u2026, 6)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002732_j.engfailanal.2021.105451-Figure20-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002732_j.engfailanal.2021.105451-Figure20-1.png", + "caption": "Fig. 20. Position of strain gauges T6, T7 and T9, radial plate and cylinder of slewing platform.", + "texts": [ + " Thus, line is drawn along in the vertical direction in the rosette zone, which practically imitates the T2 strain gauge. According to the table, a high matching between experimental results and numerical calculation can be noticed. It can also be noticed that everything that can be measured with strain gauges, can also be measured with a system for non-contact measurement of A. Petrovic\u0301 et al. Engineering Failure Analysis 126 (2021) 105451 displacement and deformation, faster and more efficiently. Strain gauges T6, T7 and T9 are positioned as shown in Fig. 20. The measurement results at MM 6 are given below. In Fig. 21, a field of vertical displacements that is fully aligned with the field obtained by the numerical calculation can be seen. This is the point around which the pylons bend slewing platform. The measurement results at MM 9 are given below. At this point, a considerable difference between the numerical calculation and the experimental results was observed. The reason for this are the massive welded joints near this place, because the cylinder, the upper plate of slewing platform and the knot plate of sprit join there" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000886_pime_conf_1965_180_339_02-Figure10.14-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000886_pime_conf_1965_180_339_02-Figure10.14-1.png", + "caption": "Fig. 10.14. Pieces of bearings which have sufleered damage by test in liquid sodium-", + "texts": [], + "surrounding_texts": [ + "Another shortcoming of Fig. 10.3 is that it makes no reference to the material of the mating surfaces. T o complete the picture it would be necessary to include such features as hardness, softening temperature, and melting temperature as well as chemical characteristics such as proneness to oxidation. In the ideal case of hydrodynamically lubricated plain bearings the bearing material should have no effect on performance being separated by a complete fluid film. However, in practice, the performance may be markedly affected by material properties. The existence potassium eutectic lubricant Proc Instti Mech Engrs 1965-66 VO! I W Pt 3K at UNIV CALIFORNIA SANTA BARBARA on July 22, 2015pcp.sagepub.comDownloaded from Proc Instn Mech Bngrs 1965-66 THE INVES'I'IGAI'ION OF UNUSUAL BEARING FAILURES at UNIV CALIFORNIA SANTA BARBARA on July 22, 2015pcp.sagepub.comDownloaded from 290 Proc Instn Mech Engrs 1965-66 F. 1\u2019. RARWELL AND D. SCOTT Vo1180 Pt 3K at UNIV CALIFORNIA SANTA BARBARA on July 22, 2015pcp.sagepub.comDownloaded from THE INVESTIGATION OF UNUSUAL BEARING FAILURES 29 1 of hydrodynamic conditions under practical degrees of misalignment and departures from true cylindrical form is often only rendered possible by the conformability of the lining material and the capacity to absorb particles of foreign material may be of critical importance. An unusual type of cement-kiln bearing failure was described as \u2018wire-drawing\u2019 (33), because the first manifestation was the emergence of wire-like debris from the bearing. The gun-metal bearings and steel shafts were both heavily scored, Fig. 10.19. Bearings in progressive stages of damage were examined and a common feature was incipient scoring, some scratches terminating in an embedded ferrous particle which had a striated structure, Fig. 10.20. Micro-hardness tests revealed that the particles were extremely hard and that the surface and deformed material in the vicinity of the scores in both bearing and shaft had been work-hardened. Failure appeared to have initiated from a particle deformed and hardened by working between the bearing and the shaft. The particles appeared to score the surface of the bearing until the deformed bearing material ahead of the particle was sufficiently work-hardened to resist further deformation. The hard, partially embedded particle could then machine the shaft providing debris to initiate cumulative action and thus produce a series of ridges and grooves which in turn deformed and scored the bearing. It can be envisaged that the lateral movement allowed in such a large bearing adequately power driven could produce, by shearing of the ridges, rings of metal which would break up and work out of the bearing as debris in the form of wire. As the problem in this specific installation was solved by the substitution of oil-lubricated white metal bearings in place of the grease-lubricated gun-metal bearings, thus enabling stray particles to be deeply embedded in the white metal to prevent shaft wear, no Proc Insrn Mech Engrr 196546 further research effort was required. However, a similar problem termed \u2018wire-wool\u2019 type failures by Dawson (34) (35) has arisen in the white metal bearings in large steam turbines and generator rotors with the introduction of chromium-molybdenum steels as improved material for the journals. The phenomenon has also been referred to as \u2018black mark\u2019 bearing failure as the initial stage appears to be the impregnation of the white metal with debris to give a black surface appearance. The problem is being investigated, contributory causes and remedial measures suggested, and palliatives assessed by laboratory service simulation tests (35). The phenomenon does not appear to be connected with the tin oxide corrosion problem of power plant turbines and some marine bearings (36) (37), the mechanism of which is as yet not fully understood. An unusual type of bearing failure was brought to the attention of the National Engineering Laboratory when an investigation was requested to determine if a failed shaft had at any time been built up by welding or electrical deposition. No apparent trouble had been experienced with the large lift shaft during almost continuous service for a considerable period. An observant operator, however, noticed slivers of steel when changing the oil and initiated an inspection. The shaft exhibited a curious type of wear extending over about one-half of the length of one of the two main cast-iron bearings, Fig. 10.20. The remainder of the shaft bearing surface was slightly pitted and scored, Fig. 10.21, but relatively unaffected compared with the damaged area. The shaft was sectioned transversely through the damaged area. The damaged surface was patterned in a regular series of cracks and the material between the cracks had suffered severe plastic deformation, Fig. 10.23. There was no evidence of welded, electro-deposited, or case-hardened material. Yo1 180 Pr .1K at UNIV CALIFORNIA SANTA BARBARA on July 22, 2015pcp.sagepub.comDownloaded from 292 a Shaft. b Bearing. F@. 10.19. ' Wire-drazuing' type bearing ,failiire Proc Insrn Mrch Engrs 1965-66 F. T. BARWELL AND D. SCOTT at UNIV CALIFORNIA SANTA BARBARA on July 22, 2015pcp.sagepub.comDownloaded from THE INVESTIGATION OF UNUSUAL BEARING FAILURES Proc Instn Mech Engrs 1965-66 293 Vol180 Pr 3K at UNIV CALIFORNIA SANTA BARBARA on July 22, 2015pcp.sagepub.comDownloaded from 294 Proc Inrrn Mech Erigrs 1965-66 F. T. RARWELL AND D. SCOTT V01 180 Pr 3K at UNIV CALIFORNIA SANTA BARBARA on July 22, 2015pcp.sagepub.comDownloaded from Proc in st^ M ech Etzgrs 1965-66 20 T H E IN V E ST IG A T IO N O F U N U SU A L B E A R IN G FA IL U R E S x X B d d .- 9 295 i\u2018uJ 180 P t 3X at U N IV C A LIF O R N IA S A N T A B A R B A R A on July 22, 2015 pcp.sagepub.com D ow nloaded from 296 F. T. BARWELL AND D. SCOTT Micro-hardness exploration showed that the hardness of the deformed material varied from 189 to 287 HV, whilst the hardness of the undeformed surface shaft material was 110 HV 30. Interrogation of lift operators revealed that throughout the service life, periods of judder and vibration had been experienced and it would appear that incipient failure had occurred at an early stage, but that the powerful driving motors had forced shaft movement deforming the broken surface to conform with the reduced bearing clearance, debris being rolled along between the surfaces. Cumulative action had led to the curious surface pattern and to the production of metal slivers in the lubricant which was castor oil. This action destroyed evidence of the initial failure but it is likely that this may have been due to surface fatigue brought on by corrosion." + ] + }, + { + "image_filename": "designv11_83_0000474_bfb0031450-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000474_bfb0031450-Figure1-1.png", + "caption": "Fig. 1 : Frames and transformations", + "texts": [ + " The method is assuming that tactile sensor arrays provide the contact point location relative to the local phalanx frame together with the local curvature of the object. Some experiments [6] with tactile sensors showed already that it is possible to use tactile sensors for this purpose. 3. MATHEMATICAL MODEL OF OBJECT AND MULTI-FINGERED HAND 3.1 General equations. For describing the object-finger interaction, the formalism as proposed by J.C. Trinkle [9] and J.R. Kerr [8] is used. The multi-fingered hand, considered here, is consisting of a 3 d.o.f. index and a 2 d.o.f, thumb. Figure 1 shows an object with frame B0 fixed to it. The origin of B0 coincides with the centre of gravity Z of the object to be grasped. Similarly, the i,f h link of the hand has the frame Bij fixed to its centre of gravity. The subscript i and j denote the number of the finger and the number of the phalanx. The palm is considered to be the base link of all fingers. The frame Bo,o is fixed to the centre of gravity Z of the palm. The world frameX is any conveniently chosen inertial frame. The frame Ti,o is fixed to the palm and defines the first joint of the i th fmger" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002445_tim.2021.3062163-Figure9-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002445_tim.2021.3062163-Figure9-1.png", + "caption": "Fig. 9. (a) Test rig of the rotor system [54] and (b) diagrammatic sketch, and (c) enlarged view of rubbing structure.", + "texts": [ + " For these faulty characteristics, it is difficult to separate their weak harmonic components using traditional methods such as EMD, LMD, and VMD, and mode aliasing is easy to happen. As a result, the frequency feature cannot be extracted well by these methods. In this part, NCCD is applied to a rotor rubbing experiment and compared with other algorithms to verify its effectiveness in processing fast-time-varying signals. Yu et al. [49] from the School of Mechanical Engineering of Northeastern University collected the experimental data. Fig. 9 displays the layout of the rotor test bench. The motor drives the entire rotor system to rotate at a constant speed of 5100 rpm through the elastic coupling. Two bearings on both sides support the rotor, and there is a disk in the center of the rotor system. On the bracket on one side of the disk is a friction rod with an adjustable position. We can simulate the rotor system\u2019s rubbing phenomenon by changing the distance between the friction rod and the disk. Besides, we use the electric eddy current sensor with a sampling rate of 2000 Hz to collect the radial displacement information of the shaft" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000807_07ias.2007.206-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000807_07ias.2007.206-Figure5-1.png", + "caption": "Figure 5. Cross section of the five-phase outer rotor PM machine", + "texts": [ + " )()()()()(),,( 221121 igipigipihppif ++= iFipQiKTipii TTT .)())(( 2 ' 12 1 +\u2212+= (12) Where )(1 ip and )(2 ip are Lagrangian multipliers. The currents in the healthy phases for fault-tolerant operation can be obtained as below: QFQKKFQKKn QKFFQKnTi TTT H T H +\u2212 +\u2212= '* (13) The output torque should be adjusted accordingly to keep the stator phase currents under the maximum current limit. IV. SIMULATION RESULTS A five-phase, eight-pole outer rotor PM machine is considered for the verification of the proposed method. The machine cross section is shown in Fig 5. The permanent magnets are NdFeB, with TBr 2.1= . The stator has concentrated windings. Each stator winding is split into two equal halves. Flux linkage in the stator phase windings obtained from Finite Element Analysis (FEA) of the machine is presented in Fig 6(a). The static output torque of the machine when supplied with optimal current waveform under normal operation is shown in Fig 6(b). The average torque is 14.6 N.m and the torque pulsation is about 6%. The rms value of the stator phase currents is 10A" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003407_j.optlastec.2021.107425-Figure11-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003407_j.optlastec.2021.107425-Figure11-1.png", + "caption": "Fig. 11. Change curves of surface sintering pool structure and central temperature of powder bed under different scanning speeds (a) V = 1.6 m/s (b) V = 1.8 m/s (c) V = 2.0 m/s (d) V = 2.2 m/s (e) Change curves of temperature (f) Change curves of temperature (enlarged drawing).", + "texts": [ + " The main reason is when the laser power is small, less calories absorbed by powder bed, powder bed surface temperature is low, can not melt all the powder in laser radiation area. With the laser power increasing, more calories absorbed by powder bed, powder bed surface temperature is high, melting amount of the powder in laser radiation area increase. When all the powder melt, continue to increase the laser power, causing small melting amount of powder, mainly due to its small thermal conductivity coefficient. Fig. 11 shows the change curve of sintering pool structure and central temperature on the powder bed surface at different scanning speeds (1.6 m/s, 1.8 m/s, 2.0 m/s, 2.2 m/s) when the laser power is 10 W, the layer thickness is 0.1 mm, the preheating temperature is 80\u2103 and the scan spacing is 0.2 mm. From Fig. 11 (a), (b), (c) and (d), it can be seen that as the scanning speed increase, structure size of sintering pool is decreasing, but the overall shape of sintering pool has no obvious change. With the scanning speed increasing from 1.6 m/s to 1.8 m/s, the sintering depth and width decreases significantly. With the scanning speed increasing from 1.8 m/s to 2.2 m/s, the sintering depth and width decreases, but the reduction is small. With scanning speed increasing, the center temperature of sintering pool declines, but temperature field distributions are basically the same, the isotherms closer to the edge of Y. Yu et al. Optics and Laser Technology 144 (2021) 107425 the sintering pool are denser, the sintering temperature outside pool edge presents the uneven gradient diffusion. Fig. 11(e) and (f) show that during the whole laser sintering process, the center temperature of the sintering pool increases rapidly from the preheating temperature and gradually cools down to room temperature of 20\u2103 after the laser sintering stops. When scanning speed increases from 1.6 m/s to 2.2 m/s, the center temperature of sintering pool changes significantly and gradually decreases, and the maximum value of the center temperature curve is 14.3 \u2103 and the minimum one is 9.5 \u2103. The main reason is when scanning speed is small, time of laser energy acting on the powder bed surface increases, causing heat absorption increases, powder bed surface temperature is high, prompting full melting of WSPC powder, melting amount increased, thus the structure size of sintering pool increases" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002382_j.medengphy.2021.03.002-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002382_j.medengphy.2021.03.002-Figure3-1.png", + "caption": "Fig. 3. Mechanical configuration of the instrumented ratcheting dial. A resistive rotary sensor is mounted within a custom housing that also holds the ratcheting element.", + "texts": [ + " Instrumented ratcheting dial The purpose of instrumenting the ratcheting dial is to track the ocket size adjustments made by users and to store that data with elease/relock data on the electronics board in the TARPIN housng. The combination of data from the ratcheting dial and the reease/relock system monitor user socket size adjustments, doffs, nd partial doffs over time. These data may be used for later analsis by patients and their prosthetists. To create the instrumented ratcheting dial, we modified a comercially available ratcheting dial from its original configuration by i T h t i a r t f t h w t f 7 t 1 b r 2 m r m a a m t n t i fi nstrumenting it with a sensor to track rotation during use ( Fig. 3 ). his modification requires a custom dial housing. The housing olds a single resistive rotary sensor. The ratcheting dial attaches o the rotary sensor using a custom shaft adaptor that is press fit nto the grooves on the bottom of the dial. The shaft end of the daptor is press fit into the rotating center of the sensor. As the atcheting dial is rotated by the user, the inside of the sensor roates relative to the dial housing and outputs a signal. Wires run rom the rotary sensor through tubing in the socket to the elecronics board within the TARPIN mechanism described above" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003436_j.apacoust.2021.108345-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003436_j.apacoust.2021.108345-Figure6-1.png", + "caption": "Fig. 6. Experimental set-up, front and top images.", + "texts": [ + " The sound radiation model of [26] considers the presence of two solids as equivalent cylinders, but therefore does not take into account the diffraction in geometric details such as gear teeth. Fig. 5. shows an example of an output of the acoustic model where the data of Fig. 3. is used as an input. For each of the impacts marked in Fig. 3, the sound emitted by the collision of a pair of 0 t tc \u00f010\u00de 1t\u00de \u00fe Ysin \u00f0xc l1\u00f0 \u00detc \u00fe l1t\u00de e l2 t tc\u00f0 \u00de \u00fe \u00bdEcos l1t\u00f0 \u00de \u00fe Fsin l1t\u00f0 \u00de e l2t\u00de For t > tc teeth is generated by the acoustic model, thus obtaining the acoustic pressure signal in function of time. The assembly used to validate the sound prediction is shown in Fig. 6. It consists of a gear pair driven by an electric engine with a transmission pulley of 2 to 1. The gear pair is mounted on shafts fastened on ball bearings, so the gears could roll free of load. The system is powered by a variable frequency drive. A microphone in line with the gear pair (h coordinate of the model equal to zero) through a DAQ system measures the sound signal. The set-up has the gear pair cantilevered to reduce sound reflection. In [26] the sound produced by the collision of two cylinders was validated, both in the aligned and vertical position. The aligned positioning of the microphone shown in Fig. 6. is determined for practical assembly reasons, additionally, the aligned location of the microphone allows to replicate the layout of the following approaching proposal of the MAVAS. Data collection and microphone calibration are performed using a LabVIEW script. The speed of the driving shaft is measured by a manual tachometer. Additional information about the gear pair and the measurement equipment can be consulted in Table 1. The sound produced by the pair of gears of Table 1 was measured in the experimental set-up when the driving wheel is rotating at 7" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001145_icbbe.2007.34-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001145_icbbe.2007.34-Figure3-1.png", + "caption": "Figure 3. Ichnography of screen-printed carbon paste electrode", + "texts": [ + " 2) Characterization Dynamic light scattering (DLS) is used to determine the average particle diameters and their size distributions of the grafted carbon black nano-particles. Fig. 2 shows the typical particle size distributions of CB-g-PSS. For CB-g-PSS, their average particle diameter was about 80nm because PSS was polyelectrolyte with negative charges supplying steric and static stabilization in water. Data and adscription (\u03b4, ppm): 7.93-6.20 (-C6H4), 2.25-0.96 (-CH2-CH (C6H4)). Get appropriate CB-g-PSS powder, made in 5mg/mL solution for using. Fig. 3 shows the structure of screen-printed carbon paste electrodes. Two carbon paste electrodes are work electrode and reference electrode respectively. Through wire, they can be conducted with electrochemical work station. Reaction space is limited by chamber volume. So sample capacity, enzyme quantity and reaction area can be controlled accurately. 1) Facture of Sensors No Modified Film former, K3Fe(CN)6 and stabilizer were added in citric acid buffer, mixed equably and printed on carbon paste electrodes; Then modified by appropriate enzyme, aired naturally, made chamber, waited for testing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002031_inmic.2008.4777752-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002031_inmic.2008.4777752-Figure3-1.png", + "caption": "Fig. 3 Justification of the positive sign in 3", + "texts": [ + " The Lambert scheme is an explicit scheme, which generates a suitable trajectory under the influence of an inverse-square-central-force law (gravitational This work was made possible, in part, by Dean's Research Grant awarded by University of Karachi, which is, gratefully, acknowledged. ApPENDIX A: ASTRODYNAMICAL TERMINOLOGIES Down-range error is the error in the range assuming that the vehicle is in the correct plane; cross-range error is the offset of the trajectory from the desired plane. An unwanted pitch movement shall produce down-range error; an unwanted yaw movement shall produce cross-range error. For an elliptical orbit true anomaly,.h is the polar angle measured from the major axis (LPFX in Fig. 3). Through the point P (current position of spacecraft, mPF = r, the radial coor-dinate) erect a perpendicular on the major axis. Q is the intersection of this perpendicular with a circums-cribed auxiliary circle about the orbital path. The angle, LQOF (cf. Fig.3) is called the eccentric anomaly, E. ercosf = a(l-e2)-r r(1 +cos f) = a(1- e)(1 +cos E) (C1) Adding er to both sides and using (Bla) on the right-hand side, one gets For this orbit, pericenter, the point on the major axis, which is closest to the force center (point A in Fig. 3), is chosen as the point at which f = o. Apocenter is the opposite point on the major axis, which is farthest from the force center (point A in Fig. 3). The line joining the pericenter and the apocenter is called the line of apsides. (BIe) (B2e) (B2d) (B2c) (C2) (1 + e)(I- cosE) (1- e)(1 + cosE) l-cosf 1+cosf r(1- cos f) = a(1 + e)(I- cos E) Dividing (C2) by (C 1) Using the identities 1- cos f = 2 sin 2 f 2 1+ cos f = 2 cos2 f 2 with similar results for the expressions (1- cos E) and (1 + cos E) , one sees that the above equation reduces to 2f l+e 2 E tan -=--tan - 2 l-e 2 which implies f J\u00a7+e Etan-=\u00b1 --tan2 l-e 2 Below, it is justified that only positive sign with the radical gives the correct answer. Consider ~ ORF (cf. Fig. 3). One notes that, Subtracting er from both sides and using (B1a) on the right hand side, one gets 1t E 1t -H~E~H =::} --~-~- 222 Further, E ~ 0 =::} f ~ 0; E < 0 ~ f < 0 . Therefore, f and E have the same sign. When 2 2 _!!.. ~ E < 0 tan E < 0 tan f < 0 which implies that 2 2 ' 2 ' 2 ' positive sign with the radical should be chosen. Similarly, E H E f O~-~-, tan-~O, tan-~O 2 2 2 2 and, hence, positive sign with the radical is the correct choice. Two-body problem can be handled elegantly by using the elliptic-astrodynamical-coordinate mesh [14, 15]", + " ApPENDIX C: RELATION CONNECTING ECCENTRIC ANOMALY TO TRUE ANOMALY Some useful relationships among radial coordinate, eccentric anomaly, eccentricity and semi-major axis for an elliptical orbit are listed below: ApPENDIX B: ASTRODYNAMICAL RELATIONHIPS f cosE-e cos =---- 1- ecosE E cosf +e cos =---- 1+ ecosf . f E sinE SIn =---- l-ecosE . E E sinf SIn =---- l+ecosf 3tan f =tan E 2 2 In Appendix C, the last relation is proved and a justi-fication is given for the positive sign taken in front of the square root appearing in the expression for 3. In Fig. 3, semi-minor axis of the ellipse, b, is related to a by b = a E. Do not confuse the point Q in Fig. 3 with the quantity Qdefined in (9). Using the relations r = p l+ecosf and p = a(l- e2) , one may write, r(1 +ecos f) =a(1- e2) . Rearranging REFERENCES [1] H. Goldstein, Classical Mechanics, 2nd ed., Reading, MA: Addison Wesley, 1981, pp. 70-102 [2] 1. B. Marion, Classical Dynamics of Particles and Systems, 2nd ed., New York: Academic, 1970, pp. 243-278 [3] 1. T. Wu, \"Orbit detennination by sol ving for gravity parameters with multi ple arc data,\" Journal of Guidance, Control and Dynamics, vol. 15, pp" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000993_1.3653089-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000993_1.3653089-Figure7-1.png", + "caption": "Fig. 7 Schematic d iagram of response orbit", + "texts": [ + "3929 + iO.1945 -0.9434 - i0.5948J ( a* ) (0.5539) ~ \u00b0 (0.34591 (52) Numerical inversion of equation (55) yields e* = (0.0806 - *0.0541)o- e0a* = (0.01482 - i0.1643) (55) ) (56) This represents an elliptical orbit with major and minor radii equal to 0.239crC and 0.0217 2 rs (see ig. 3 (b)). This, in turn, simplifies the detection of segment\u2013 egment collisions (see below). aximum joint velocity. To achieve a smooth motion of the robot, he change of joint angles between subsequent iterations of the ptimizer is limited (see Fig. 8 (b)). Without load, the selected ervo drives (see 2.1) can reach speeds up to 354\u25e6/s according to pecification. Under load, however, the maximum speed will be ower. At a controller frequency of 20 Hz, the maximum allowable oint angle change between two iterations of the controller was et to \u2206\u03b8min/max = 10\u25e6, which corresponds to an angular velocity f 200\u25e6/s. tability. On flat terrain, stability is usually ensured as long as t least two segments have ground contact. When rolling on teep slopes or when climbing over obstacles, the stability of the ystem must be supervised to prevent forward/backward tilting see Fig. 8 (c)). This is explained in more detail in Section 3.6. he stability measure is represented by the weight-normalized nergy that would be required to topple the robot over. Thus, he constraint is defined by a minimal stability smin without a aximum limit. egment\u2013segment collisions. Due to the high number of DoFs, ollisions between segments can easily occur, e.g., if the robot rosses an obstacle (see Fig. 8 (d)). Therefore, the minimal disances between all segments must be checked to maintain a efined clearance between each other. The two step concept y which the distances between segments are computed, is exlained in Section 3.7.1. The penalty added for the respective onstraint in the sense of Eq. (2) is proportional to the amount of d c c i s 3 d e i t k t r c o t I s j j m n T t m o f 3 m i s p e h( A i i t s p s g e a potential overlap. Also for this case, a minimum limit \u03c7min = min but no maximum limit \u03c7max (in the sense of a maximum learance dmax) is defined. Collisions with obstacles. If one of the segments of the free sequence detects a collision with an obstacle, this likely denotes a low-hanging ledge (see Fig. 8 (e)). However, since the robot can change its shape, it can duck down and move underneath the ledge. For this purpose, the location of the obstacle must be memorized in order not to collide continuously with it. Moreover, the controller might find solutions for the robot shape that do not collide with a memorized obstacle but that enclose the obstacle (see Fig. 8 (f)). Since this is physically impossible, it must be ensured that none of the memorized obstacles is located within the robot area. The obstacle avoidance is explained in Section 3.7.2. For this constraint, the same minimum limit as for the segment\u2013segment collisions is used. Closed kinematic chain. The joint angles must be chosen such that the chain of segments remains closed. Since the segments are mechanically connected, they cannot be ripped apart by the rather weak servo drives. However, a kinematically imperfect solution would result in mechanical tension and thus in an increase of joint torques and energy consumption" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001340_10402000801926471-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001340_10402000801926471-Figure4-1.png", + "caption": "Fig. 4\u2014Lateral view of a misaligned journal.", + "texts": [ + " Equations [4] and [6] will be used in the calculation of the velocities of points Pb and Pj , and it is noteworthy that none of the aforementioned authors provided any kind of description regarding the \u201celliptization\u201d of the journal cross section due to the relative angular misalignment. Except Smalley and McCallion (16), who mentioned the \u201celliptization\u201d and considered it negligible without further analysis, no other comments about the subject were found. So as to make the visualization easier, Fig. 4 shows a particular situation in which the journal is rotated around axis Zj with Ar > 0, so that the major axis of E\u2032 is r \u2032 = r/ cos Ar ; the same holds for a generic rotation \u03b4r defined in Eq. [3]. Thus, in the context of small angles, it is correct to write r \u2032 = r ; likewise, all previous authors seem to have done this arbitrarily. Similarly, other geometric features, such as circumferential out of roundness, were implicitly ignored in the mathematical model. Nevertheless, the consideration of small angles in Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002087_1.5061219-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002087_1.5061219-Figure2-1.png", + "caption": "Figure 2: Principle of the hybrid laser\u2013MIG/MAG welding process [13]", + "texts": [ + " Principles and Applications In many welding applications the occurrence of gaps is hard to avoid. This usually results in the usage of additional filler material. The hot crack affinity of extrusion compound alloys is another reason to use filler material. With regard to those applications, the biggest potential of the hybrid welding technique is expected to be in the area of using additional filler material and thus the combination of laser and MIG/MAG welding which is currently the most preferred laser\u2013arc hybrid welding process [5\u201312, 20]. The fundamentals of the coupled process (see Figure 2) are nearly the same for both, the CO2 laser and the solid state lasers. The laser process and the arc process have a common process zone and weld pool. The process can be controlled in such a way that the MIG/MAG part provides the appropriate amount of molten filler material to bridge the gap or fill the groove while the laser is generating a keyhole within the molten pool to ensure the desired penetration depth. This can be reached at high speed. By combining the laser beam and the arc, a larger molten pool is formed compared to the laser beam welding process" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003572_s10015-021-00694-y-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003572_s10015-021-00694-y-Figure3-1.png", + "caption": "Fig. 3 Folding mechanism for 1 pair", + "texts": [ + " The combination of wall-press type and the crawler type are used to achieve an expandability and mobility. There are 3 pairs of folding mechanisms, all 3 pairs are driven by a motor (24\u00a0V 22\u00a0W 52.1mNm 4600\u00a0rpm), the rotation is transferred through the gear train (gear ratio 1875:1) to drive each linkage. By this way, the robot can be adjusted the diameter and contact force. Thanks to the worm gear, the mechanism be able to lock itself automatically, while the motor is turned off. The mechanism for each pair is illustrated in Fig.\u00a03. 1 3 The relationship between the link angle (\u0275) and robot diameter (Dr) are expressed by Eq.\u00a0(1). Where origin diameter (Do) is 130.2\u00a0mm, length of link (l) is 90\u00a0mm and link angle (\u0275) are 5\u00b0 \u2264 \u0275 < 90\u00b0: Using the current control algorithm. The robot achieves a self-locking (S-L) feature to keep the contact force and lock itself with the pipe, while the manipulator is operating. The pipe is setup in horizontal plane. The inner surface of the pipe is slippery and has low friction. Therefore, only normal force occurs at the contact point" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000854_ac60232a014-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000854_ac60232a014-Figure4-1.png", + "caption": "Figure 4. Absorbance of silver bromide suspensions as a function of concentration for KBr solutions of known concentration", + "texts": [ + " Total bromine : ,4 calibration curve was prepared using several aliquots of a stock solution of KBr (ACS, 99.0%) containing 18.6 mg. per liter. Aliquots were added to 25-ml. volumetric flasks to each of which was added 1 ml. of Na2S03 solution (1 gram/100 ml.). After mixing and diluting to approximately 20 ml., 1 ml. of solution (1 gram/100 ml. of 1X HNOa) was added to each. The solutions were made up to 25 ml. and aged 1 hour before measurement of the absorbance a t 500 mM. h plot of these data is shown in Figure 4. To establish their concentrations, similar measurements of total bromine were made on 20-ml. aliquots of a series of Brz solutions which contained up to 10 drops of saturated (room temp.) bromine-water per liter. Aliquots of these same Bra solutions were used for measurement of the starting concentration of oxidant by the double extrapolation as described below, thus giving the required data for standardizing and relating the two analytical approaches. Oxidant (HBrO) was determined on each of the above series of Br2 solutions as follows: 5 ml", + " The averaged values of absorbance were plotted as a function of time and the curve extrapolated back to the time of preparation of the o-tolidine Brz mixture. By repeating this procedure at hourly intervals (usually 4) after preparing a Brz solution, a curve showing the change in oxidant concentration with time was obtained and extrapolated back to find the initial oxidant concentration (in terms of absorbance) at the time of preparation of the Brz solution. VOL. 37, NO. 13, DECEMBER 1965 1691 RESULTS AND DISCUSSION The calibration curve of Figure 4 shows a small amount of curvature, whereas Figure 3, which is plotted from essentially the same kind of data, is a straight line. No reason can be given for the difference in behavior, unless perhaps there was a residual, uncompensated dependence of absorbance on aging time. The curve of Figure 4 does not pass through zero a t the Data obtained from Figurer 4 and 5 abscissa, which indicates that probably some AgBr particles were too small to scatter effectively a t 500 mp. The same reasoning applies to Figure 3. The solubility of AgBr a t 20\u2019 C. is given by Hillebrand and Lundell (6) as 0.107 rng./liter which can account for about a third of the unmeasured AgBr. The points on Figure 4 are of single (not replicate) samples and give an estimated reproducibility of about 0.001 absorbance unit (about 0.035 mg./ml.). The accuracy of weighing for the preparation of the KBr for the solution of known concentration was about 10.25%. The colorimetric sensitivity for the determination of the o-tolidine oxidation product was considerably higher than that for the determination of AgBr suspensions. This was partially compensated for by using 20-ml. aliquots of the dilute Brz solution diluted to 25 ml" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002752_012051-Figure11-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002752_012051-Figure11-1.png", + "caption": "Figure 11. Dynamic model of two gear pairs in planetary set [7].", + "texts": [ + " This model is presented as a nonlinear system with 6 degrees of freedom, consisting of a gear pair, represented by solid disks with masses m1 and m2 and moments of inertia J1 and J2, respectively. The meshing of the teeth is described by an elastic-damping connection with variable stiffness kz(t) directed along the line of action of the force. The bearing arrangements of the toothed shafts are characterized by the bearing stiffness kbx and kby in accordance with the directions of the selected orthogonal coordinate system for each gear. The system is balanced by torque applied in opposite directions to the gears. Figure 11 shows a schematic diagram of the dynamic model used. 1 - ring gear, 2 - sun gear, 3 \u2013 planetary gear, \u03b1 - pressure angle. The combination of analytical methods and the finite element method used in the dynamic model made it possible to obtain accurate estimates of the dynamic forces in gears at low time costs, achieved by simulating the process of gearing of gears in the finite element method and taking into account the possible loss of contact between the teeth. Since only a pair of gears was considered in the calculation, changes related to the kinematic scheme of the gearbox were introduced in the used dynamic model - there is no elastic-damping connection with variable stiffness ke(t)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001125_aero.2007.352851-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001125_aero.2007.352851-Figure2-1.png", + "caption": "Figure 2. General formation control configuration", + "texts": [ + " This controller is used for helicopters marching in a single file (for example in a Figure 3. 1- a control configuration line formation) or at an edge ofthe formation geometry. Note that the I - a controller alone cannot uniquely define a general formation. When a helicopter is constrained by more than one neighbour in the formation, a second controller is needed to control the three-dimensional distances ofthe helicopter from two neighbouring helicopters. This controller is called the I - I controller. These two local controllers are necessary to define a general formation (Fig. 2). Usually the helicopters at an edge of the formation geometry control their distance with their immediate front helicopter using the I - a controller. The other helicopters control their distances to their immediate front and side helicopters using the I - I controller. This is necessary so that a helicopter can also avoid its side helicopter. More details about the possible formation structures based on the graph theory can be found in [2]. Design ofthe I - a Controller In Fig. 3, a system oftwo neighbouring helicopters in the formation is shown" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001092_1.2785577-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001092_1.2785577-Figure2-1.png", + "caption": "Figure 2. (a) Illustration of nematic molecular orientations (director field) influenced by a rough surface. (b) The pretilt angle of a liquid crystal depends on its chemical structure and on the surface\u2019s composition and morphology.", + "texts": [ + " We have also defined a statistical measure of surface morphology that clearly indicates the surface asymmetry that is likely responsible for pretilt, a vector property of the SiO2 surface. 2. Alignment direction and pretilt The relative roles of surface roughness anisotropy and molecularscale anisotropy (e.g. polymer chain orientation) in determining liquid crystal alignment depends on the specific surface under consideration and is sometimes uncertain [7]. In the case of amorphous SiO2 we start with the assumption that surface morphology is the dominant mechanism. In the simple model of a nematic liquid crystal in contact with a rough surface (Figure 2(a)), we assume that the bulk liquid crystal is characterized by a uniform director field n that is parallel to the mean plane of the surface. Close to the surface the director field becomes increasingly non-uniform; we assume strong anchoring such that molecules in contact with the surface lie flat (planar, homogeneous alignment) and that the local direction n is parallel to the surface. The elastic energy density per unit area of surface F is usually calculated for the one-dimensional case of only a single spatial frequency [8, 9], e", + " This model assumes that variations in the direction of n with position are slow and confined to a plane containing the z-axis and \u2329n\u232a. Energy density F is minimized by orienting the liquid crystal director field \u2329n\u232a along the direction of least roughness. In this SID 07 DIGEST \u2022 1405ISSN/007-0966X/07/3802-1405-$1.00 \u00a9 2007 SID approximation F depends on the spatial frequency spectrum of the surface but not on the surface shape (i.e. it doesn\u2019t depend on phase relationships between frequency components). Another important property of LC-surface interactions, especially for FLC alignment [3, 4], is pretilt (Figure 2(b)). Molecular orientation at a surface can vary from perpendicular (90\u00b0 pretilt) to parallel (0\u00b0 pretilt). An isotropic surface can induce non-zero azimuthally degenerate pretilt. On anisotropic surfaces (e.g. rubbed or obliquely deposited) the pretilt generally exhibits a vector character, i.e. tilting molecules toward or away from the surface\u2019s easy axis direction (e.g. toward or away from the rubbing direction). The FLCs of interest here posses an I-N-SmA-SmC phase sequence upon cooling; the SmC layer normal is generally expected to be parallel to the director orientation established during the nematic phase" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001848_robot.2007.363783-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001848_robot.2007.363783-Figure1-1.png", + "caption": "Fig. 1: Manipulator Configuration and Coordinate Description", + "texts": [ + ". INTRODUCTION ue to their high stiffness, high speed, large load carrying ability and high precision positioning capability, parallel mechanisms have become very popular in the past decade. There are numerous parallel manipulators with different structure, different way of actuations and different number of the degree of freedom (DOF). The parallel manipulator proposed in [1] which is the subject of this paper as shown in Fig.1, has a different structure than the others. The specific feature of this manipulator is that with only two legs it has 6 DOF. Fewer legs leads to smaller required space for manipulator\u2019s installation, decreases the chance of the leg\u2019s collision during maneuver, and also means fewer moving part. However, although the manipulator has the above advantages, it suffers from the small load carrying capacity which is a direct consequence of its actuation. It is to be noted that, the manipulator has two base-mounted spherical actuators", + " The different possible numbers of solution for the inverse and forward pose kinematic analysis are given. A closed form solution for the rate kinematic is proposed. The two different categories of the singular points for the new mechanism with their geometrical interpretation are introduced. It is worth noting that in one category the mechanism loses one or more DOF while in the other one it gains one or more DOF. The parallel manipulator presented in this article consists of a moving platform connected to the base frame by two legs as shown in Fig. 1. Each leg is composed of the spherical (S), prismatic (P) and universal (U) joints, which is called a SPU leg. These joints in a serial manner construct each leg. Although, it consists of two legs, it has 6 DOF[1]. Moreover it should be mentioned that the actuators of the mechanism are at the spherical joints and are base-mounted. To specify the location and orientation of the moving platform with respect to (w.r.t) the base, a coordinate frame is attached to the moving platform in which its origin is at the center of the platform. The moving platform coordinate and the base coordinate are schematically shown in Fig. 1. The transformation matrix of the moving platform D 1-4244-0602-1/07/$20.00 \u00a92007 IEEE. 175 coordinate w.r.t the base frame coordinate using X-Y-Z Eulerian angles of Rotation is =TB P \u2212 \u2212+ +\u2212 1000 zCCSCS ySCCSSCCSSSCS xSSCSCCSSSCCC \u03b3\u03b2\u03b3\u03b2\u03b2 \u03b3\u03b1\u03b3\u03b2\u03b1\u03b3\u03b1\u03b3\u03b2\u03b1\u03b2\u03b1 \u03b3\u03b1\u03b3\u03b2\u03b1\u03b3\u03b1\u03b3\u03b2\u03b1\u03b2\u03b1 For the inverse and forward pose kinematic analysis, the standard Denavit-Hartenberg parameters are required. The selected intermediate coordinates based on the DenavitHartenberg notation are shown in Figure 2. The spherical, prismatic and universal joints need three, one and two coordinate frames, respectively", + " IV. INVERSE POSE KINEMATIC In the inverse pose kinematic the desired actuator variables which are the spherical joint variables, should be calculated having TB P . First the basic idea which leads to the calculation of the active joint variables with no need for the evaluation of the passive joint ones will be discussed and later its equivalent mathematical repression will be introduced. Having the platform\u2019s location and orientation in inverse pose process, location of the points A\u2032 and B\u2032 in Fig. 1 are known. Thus, the leg\u2019s direction is available. Therefore, two spherical variables 21 ,\u03b8\u03b8 which are related to the direction of the leg as shown in Fig. 3 can be evaluated. To have 3\u03b8 , rotation of the leg around itself has to be found. From Fig. 1, it is clear that the direction of 5z , solely depends on the rotation of the leg around itself. However, 5z for each leg is perpendicular to its corresponding 6z and leg. Since the direction of leg and 6z is known, the direction of 5z will be available by cross producting 6z and the leg\u2019s direction. It is to be noted that, the direction of the leg is available having 21 ,\u03b8\u03b8 . Also, the 6z direction is the same as the direction of the normal vector of the manipulator which is known. In the following the mathematical equivalent of the above geometrical interpretation is provided", + " The same procedure can be used to have the active variables of the right leg, except that TT BP 06 , for the right leg are: \u2212 \u2212 \u2212 = 1000 0100 0010 5.001 )( 0 B RBT \u2212 = 1000 0100 0010 5.001 )( 6 P RPT Since there are four possible solutions for each leg, there exist sixteen possible answers for the inverse problem of the introduced manipulator. The forward pose problem for the majority of the parallel robots is solved through the numerical analysis and rarely a closed from solution is available. However, in this paper a closed form solution for the forward pose problem of the proposed manipulator is presented. Consider Fig.1 since the spherical joint\u2019s variables are given, the direction of leftZ )( 5 and rightZ )( 5 axis are available. Since left B Z )( 5 and right B Z )( 5 always lay on the moving platform surface, their cross product, m , gives the normal direction of the platform. Moreover, line BA \u2032\u2032 which connects point A\u2032 of the left leg to point B\u2032 of the right leg, placed on the moving platform surface. Thus: =m left B Z )( 5 \u00d7 0)( 5 =\u2032\u2032\u22c5\u21d2 \u2192 ABmZ right B (9) or [ ] [ ] [ ] [ ] \u2212= \u2212+= ++= =++\u21d2=\u2032\u2032\u22c5 \u2192 1 23212211 23212211 00 mG SmCSmCCmE SmCSmCCmH GEHABm b RRRRR fffff Rf \u03b8\u03b8\u03b8\u03b8\u03b8 \u03b8\u03b8\u03b8\u03b8\u03b8 where, the relation for the \u2192 \u2032\u2032AB is given in Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002945_978-3-030-69178-3_15-Figure16.4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002945_978-3-030-69178-3_15-Figure16.4-1.png", + "caption": "Fig. 16.4 Key parameters of KUKA KR6 R700 Sixx [24]", + "texts": [ + " Its energy supply system and controller are integrated within the KR C4 controller. A Windows PC is also utilised as a platform to design and test the graphic interface and also to serve as a broker computing node connecting the AR environment and the robot. The hardware is connected using a TCP/IP connection through Ethernet ports. Unfortunately, KUKA does not support complete 3D models of industrial robots. Thus the virtual robot arm was re-created in SolidWorks based on the parameters from KUKA technical file as shown in Fig. 16.4. As Unity only supports FBX data format during importing, the assembly SolidWorks data is first imported to 3D Max for conversion into FBX format, and then transferred to Unity. The proposed system is implemented in a real industrial robot cell to evaluate the feasibility and performance. The deployment structure of the proposed system is F ig .1 6. 5 D ep lo ym en ts tr uc tu re shown in Fig. 16.5, which is compliant with the system structure design introduced in the previous section. In this research, Microsoft HoloLens is utilised as the AR device to host the interactionmodule" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000360_4-431-27901-6_2-Figure2.10-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000360_4-431-27901-6_2-Figure2.10-1.png", + "caption": "Fig. 2.10a,b. Inclined pad and journal bearings (2). a Fig. 2.9a with U1 = U and U2 = 0, b Fig. 2.9c with U1 = 0 and U j = U", + "texts": [ + " For the inclined pad bearing, 6U(\u2202h/\u2202x) is directly obtained from the wedge effect. For the journal bearing, 6U(\u2202h/\u2202x) is interpreted as the sum of the wedge effect, which is negative, and the squeeze effect (the effect of V2), which is positive with a magnitude twice that of the wedge effect. Although it can be said from this analysis that the mechanisms of pressure generation in an inclined pad bearing and a journal bearing are different, this is not necessarily true. Figure 2.9a and 2.9c can be redrawn as Fig. 2.10a,ba and 2.10a,bb; considering case 2, it is clear that Fig. 2.10a,bb is equivalent to 2.10a,ba if viewed upside down. Therefore, the difference described above is only an apparent one. In the case of journal bearings, the circumferential speed of the journal U2 is along the inclined surface, and Eq. 2.20 becomes: \u2202 \u2202x ( h3 \u2202p \u2202x ) + \u2202 \u2202z ( h3 \u2202p \u2202z ) = 6\u00b5 [ (U1 + U2) \u2202h \u2202x + 2V2 ] (2.24) It is this form of the equation that appears in Reynolds\u2019 original paper [3]. b. Stationary Surfaces and Moving Surfaces The upper surface of Fig. 2.10a,ba and the lower surface of Fig. 2.10a,bb are equivalent as stated above, and are called the stationary surface. Similarly, the lower surface of 2.10a,ba and the upper surface of 2.10a,bb are equivalent, and they are called the moving surface. When 2.10a,ba is regarded as an inclined pad bearing, the upper inclined surface is usually stationary and the lower surface is moving, in agreement with the above nomenclature. However, the meanings of \u201cstationary\u201d and \u201cmoving\u201d here are more fundamental, i.e., the stationary surface is a surface in which the distance from a fixed point (on the stationary surface) to the mating surface does not change, and the moving surface is a surface in which the distance from a fixed point (on the moving surface) to the mating surface does change" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002199_j.matpr.2020.12.064-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002199_j.matpr.2020.12.064-Figure1-1.png", + "caption": "Fig. 1. Completed CAD assembly.", + "texts": [ + " After the design phase is one among the important phases in the construction of a prototype, in which through many modifications are made to produce an efficient design and it is chosen for the build of prototype. The CAD designs are made using Solid works software. The design of gantry which supports the binder extrusion head along with the powder levelling roller which helps in the even drag of powder from the build platform to the feed platform through continuous rolling motion of roller. The feed and built platforms help in feeding and building the Green part respectively. The Fig. 1 shows the full assembly of the essential components and incorporates the major mechanisms including the XY plotter mechanism, Z-axis platforms (feed and build platforms), collecting bin, powder levelling roller, binder extrusion head. Two lead screws are used for the dynamic motion namely the up and downward motion of the feed and build platform. Hence a lead screw with an 8 mm diameter and with a pitch of 4 mm is chosen. Outer diameter = 20 mm Pitch = 4 mm Length = 800 mm No. of start = 1 Lead = 4 mm (Lead = pitch no" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001288_dscc2008-2285-Figure9-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001288_dscc2008-2285-Figure9-1.png", + "caption": "Fig. 9. Kinematic scheme of double planetary gear.", + "texts": [ + "url=/data/conferences/dscc2008/70308/ on 02/24/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use i i c 2 2 1 1 Fig. 7 shows the kinematic scheme of the bevel gear differential (cf. Figs. 1a-1c). The corresponding bond graph model is shown in Fig. 8a, and a more convenient transformed form is given in Fig. 8b [7,8]. Note that the differential model (with the bevel gear ratio i = 1) may be regarded as a special case of the planetary gear model, where the gear ratio is h = 1 (cf. Figs. 6a and 8a). The double planetary gear kinematic scheme is shown in Fig. 9 (cf. Fig. 1c). The corresponding bond graph model and its equivalent counterparts are shown in Fig. 10. Based on the ALSD kinematic scheme in Fig. 1a, and the passive differential bond graph in Fig. 8b and the clutch bond graph in Fig. 3, the ALSD bond graph is easily composed as shown in Fig. 11. Applying the straightforward bond graph modeling formalism (see Appendix and [6]) gives the following model equations: 2 21 * 22 * * * 21 2 2 fc ffifcc c i Rearranging the above torque equations yields 22 22 2 1 fi fi i i , (1) 3 Copyright \u00a9 2008 by ASME Downloaded From: http://proceedings" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003132_978-981-16-0119-4_20-Figure58.5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003132_978-981-16-0119-4_20-Figure58.5-1.png", + "caption": "Fig. 58.5 Conceptual sketches of laser cutting machine ( Source Author)", + "texts": [ + ", it was also equipped with rapid prototyping tools like CNC, laser cutting, 3D printer, and MCB printer. Fab-Lab thus proved an ideal place to develop the project. The prototype was designed and after a practical research it was finalized that the bamboo will rotate as per the design and the laser could move in single plan, i.e. parallel to the ground. The conceptualization started through sketches, but that was not enough as imagining the product and its working was not that easy for the producers (See Fig. 58.5). Prototyping Mentors at Fab-Lab, Nagpur guided students to use \u201cFusion 360\u201d, an integrated software built to bridge designer and producer. The project was first created virtually in the software clearing the details in the design of outer body and the machine (See Fig. 58.6). Frommajor working like access toworking spacemachine, bamboo fixing, etc. To smaller details like display and interaction panels, the 3Dmodel became the reference for prototyping. With use of power tools, the body of machine were made with ply and acrylic" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001844_cp:20080483-Figure12-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001844_cp:20080483-Figure12-1.png", + "caption": "Figure 12. Locus of input and load voltage vectors under unbalance", + "texts": [ + "0 and the parameters of the circuit are given in Appendix A. supply voltages: (a) Standard MC with passive compensation, (b) proposed hybrid MC&VSI (20% unbalance). 4.4 Operation under unbalanced supply conditions The effectiveness of compensating voltage unbalance of 20 % without any loss in the output voltage generation capability of a hybrid MC&VSI is shown in Fig. 11 and 12. Two strategies are presented: Strategy 2.1 - ZV are completely removed in MC Strategy 2.2 - ZV are used in MC The locus of input and load voltage from Fig. 12 shows that the hybrid MC&VSI is able to deliver balanced voltage to load unaffected by the heavy unbalance of the supply voltage which is revealed by the fact that the load voltage exceeds the supply voltage in areas where the supply voltage is minimum. (9) VTR 2 Nmain + k v-aux . Naux=J; kVAout A comparison of the installed power in the switches based on the assumption that the supply voltage and the load current remain the same is presented in Table I. The power installed in the switches increases for the hybrid approach but in a greater proportion so does also the output power, which means that the cost of processing 1 kW of power is slightly smaller for the hybrid MC&VSI than for a standard MC" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002838_012021-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002838_012021-Figure4-1.png", + "caption": "Figure 4.(Step3)", + "texts": [ + " HSLA Steel Properties of selected materials for the analysis of rim as shown in Table 1. 1.2. Modeling of Rim FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 Rim was designed by using 3D modeling software solid works, 2019 version [14]- [15]. Various steps used to create the 3D model of rim modelling. As figure 2. indicates the flange part of the rim which is a design by using revolve command whereas figure 3. illustrate the flange part with the hub designed by using extrude command and Figure 4. gives the complete design of the rim and its cuts by using extrude cut command. The paper aims to analyze different properties in different materials selected for the rim. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 2. Finite Element Analysis of Rim: 2.1. Boundary Conditions Fixed support is given to the five stud holes because the wheel fixed upon axle through bolts on stud holes. By considering the diameter of the rim as 482 mm, the rotational velocity of the rim considered as 4000 rpm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000869_piee.1965.0259-Figure9-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000869_piee.1965.0259-Figure9-1.png", + "caption": "Fig. 9", + "texts": [ + " From our analysis, both the magnetic flux-density distribution due to the field dependent on the load current, and part of that due to the field dependent on the voltage, agree with the assumptions of the travelling-field theory. This justifies reference to that theory when interpreting the phenomena occurring in the meter. When the conditions of the travelling-field theory are fulfilled, the magnetic fields in the meter air gap, whose amplitude distribution agrees with the assumptions of the theory under consideration, form a forward-travelling field in the form of a sine wave running in the same direction as the rotor, within the longitudinal limits of the air gap. The idealised model shown in Fig. 9 leads to another, maybe more directly convincing, way of forming such a travelling field. This model consists of a stator, and two rotors which revolve synchronously with a uniform speed in the directions marked with arrows (Fig. 9). To the circumference of each rotor, identical permanent magnets are attached, with their poles alternating. The appropriately shaped air gap of the stator contains a disc rotor, just as in a meter; a portion of the disc rotor is acted upon by the field due to the permanent magnets as they pass. The similarity, between the travelling field thus produced and the forward-travelling field in the air gap of a meter, increases the greater the distance of the revolving magnets from their centre of rotation compared with the length of the segment of the disc rotor upon which the travelling field acts. 1599 On comparing the concept of the travelling field as illustrated in Fig. 9 with that of References 1 and 2, in the case of the field of a pair of magnets revolving coaxially with the disc rotor, one immediately becomes aware of the essential Conceptual model of device for producing the forward-travelling field in a meter (i) Stator (ii) Rotors with permanent magnets (iii) Permanent magnets (iv) Disc rotor of meter difference between the two interpretations of the travelling field in the meter. This is because the papers quoted identify the travelling field in a meter as the field which rotates coaxially with the disc rotor, whereas our analysis of the magnetic fields provides evidence only for the existence of the travelling field within the limits of the air gap" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000527_978-3-540-79016-7_13-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000527_978-3-540-79016-7_13-Figure4-1.png", + "caption": "Fig. 4. A gravity-flow tank/pipeline system", + "texts": [ + " The second-order sliding mode occurs if s = s\u0307 = 0 and the sufficient condition s\u0307 = \u2212\u03b2|s|p \u2212 \u03b1 \u222b sgn(s)dt (48) where \u03b1, \u03b2 > 0 and 0 < p \u2264 0.5, is satisfied [4]. The following control law satisfies the condition (48) u = 1 Gn(y, y\u0307, . . . , y(n\u22121)) ( \u2212 n\u22121\u2211 i=1 kiy (i) \u2212 Fn(y, y\u0307, . . . , y(n\u22121) \u2212 \u03b2|s|p \u2212\u03b1 \u222b sgn(s)dt) \u2212 \u03c6T n (y, y\u0307, . . . , y(n\u22121))\u03b8\u0302 ) (49) A gravity-flow/pipeline System is a liquid system in which the water supply is higher than all points in the pipeline and no pump is normally required (see Fig. 4). It is assumed that the flux cannot be reversed. Consider the following gravity-flow/pipeline system including an elementary static model for an \u2018equal percentage valve\u2019 [39] x\u03071 = Apg L x2 \u2212 Kf \u03c1A2 p x2 1 x\u03072 = 1 At ( FCmax\u03b1\u2212(1\u2212u) \u2212 x1 ) + \u03b8f(x1, x2) (50) with x1 : volumetric flow rate of liquid leaving the tank x2 : height of the liquid in the tank FCmax : maximum value of the volumetric rate of fluid entering the tank g : gravitational acceleration constant L : the pipe length Kf : friction of the liquid \u03c1 : density of the liquid Ap : cross sectional area of the pipe At : cross sectional area of the tank \u03b1 : rangeability parameter of the value u : control input, taking values in the closed interval [0, 1] \u03b8 : an unknown parameter f(x1, x2) : a known perturbation function depending on the waves produced by entering the liquid" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002772_978-3-030-67750-3_14-Figure30.7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002772_978-3-030-67750-3_14-Figure30.7-1.png", + "caption": "Fig. 30.7 Side vie of the wheelbarrow loader bucket", + "texts": [ + " The simple 2D sketches equipped with the specifications and parts from the data collected in the previous design method process had been analyzed later by using a weighted objective analysis method. Therefore, the final design concept is available after finishing up all the designs steps. The final conceptual design was made by using SolidWorks, Fig. 30.6a displays the final conceptual design of the wheelbarrow loader bucket, while Fig. 30.6b is the exploded view of the design. The side view of wheelbarrow bucket loader is shown in Fig. 30.7. The design is based on the existing wheelbarrow but has a modification by using the six-point linkage based on the existing loader bucket design. The pneumatic actuator attached at the loader linkages is to make ease the loading and unloading process of the bucket. Themainmaterial used for the loader bucket is using Aluminiumwhile the wheelbarrow body is maintains to use the existing material, which is steel. The wheelbarrow loader bucket is attached to a loader linkage that is attached to shafts that connect to a linear actuator that acts as the arm for the bucket to move downwards or upwards" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002167_jae-201576-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002167_jae-201576-Figure1-1.png", + "caption": "Fig. 1. Main MG structure: (a)\u00a0Concentric, (b)\u00a0harmonic, (c)\u00a0planetary.", + "texts": [ + " All rights reserved magnetic gear (MG) offers exciting features such as low acoustic noise, minimum vibration, maintenancefree, improved reliability, inherent overload protection and no physical contact between the input and output shafts. Besides, MG does not produce debris and does not require lubricant and is, therefore, suitable to be used in harsh environments\u00a0[3\u20135]. Three main MG structures were developed in early 2000; concentric, harmonic and planetary, using rare-earth to improve the torque density\u00a0[6\u20139]. Figure\u00a01 shows the three distinct MG structures. Ferrite magnet does not acquire enough flux density remnants to be competitive in this industry. Torque density comparison of different rare-earth and Ferrite shows that Neodymium Iron Boron (NdFeB) delivers the highest torque density nearly 15 times that of ferrite\u00a0[10,11]. If cost-effectiveness is emphasized in CMG, Ferrite may be a better choice. Still, for our EVs application, however, without rare earth design in the design of the magnetic gear, the torque density could not match its mechanical gear counterpart\u00a0[12,13]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002266_j.ymssp.2020.107484-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002266_j.ymssp.2020.107484-Figure3-1.png", + "caption": "Fig. 3. Coupled system consisting of front-axle and steering subsystems. Rims and tires are not shown explicitly, although these parts are included at the front-axle subsystem during the analysis [9].", + "texts": [ + " It is the aim of this work to break those targets down to the corresponding characteristics of the main subsystems, which are part of the overall transmission path. The required subsystem properties must be useful for the parallel development of the substructures throughout their design and verification process. For a completed vehicle, i.e., at the global system level, the transmission path of mechanical vibration starts at the contact patch of the tires on the road and ends up at the steering wheel, as shown in Fig. 3. Fluctuating tire forces at the front left and/or front right wheel are propagated via the respective tie rod to the steering rack. At the steering rack, the assisting force of the electric servo line is added to the steering wheel torque according to a specific control strategy. The effective rack force is transmitted by the rack and pinion gear to the upper steering column up to the steering wheel. In order to separate the steering gear from the front axle, the connection of the rack to the tie rods is chosen as a mechanical interface between the two subsystems. Assuming the rack bar, on which the tie rods are acting, to be rigid, the left and right tie rod force can be summarized to only one effective interface force between the front axle and the steering. In addition, we assume that the right and left halves of the front axle behave identically, which is indicated in Fig. 3, which shows the left half of the front axle grayed out. This simplification is justified by the assumption that the small masses and high stiffnesses of the ball joints and tie rods have a low impact on the global system dynamics up to 30 Hz. Regarding the wheel assemblies, please note that the rims and the tires are not shown owing to better visibility though they are part of the frontaxle subsystem. Fig. 4 shows each of the subsystems represented by one mechanical four-pole element connected in series to the other" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003132_978-981-16-0119-4_20-Figure78.1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003132_978-981-16-0119-4_20-Figure78.1-1.png", + "caption": "Fig. 78.1 Conceptual design of propeller with saw tooth cut at one TE and one LE", + "texts": [], + "surrounding_texts": [ + "enhancement technique and its positions on the propeller for this current work [9]. While coming to the methodology section, the two engineering approaches have beenused to test the aerodynamic behavior on the relevant components. The methodologies are experimental test andComputational FluidDynamics (CFD)based simulation, in which most of the aerodynamic analyses were executed with the help of CFD simulations. Because of these huge implementations, the pre-processor steps are available to everyone [11]. Especially, the boundary conditions such as type of solver, type of turbulence model, quantity and quality of turbulence model are easily available data to the new researchers. The high amounts of analyses were used velocity inlet to their problems. Thus, with these inputs, the CFD simulations are executed in these comprehensive aerodynamic analyses on various propellers [13]. b. Solution Techniques \u2013 CFD Analysis In this work, analyses the comparative aerodynamic performance on a UAV\u2019s propeller by using CFD tool, i.e., Ansys fluent. The fundamental aim of this work is to select the suitable lift enhancement technique for UAV\u2019s propeller. In this regard, six different designs are modeled, in which five conceptual designs are comprised of a propeller with edge modifications, and the other one is the conceptual design of base propeller. The techniques implemented in the 5-inch diameter propeller are curvy cut saw tooth cut, aero cut, etc., and in general, the profile modifications in the UAV\u2019s propeller is executed for noise reduction. From the literature survey, it was clearly understood that the noise induced due to the abnormal environment is reduced [1, 2]. Nowadays, UAV industry needs a quite propeller so the design modifications-based propellers are suggested a lot for the construction of quite UAV. But the problem along with these types of profile modified propellers may have a chance to generate low-aerodynamic forces. Hence, the conduction of an integrated study is very important in the updated propellers to increase the implementation of the UAVs in real-time applications. The steady- and pressure-based turbulent flow is used as fundamental behavior to the working fluid for all these analyses. The aerodynamic parameters such as lift, drag, CL, CD are used as selection parameters for this comparative analysis [15, 16and17]. c. Conceptual Design of various propellers UAV\u2019s propeller and its designs are the key role of this aerodynamic performance investigation. From the previous work, it was understood that propeller with saw tooth cut has been provided the low turbulence noise than base propellers. Thus, in this work, the aerodynamic performance of low acoustic profiled propellers is computed, in which the conceptual design of all the propellers are used from the literature survey [18, 19,and20]. Figures 78.1, 78.2, 78.3, 78.4, and 78.5 are revealed the conceptual design of propellers with saw tooth cuts, in which the locations of the saw tooth cuts are formed at the various edges of the propellers such as both leading edges, both trailing edges, one leading edge cum one trailing edge, only one trailing edge. Apart from these saw tooth cuts, the one more relevant cut is located at the leading edges of the propeller, which is v-cut. 78 Comparative Aerodynamic Performance Analysis on Modified \u2026 973" + ] + }, + { + "image_filename": "designv11_83_0002191_j.matpr.2020.12.112-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002191_j.matpr.2020.12.112-Figure3-1.png", + "caption": "Fig. 3. Meshed Model.", + "texts": [ + " This point bargains about the compelled part appearing of the posts and imitating them by contributing specific physical representations and limit situations to reenact the fascinating situation plans ended likely. The segment stayed demonstrated utilizing ANSYS 14.0 software Plan Modeler programming as a strong perfect and ANSYS 14.0 worktable was utilized aimed at the assess- ment in the static examination. The key work method of a Limited Component Examination system is tended to like in Fig. 1(A) Fig. 2 Fig. 3 Fig. 4 Fig. 5 Fig. 6 Fig. 7 Fig. 8 Fig. 9 Fig. 10. The stainless harden fabric property Bulk (q): 7950 kg/m3 Young\u2019s modulus (E):206000Mpa Poisson\u2019s ratio: 0.3 The CAD typical of the bar is fit hooked on a limited quantity of components utilizing ANSYS 14.0 software inherent lattice calculation. The pillar contact locale is fit and interlinked to empower estimation of the power collaboration between them limited component investigation or FEA representing to a genuine task as a \u2018\u2018work\u201d a progression of little, consistently formed tetrahedron associated components, as appeared in the above fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002654_s42417-021-00299-6-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002654_s42417-021-00299-6-Figure4-1.png", + "caption": "Fig. 4 Cantilever beam model of the helical gear tooth", + "texts": [ + " (3) { u\u0308 + 2\ud835\udf0911u\u0307 + 2\ud835\udf0912v\u0307 + \ud835\udf0511(\ud835\udf0f)g(u) + \ud835\udf0512(\ud835\udf0f)g(v) = f1 + e\u03081(\ud835\udf0f) v\u0308 + 2\ud835\udf0921u\u0307 + 2\ud835\udf0922v\u0307 + \ud835\udf0521(\ud835\udf0f)g(u) + \ud835\udf0522(\ud835\udf0f)g(v) = f2 + e\u03082(\ud835\udf0f) g(x) = \u23a7\u23aa\u23a8\u23aa\u23a9 x \u2212 1 0 x + 1 x > 1 \u22121 < x \u2264 1 x \u2264 \u22121 , and g(x) is the dimensionless nonlinear backlash function. Meshing stiffness is a time-varying parameter, and it changes along with the meshing position. The slice-integral method is an efficient calculation method widely used by researchers to calculate the time-varying meshing stiffness of helical gears. The helical gear is approximated as a series of independent spur gear slices whose face width is relatively small along the axial direction, and each slice can be regarded as a cantilever beam model, as shown in Fig.\u00a04. The accurate time-varying meshing stiffness of the helical gear pairs can be obtained by accumulating the stiffness of the sliced spur gear pairs [21, 22]. The time-varying meshing stiffness of the gear pairs includes five parts, bending stiffness kb , shearing stiffness ks , axial compressive stiffness ka , Hertzian contact stiffness kh , and fillet-foundation stiffness kf. Equations (4\u20138) show the computing equations on the above five parts of the stiffness [23\u201326]: (4)kb = N\u2211 i=1 1 /( \u222b 3 \u2212 \ufffd 1 3 ( 2 \u2212 ) cos { 1 + cos \ufffd 1 [( 2 \u2212 ) sin \u2212 cos ]}2 2E\u0394l [ sin + (( 2 \u2212 ) cos )]3 d ) (5) ks = N\u2211 i=1 1 /( \u222b 3 \u2212 \ufffd 1 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002563_012068-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002563_012068-Figure3-1.png", + "caption": "Figure 3. BLDC motor construction", + "texts": [ + " From Figure 2 there are 3 axes (x, y, z), each of which has the following functions: Axis x Surge The ROV moves forward / backward in the direction of the x axis Roll The ROV rotates about the x axis Y axis Sway The ROV moves sideways along the y axis Pitch ROV rotates about the y axis Sumbu z Heave OV moves up / down following the z axis Yaw ROV rotates about the z axis To be able to drive the ROV and convert energy from electricity to force, a BLDC motor equipped with a propeller is used. Basically, the BLDC motor works by using the principle of the attractive force between two magnets with different poles or the repulsion between two magnets with the same poles [4]. The rotor on a BLDC motor is composed of permanent magnets so that the poles are fixed while the stator is made of windings so that the magnetic poles can change depending on the polarity of the stator winding currents given. In Figure 3 we will explain the construction of a BLDC motor using 12 stator windings and 8 magnetic poles on the rotor. International Conference on Technology and Vocational Teachers (ICTVT) 2020 Journal of Physics: Conference Series 1833 (2021) 012068 IOP Publishing doi:10.1088/1742-6596/1833/1/012068 In this BLDC motor, the implementation uses a DC source as the main energy source which is then converted into AC voltage using ESC (Electronic Speed Control). To be able to rotate in two directions, the ESC must be modified with the addition of a relay, which can be seen in Figure 4 [6]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002783_09544062211012724-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002783_09544062211012724-Figure4-1.png", + "caption": "Figure 4. Configuration of a 3-RPR parallel mechanism.", + "texts": [ + " Then, the elastic force of each limb can be computed as F01 F02 F03 0 @ 1 A \u00bc H 1 F M (25) Then the deformation in each limb is s0i \u00bc F0i ki ; i \u00bc 1; 2; 3 (26) The stroke of the linear actuator of each limb needed for error compensation is d0i \u00bc s0i rf\u00f0s0i\u00de; i \u00bc 1; 2; 3 (27) where f\u00f0 \u00de is the signum function defined in equation (15) and r is the magnitude of the nominal clearance. Using additional limb stroke \u00bdd01; d02; d03 >, the pose error arisen form clearances and deformation is eliminated. Case study and singularity analysis An example of the 3-RPR PM is used to validate the proposed accuracy analysis method, as shown in Figure 4. A force pair W \u00bc \u00bdF;M > \u00bc \u00bd300N; 100N; 400Nmm > is employed at the center of the manipulator in the global system. The initial coordinates of the connection points at the manipulator(A1;A2;A3) and at the base(B1;B2;B3) are 200cos 3p 2 np 3 200sin 3p 2 np 3 0 BBB@ 1 CCCA and 400cos 4p 3 2np 3 400sin 2p 2np 3 0 BBB@ 1 CCCA respectively, where n\u00bc 1, 2, 3. The global system O \u2013 xy and the body system O1 x1y1 coincide initially. Then the initial pose g0 \u00bc I3 is the 3 3 identity matrix. The rest of the parameters essential to calculation are presented in Table 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000239_iecon.2006.347435-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000239_iecon.2006.347435-Figure6-1.png", + "caption": "Fig. 6. Three marks is attached to a crane", + "texts": [ + " In this section, we explain how to calculate the working radius of a crane. To measure the working radius, we must measure two positions which are the center of rotation and the crane hook. The position of the crane hook is measured by the mark which attached to the crane hook. However, it is difficult to measure the center of rotation directly. Then, first we explain about measurement of the center of rotation of a crane. In order to measure the center of rotation, we attached three marks to the crane body as shown in Fig. 6. The position relationship between the center of rotation and the mark1 which shown in Fig. 6 is well-known. However, we must consider the posture of the crane in order to measure the position of the center of rotation on the basis of the observation device. Then, by using three marks, we measure the posture of the crane. First, the slant of the crane about x axis is measured by mark 1 M1 = (x1, y1, z1) and mark 2 M2 = (x2, y2, z2). The x vector of the crane coordinate X = [lx mx nx]T is calculated by the following equation. dx = \u221a (x2 \u2212 x1)2 + (y2 \u2212 y1)2 + (z2 \u2212 z1)2 lx = x2 \u2212 x1 dx , mx = y2 \u2212 y1 dx , lx = z2 \u2212 z1 dx (15) Next, by using mark 1 and mark 3 M3 = (x3, y3, z3), we measure the slant about y axis Y = [ly my ny]T " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002905_s11277-021-08619-5-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002905_s11277-021-08619-5-Figure2-1.png", + "caption": "Fig. 2 Navigation illustration signified with obstacle presence", + "texts": [ + " The updated robots environment position ( \u2322xk ) is as per the subsequent equations, where W\u0303kQ\u0303kW\u0303 T k and V\u0303kR\u0303kV\u0303 T k denotes the new independent random variables having zero mean and covariance matrices. The EKF is a way of including measurements from multiple sensors to make more robust estimations. This result the accurate positions of obstacle in the region for the mobile robot navigation and the proposed EKF can deal with sensors presenting measurements in different dimensions. After the estimation of obstacle positions then the distance between the obstacles and robot and the robot\u2019s angle for the next moving position is calculated in the next section. Figure\u00a02 illustrates the R\u0303 is mobile Robert the position (sj, tj, uj) and the obstacle distance estimation using Dj (8) angle estimation using \u03b8 (9) of the mobile robot in the coordinate operation. The desired target is represented by the coordinates (sj\u22121, tj\u22121, uj\u22121). (4)\u2322 p \u2212 k = A\u0303k \u2322 p \u2212 k\u22121 A\u0303T k + W\u0303kQ\u0303kW\u0303 T k (5) \u2322 k \u2212 k = \u2322 p \u2212 k H\u0303k(H\u0303k \u2322 p \u2212 k H\u0303T k + V\u0303kR\u0303kV\u0303 T k )\u22121 (6)\u2322 xk = \u2322 x \u2212 k + \u2322 kk( \u2322 zk \u2212 h\u0303( \u2322 x \u2212 k , 0)) (7)\u2322 pk = (1 \u2212 \u2322 kkH\u0303k) \u2322 p \u2212 k 1 3 Distance of obstacle position and angle is needed to calculate for the wheel velocity determination by the proposed AKH-NFIS" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000364_978-3-540-37275-2_15-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000364_978-3-540-37275-2_15-Figure1-1.png", + "caption": "Fig. 1. The laboratory set-up TRMS", + "texts": [ + " The Two Rotor MIMO System (TRMS) is a laboratory set-up deigned for control experiments. In certain aspects, its behavior resembles that of a helicopter. From the control point of view it exemplifies a high order nonlinear system with significant cross-couplings. The approach to control problems connected with the TRMS proposed in this paper involves some theoretical knowledge of laws of physics and some heuristic dependencies difficult to express in analytical form. A schematic diagram of the laboratory set-up is shown in Fig. 1. The TRMS consists of a beam pivoted on its base in such a way that it can rotate freely both in the horizontal and vertical planes. At both ends of the beam, the rotors (the main and tail rotors) are driven by DC motors. A counterbalance arm with a weight at its end is fixed to the beam at the pivot. The state of the beam is described by four process variables: horizontal and vertical angles measured by position sensors fitted at the pivot and two corresponding angular velocities. Two additional states variables are the angular velocity of the rotors measured by tachogenerators coupled with the driving DC motors. In a normal helicopter, the aerodynamic force is controlled by changing the angle of attack. The laboratory set-up from Fig. 1 is so constructed that the angle of attack is fixed. The aero dynamic force is controlled by varying the speed of the rotors. Therefore, the control inputs are supply voltage of DC motors. A change in the voltage value results in a change of the rotation speed of the propeller which results in a change of the corresponding position of the beam. A system performance index is used for fitness function in the RGA. It is an optimization criterion for parameters tuning of control system, which is suitable for the RGA" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003334_icuas51884.2021.9476719-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003334_icuas51884.2021.9476719-Figure1-1.png", + "caption": "Fig. 1 Rigid body model considered for trajectory generation", + "texts": [ + " Simulations are presented in Section IV, where we consider the full six degrees of freedom tracking control of a quadcopter UAV, and show the application of the proposed scheme to trajectory generation followed by trajectory tracking for this UAV. Finally, the research findings of this paper and possible future work are summarized in section V. II. Dynamics Model The rigid body model considered in this paper has four control inputs for the six degrees of freedom. These control inputs are three control inputs that generate a torque for the three degrees of freedom of rotational motion, and one thrust along a body-fixed thrust vector as shown in the figure 1. This model is identical to that used in [16, 19]. This 844 Authorized licensed use limited to: QUAID E AZAM UNIVERSITY. Downloaded on September 02,2021 at 06:19:23 UTC from IEEE Xplore. Restrictions apply. model can be applied to several unmanned vehicles, and the particular case of a quadrotor UAV is considered in section IV for numerical results. Pose is the combination of position and orientation of a rigid body and can be represented as: G = [ R b 0 1 ] \u2208 SE(3), (1) where b \u2208 R3 is the rigid body\u2019s position vector expressed in an inertial coordinate frame and R \u2208 SO(3) is the rigid body\u2019s attitude (orientation) expressed as the rotation matrix from inertial frame to body-fixed frame" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001713_icelmach.2008.4799970-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001713_icelmach.2008.4799970-Figure1-1.png", + "caption": "Fig. 1. Finite element mesh of the four-pole 200 kW induction motor used in the analysis and tests. The mesh consists of 2094 linear triangular elements and 1081 nodes. The FEM motor model calculation is carried out with 25 \u03bcs time steps which is the same as the DTC control cycle.", + "texts": [ + " The measured currents were compared with simulation results obtained using analytical motor model presented with \u0393-equivalent circuit [5] and FEM-motor model. The frequency converter model has analytical models for diode rectifier, the intermediate circuit, and the inverter bridge. The switches of the inverter bridge model are ideal, expect the blanking time. The state of the two switches in an inverter leg cannot be changed simultaneously. All the simulations were carried out with 3 \u03bcs blanking time. The mesh used in the FEM computation is presented in Fig. 1. Because of the symmetry, the finite element mesh of the motor covers only a quarter of the cross-section. r sR RR s Rmj \u03c8\u03c9 s si s Mi \u03c3L MLs su s Ri Fig 2. Dynamic \u0393-equivalent circuit for a squirrel cage induction motor. The FEM computation is carried out with 25 \u03bcs time steps, which is the minimum control pulse length in direct torque control in this case. The analytical motor model is presented in the magnetizing inductance r\u03c3m 2 m M LL LL + = (2) and the total leakage inductance r\u03c3m r\u03c3m s\u03c3\u03c3 LL LL LL + = " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000017_icsmc.1988.754312-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000017_icsmc.1988.754312-Figure1-1.png", + "caption": "Figure 1: The 2-link Manipulator for Simulaaon", + "texts": [ + " can be made linear in terms of a suitably selected vector a of equivalent dynamic parameters 7 = y(q, 4, ;i) a (2) As an example, the dynamic equations of the 2-DOF manipulator in F i g m 1 be explicitly written at F I + (a3c21+a4szl) 6 2-a3szl 4 22 +a4cZ1 42 = (a3c21+qszt);j +a242+a3s214 1z-a4cz14 12=t2 where ~ 2 1 = c o s ( ~ ~ Y ~ ) . spl =sin(qZ-q,). The joint torques are clearly linear in terms of the four parameters ul,a2,u3,u4, which are related to the physical parameters of the links in Figure 1 through aI = Jl + Jn+m2Il2 a3 = m21i211cos8 - moh,rn o2 = Jz + J,, + mora2 a4 = m2li2Itsin8 with the load treated as part of the second link. Note that the number of the equivalent parameters in (2) may be much smaller than that of physical parameters, since the equivalent parameten are their nonlinear combination% This means that the physical paranletem themselves may be unidentifiable. This does not represent a difficulty from a control point of view, since Only h e equivalent parameters affect the dynamics and control", + " The concept of mixed adapthe control, i.e.. combining both the joint motion tracking emor and joint toque predidion m r for parameter adaptation, was recently proposed by the authors as an effective approach to further improve the performancr. of the our globally convergent, direct adaptive robot conholler [Slotiae and Li, 1986, 1987a, b]. The gain-adjusted-forgetting estimator has been used for both purposes with good results. anyway. P-'(t)=-h(t) [ P-' +-I J + WT w 4. SIMULATION RESULTS The manipulator in the simulation, shown in figure 1, is the 2-DOF in our recent adaptwe control expenments. The four parame an? assumed to have the following true values Assume that the arm IS controlled by a joint PD controller to folloy the @.e.) desired uajectory for the joints oI = 0 IS a,4.04 a3=0.03 u4=0 025 I qdl(t) = n(4 + 2(1-cos3t) q&) = n(6 + (I-cos5t) qd has one frequency in each jo trajectories qi(t). q2(t) (very cloy t0 qal and h) and correspondang vel from the c o n t d e d joints are used for parameter estimahon. The init& parameters in the estimators are taken to be all zero and the initial gain matrix to be P(0)=0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000474_bfb0031450-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000474_bfb0031450-Figure5-1.png", + "caption": "Fig. 5 : Singular position", + "texts": [ + " According to equation (6) the two translational velocity components can be found using the velocity transformation Tv(U13,22). When rewriting every factor in function of the coefficients of U13,22, equation (13) becomes : [ /12,1.Ul,a--Ul,l.U2,3 1 | Ul,I.Ul,3 + U2,I.U2,3 [ CWphal, l,3 =\" Oobj . ]( (U2,2_ lll,3) . U13 _ l12,32 31 (17) L[rx3 (Ul,X.//1,3 +/12,1. \"2,3))J The joint velocities can be calculated by inverting equation (14): 41 = ( CJt ,3)-I \u2022 CWphal, l,3 (18) In one singular case, the jacobian cJi,3 can not be inverted. This can be easily understood from figure 5. When the three finger-joints are coUinear, the velocity component parallel to the phalanx and the rotational velocity in frame C1,3 are not independent. Hyper-extension of an inter-phalangeal joint is another interesting phenomenon. In the case of figure 5, all elements of the instantaneous contact point velocity vector are equal to zero, so equation (14) will stand for a homogeneous 3 by 3 set of equations of rank 2. This system can be solved by choosing a parameter value for one unknown, in this case a particular joint velocity. Figure 6a and 6b show the resulting movement after an infinitesimal time step. It shows up that there is always one joint in hyper-extension. One of the two solutions (figure 6a) can be observed in the human hand. ha the solution of figure 6b, the object will jam against the second link of the index" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001187_imece2007-43017-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001187_imece2007-43017-Figure3-1.png", + "caption": "Fig. 3 Coordinate systems for the Gleason Phoenix six-axis machine", + "texts": [ + " The above mathematical model can be applied to simulate both generating and nongenerating face-hobbing processes, including Oerlikon\u2019s Spiroflex and Spirac, and Gleason\u2019s TriAC\u00ae cutting systems. It can also be easily simplified to simulate face-milling cutting, including most existing flank modification features. 3. Mathematical Model of the Cartesian-Type Hypoid Generator The DOF of the proposed Cartesian-type hypoid generator is arranged based on the Gleason Phoenix machine, which has six axes: three rectilinear motions ( , , ) x y z C C C and three rotational motions ( , , ) a b c \u03c8 \u03c8 \u03c8 (see Fig. 3). Such a machine configuration is recognized to have a minimum number of movable axes for the operation of spiral bevel gear cutting. Its coordinate systems ( , , ) t t t t S x y z and 1 1 1 1 ( , , )S x y z , respectively, are rigidly connected to the cutter head and work gear, whose relative positions are described by auxiliary coordinate systems from a S to d S . Here, a \u03c8 and c \u03c8 are the rotation angles of the work gear and cutter, respectively, and b \u03c8 denotes the machine root angle. The horizontal motion x C and the vertical motion y C are used for cutter positioning, while z C is the sliding base for controlling cutting depth" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001691_icnc.2007.678-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001691_icnc.2007.678-Figure2-1.png", + "caption": "Figure 2. Schematic diagram of the motor", + "texts": [ + " The fitness is evaluated by the reciprocal of the error of mean square, namely \u2211 +\u2212= \u2212== k pki d k k iyiyp E f 1 2))()((1 (18) where kf is the fitness value at time k , p is the width of the identification window, )(iyd and )(iy are the expected and actual outputs at time i , respectively. Besides, the fitness function is linearly adjusted to avoid early convergence and ensure the variety of individuals. Numerical simulations are performed using the proposed method for the speed identification of a longitudinal oscillation USM [6] shown in Figure 2. Some parameters on the USM model are taken as: driving frequency 27.8kHZ, amplitude of rated driving voltage 300V, rated output moment 2.5kg\u00b7cm, rotation speed 3.8m/s. Besides, the initial ranges of all the weights are in [-5, 5], the initial value of the self-feedback coefficient \u03b1 is taken as 0.4, and the initial number of the neurons of the hidden layer is 10. The Block diagram of identification model of the motor is shown in Figure 3. In the following section, we present and discuss an identifier to perform the identification of non-linear systems based on the AGA-based Elman Neural Network proposed in the previous section" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001578_iros.2008.4650591-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001578_iros.2008.4650591-Figure1-1.png", + "caption": "Fig. 1. Schematic figure of an insect leg (right front leg). It consists of three hinge joints. \u03b1-joint moves the leg for- and backward (protraction means positive direction). \u03b2- and \u03b3-joint operate in the leg plane, meaning their axes are parallel to each other and are perpendicular to the drawn leg plane. Lifting the leg equals a positive movement in the \u03b2-joint, an outward going movement produced by the \u03b3-joint (extension) means a positive movement. The origins of the joint coordinate systems are set corresponding to leg positions in a standing walker (\u03b1 is in a middle position, while \u03b2- and \u03b3-joints are in a position in which the femur is approximately parallel to the ground and the tibia is nearly orthogonal to the femur.", + "texts": [ + " The controller structure has been tested in a dynamic simulation environment and some of the simulation results are presented to stress the feasibility of the LPVF approach for six-legged walking. The simulations has been performed in BREVE (version 2.6 [14]). BREVE is an integrated environment using OpenGL for visualization which incorporates the ODE library for simulating the dynamics. The walker is a scaled-up version of a stick insect (ten times) and matches the robot TARRY IIB on which the LPVF approach shall be implemented next. The control of walking of a hexapod walker, like the stick insect, involves the control of six legs with three degrees of freedom each (see Fig. 1). The Walknet has been introduced to control such an insect-like artificial structure consisting of six legs with three hinge joints. The configuration of this robot mimics the stick insect with respect to its dimensions, position and orientation of the joints. This results in a robot with a total of 18 degrees of freedom which have to be controlled in parallel. It seems natural to divide this control problem up into smaller parts. Biological findings [7], [11], [4] suggest that each leg has its own controller structure which only has to deal with three joints (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002646_s11517-021-02347-5-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002646_s11517-021-02347-5-Figure2-1.png", + "caption": "Fig. 2 Prostate intervention robot: (a) CAD model, (b) photography", + "texts": [ + " The optical encoder, based on the principle of grating diffraction, is widely used in the robotic systems operating in the MR environment [5, 8]. The HKT30-C03 two channel kit incremental encoder (RepAvago Wuxi, China), with an encoding resolution of 0.087\u00b0/ count for rotary motion, is selected as the optical encoder in the robot. The material of the optical grating is mylar and the signals acquired by the optical encoders are transmitted through shielded, twisted pair cables to the encoder interfaces, which are placed inside an electromagnetic interference (EMI) shielded enclosed controller box. Fig. 2 depicts the 6-DOF serial prostate intervention robot with overall dimensions of 220 mm \u00d7 200 mm \u00d7 170 mm. Each joint is actuated by an independent PMR30 with a rotary optical encoder to feedback the joint variable. The proposed robot contains a 3-DOF cartesian module, a 2-DOF orientation module, and a 1-DOF needle insertion module. The details are described as follows. The cartesian module is secured on the MRI scanner bed and guarantees the accuracy and stability of the robot. The cartesian module employs 3 linear guide systems to formulate 3 orthogonal translational DOFs in X-Y-Z directions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001781_20070625-5-fr-2916.00007-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001781_20070625-5-fr-2916.00007-Figure1-1.png", + "caption": "Figure 1. EFIGENIA S/VTOL-UAV Airplane.", + "texts": [ + " The EFIGENIA was designed and developed to validate and demonstrate the flying qualities and performance characteristics of a short or vertical takeoff and landing (S/VTOL) unusual experimental unmanned aerial vehicle that gives to the aerial Vehicle the vertical flight capability and low speed 1Electronics Engineer, Aeronautical Engineering researcher in Unmanned Aerial Vehicles UAV. cordoba@efigenia-aerospace.com flight characteristics of a helicopter1 and the horizontal cruise speed of a conventional aircraft2, 3 (Figure 1). 3. EFIGENIA UAV Aircraft Design The EFIGENIA UAV is built of robust, lightweight, and high-strength materials. EFIGENIA aerospace design introduces an S/VTOL Rotor and Tailless Forward Swept Wing Concept with the purpose of allowing to the air vehicle an excellent aeromechanical behavior4, 5. The EFIGENIA is powered by two 2.0 HP engines located each one in the nose and tail fuselage respectively, and one more 2,25 HP engine inside the aerial vehicle body. In contrast, the tail engine has been adapted for conform a thrust vectoring unit to aimed high performance flight control system, maneuverability and agility at low speeds" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001650_pes.2007.386081-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001650_pes.2007.386081-Figure2-1.png", + "caption": "Fig. 2. Voltage phasor diagram for ground fault with resistance at point Y", + "texts": [ + " SE& and RE& are source voltages, SZ and RZ are source impedance respectively, F is fault point and Y is the point at the end of the reach setting. Compensated voltage is defined by the following formula, YZIUU *&&& \u2212=\u2032 , where YZ is the setting impedance, for phase-to-phase fault \u03d5\u03d5UU && = , \u03d5\u03d5II && = , for single-phase-to- ground fault \u03d5UU && = , 0* IkII Z &&& += \u03d5 . If fault occurs at point Y, the compensated voltage trajectory varying with fault resistance in voltage plane is the real arc of circle C1 as shown in Fig.2, which takes |0|FU& as chord. FU& is the fault point voltage, the subscript |0| all means the pre-fault phasor. And |0|FU& is identical with the prefault compensated voltage |0|U \u2032& . Along with the increase of fault resistance, FU& will reach Fault Component Impedance Relay with Double Circular Characteristic Shaofeng Huang, Qiankuan Liu, Huanzhang Liu, and Qixun Yang T 1-4244-1298-6/07/$25.00 \u00a92007 IEEE. point I at which C1 crosses RE& , then FU& and RE& are inphase with each other, and I is called the \u201cin-phase\u201d point", + " Changing the center and radius of C1 on the basis of (1), this paper presents a general operating equation of the compensated voltage. |||| |0||0| UkUkU RC &&& \u2032>\u2032\u2212\u2032 (2) Where Ck is defined as \u201ccircle centre coefficient\u201d and 5.0\u2265Ck , Rk is defined as \u201cradius coefficient\u201d. In voltage plane, the boundary of (2) is a circle with the end of phasor |0|UkC & \u2032 as centre, |0|UkR & \u2032 as radius. Formula (2) is called the general operating equation in this paper. Circles with different operating areas can be obtained by selected different Ck and Rk . For example, 5.0== RC kk , (2) becomes to (1) corresponding to C1 in Fig.2. 1=Ck , YIkR > , circle C2 can be drawn as shown in Fig.2, which takes the end of phasor |0|U \u2032& as centre. It can be seen that C2 includes the \u201cin-phase\u201d area, and its outside is the operating zone of the compensated voltage, so C2 can prevent the relay from the \u201cin-phase\u201d problem and can be employed as the additional blocking condition. If 1== RC kk , it follows that |||| |0|UU && \u2032>\u2032\u2206 (3) The boundary of (3) corresponds to C0 as shown in Fig.2. It can be seen from Fig.2, (3) will be failure to operate if the fault compensated voltage falls into the area between C0 and C1 along with the increase of fault resistance, this means that (3) sacrifices some abilities against fault resistance. However, C0 completely excludes the in-phase area, so the inphase problem can hardly influence on (3), whose security has been improved to a great extent. How to do everything possible to enhance the resistive tolerance while the security of (2) can be ensured has been studied and discussed in this paper", + " This relay is composed of two group criteria comprising four operating equations as follows. Criteria 1: |||5.0| |0||0| UkUU op &&& \u2032>\u2032\u2212\u2032 (4) |||| |0||0| UkUU in &&& \u2032>\u2032\u2212\u2032 (5) Criteria 2: |||| |0||0| UkUkU opop &&& \u2032>\u2032\u2212\u2032 (6) \u03b1\u03b1 \u2212> \u2032 \u2032 >\u2212 0 |0| 0 180arg360 U U & & (7) Where opk and ink are defined as \u201ccompensated coefficient\u201d and \u201cin-phase coefficient\u201d respectively to distinguish them from coefficients of the general operating equation (2), and 15.0 << opk , 1<< inkYI . \u03b1 is defined as \u201cin-phase angle\u201d between lines L and OY as shown in Fig.2, whose value has relation with the phase-angle difference, \u03b4, of two source-voltage phasors. Criteria 1 is composed of (4) and (5) while criteria 2 consists of (6) and (7), they are independent for each other. Both group criteria jointly make up an integrated relay called the fault component distance relay based on the pre-fault compensated voltage in this paper. The logic of two group criteria is shown in Fig.3. Referring to Fig.2 and corresponding to it, the operating zones of (1) and (3)-(7) in voltage plane can be drawn as shown in Fig.4. Circle C3 and dashed circle C\u20323 respectively correspond to (4) and (6). Since 5.0>opk , C3 includes the origin O while C\u20323 passes through point O. Taking the end of |0|U \u2032& as centre, |0|Ukin \u2032& as radius, C2 corresponds to (5). The outside of each circle is the corresponding operation area of the compensated voltage. L is the operating boundary of (7), whose operation area is the shadow part. Compared C3 with C1 in Fig.4, they are concentric circles, but the radius of C3 is larger than that of C1, accordingly the operating area of C3 is lessened. That is to say, the abilities against fault resistance of (4) and (6) are less than that of (1), the sensitivities of (4) and (6) are dropped while their security is enhanced. Referring to Fig.2, C2 and L can exclude the in-phase region. The outside of circle and the shadow area of L are the relay permissive operating zone, so (5) and (7) try to eliminate the influence of \u201cin-phase\u201d problem to avoid the relay unwanted operation in in-phase region. Therefore, (5) and (7) are also called the blocking criterion or condition of in-phase area. In a word, on the premise of the blocking condition of inphase area and sacrificing some resistive tolerance, the fault component distance relay based on pre-fault compensated voltage can either ensure security or improve the capability against fault resistance", + " The voltage phasor diagram for the receiving end under external fault in forward direction is identical with that for the sending end under reverse fault. The same analysis shows that the probability of undesired operation for the fault component distance relay increases along with the fault point far away from the location at the end of the reach setting. The most disadvantageous condition is that the fault point is located at the sending source, so the in-phase coefficient ink is calculated according to this case. In Fig.2, if the radius of C2 is larger than the length of YI, C2 can include all the \u201cin-phase\u201d area, so YI is the minimum radius of C2. In right triangle OIY, the minimum radius 2CR can be obtained by \u03b1\u03b1 sin*||sin* |0|2 FC UOYR \u2032== & (8) Where \u03b1 is the angle difference between |0|FU \u2032& and RE& . In the most serious condition that fault location lies at the sending source as shown in Fig.6, \u03b1 approximates the rotor angle \u03b4, then In a normal load condition, the magnitude of |0|FU \u2032& approximates that of |0|U \u2032& , so ink can be obtained by \u03b4sin\u2265ink (10) It is very difficult in a complex power system to get the proper angle \u03b1 or \u03b4, but one approximation obtained through use of swing-center voltage \u03d5cosU is as follows" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001110_detc2007-34199-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001110_detc2007-34199-Figure1-1.png", + "caption": "Figure 1: Mechanical model of rotor system with", + "texts": [ + " [7] is improved for approaching practice well. The nonlinear oil-film force model put forward by Adiletta [8] is applied and Runge-Kutta numerical method is adopted to carry out numerical simulation. At last, experiments are performed to test the simulation results by a rotor model test rig. The results show that the numerical results agree well with the experimental results. Simplified rotor system model with one end bolt looseness is considered in this paper. The sketch map of this model is shown in Figure 1. Two uniform slide bearing support the double end of rotor. Rotor lumped mass in two bearing is 1m , the equivalent lumped masses of disc and bearing pedestal are 2m and 3m , respectively. The stiffness of elastic shaft is k , damped coefficients of bearing, disc and bearing pedestal are 1c , 2c and fc , respectively, supported stiffness is fk . Assumed that Elastic shaft between disc and bearing is massless. pedestal looseness Nonlinear oil-film force model put forward by Adiletta is adopted, which simplifies the computation of oil-film force by short bearing theory" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001350_imece2008-67917-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001350_imece2008-67917-Figure2-1.png", + "caption": "Figure 2: Test fixture", + "texts": [ + " Achieving these objectives can help engineers better understand the behavior of electronic components within projectiles as well as help improve FEA of these systems. U.S. Army ARDEC Tank-automotive & Armaments Command-Armaments Research, Development & Engineering Center (TACOM-ARDEC) provided the projectile under consideration. The projectile consists of several components as shown in Figure 1. All these components are threaded or bolted together. An experimental test fixture with the same mass and mass moment of inertia as the original projectile is used in this research, Figure 2 and Figure 3. The fixture has two components: impactor and housing. Impactor has a tapered end where an impact hammer is used to send load to an electronic board. The electronic board is situated between the impactor and the housing. A 4.5\u201d-15 thread is used to connect the impactor and the housing. The outer surfaces of the impactor and housing have flat surfaces to allow tightening them together. A cylindrical board holder is used to maintain the board in place. Acceleration is measured at the center of the board using an accelerometer as shown in Figure 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001007_s1068798x08120113-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001007_s1068798x08120113-Figure5-1.png", + "caption": "Fig. 5. Phase portrait of the nonlinear dynamics of an elastic-smoothing system: (1, 3) boundaries corresponding to change in direction of the indenter\u2019s vibrational motion along the Y axis; (2, 4) boundaries corresponding to the beginning of indenter insertion in and extraction from the part; (5) self-oscillation trajectory of the indenter; (6) boundary corresponding to transition to damping indenter oscillations; (7) section where the indenter is in a stable state; (8, 9) directions of the phase trajectories to the self-oscillatory cycle; (10) direction of the phase trajectories to a stable state of the indenter.", + "texts": [ + " The discontinuous nonlinearities correspond to the switching curves at which the mapping points pass to the next sheet of the phase portrait. Each phase trajectory corresponds to the motion of the smoothing tool for specified initial values W st C0 yi hmax+( )dL 0 L \u222b\u2248 \u2013 C0 yi k2yi+( )dL 0 L \u222b C0S2 C0S1\u2013= 1 2 --C0hmaxL 1 2 --C0k2yL RUSSIAN ENGINEERING RESEARCH Vol. 28 No. 12 2008 NONLINEAR DYNAMICS OF THE ELASTIC SMOOTHING OF SURFACES 1205 of the coordinate y and the indenter\u2019s vibrational speed dy/dt (Fig. 5). Sheet 1 of the phase plane describes tool introduction in the surface and is bounded by two switching curves. Curve 2 corresponds to the moment at which the tool touches the surface and hence to the moment at which the solution switches from Eq. (7) to Eq. (4). The initial conditions for Eq. (7) are the final values of the variables calculated by computer simulation for Eq. (4). Switching curve 3 characterizes the reversal in sign of the indenter\u2019s vibrational velocity, which corresponds to switching of the solution from Eq", + " The transient process in the nonlinear dynamic system ends in a stable state (singular curve 7) or in self-oscillation (singular curve 5). The stable-state coordinates of the working part of the indenter correspond to the solution of Eqs. (1) and (2) when the derivatives are zero. The type of singular point is determined by the roots of the characteristic equation. The complex roots in the region bounded by singular curve 6 (an unstable limiting cycle) correspond to phase trajectories converging along a spiral curve to the stable-focus singular point 7 (Fig. 5). In the region bounded by unstable limiting cycle 6, elastic smoothing is stable. The stable focus 7 extends along the Y axis to the singular curve; this corresponds to a region of possible stable states of the dynamic system in the given conditions. Correspondingly, there is indeterminacy in the stable-position coordinate of the indenter\u2019s working surface within the singular curve 7 in the case of smoothing where there is dry friction in the indenter\u2019s guide pieces. Hence, singular curve 7 on the phase portrait (Fig. 5) corresponds to a stagnant indenter position; this may result in the formation of a nonperiodic wave structure at the smoothed surface. If the amplitude of indenter oscillation increases to values bounded by curve 6 under the action of random factors, the direction 10 of the phase trajectories is reversed to direction 9. Correspondingly, the stable state of the indenter gives way to self-oscillation, whose amplitude increase to values bounded by the stable limiting cycle in curve 5. From the phase portrait, we determine the boundaries of regions with different transient processes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002641_j.robot.2021.103783-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002641_j.robot.2021.103783-Figure2-1.png", + "caption": "Fig. 2. Schematic diagram of the redundantly actuated planar PM 2RPR + P.", + "texts": [ + " In the previous works focusing on the displacement coordination of the redundantly actuated PMs, the elastic deformation is not considered, but under the action of external load, the elastic deformation for the real robots or equipments based on such PMs always exist, thus the force distribution analyses of previous works are not accurate enough. In this section, the 2RPR + P planar PM will be taken as an example to illustrate the establishment of the displacement coordination equations of the redundantly actuated PMs. P a t o o i a a f u t i e Fig. 2 shows that the 2RPR + P redundantly actuated planar M consists of a moving platform, a base, two RPR actuated limbs nd a constrained limb composed of a prismatic joint, in which he axial direction of the two RPR actuated limbs and the axis f the P joint within the constrained limb are parallel to each ther. For the convenience of analysis, a coordinate system o-xyz s established at the center of the moving platform, with the yxis being parallel to the axis of the P joint in the constrained limb nd the z-axis pointing along the normal direction of the plane ormed by the mechanism" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000239_iecon.2006.347435-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000239_iecon.2006.347435-Figure1-1.png", + "caption": "Fig. 1. (a)The observation device (b)The cross mark", + "texts": [ + " And some researches are including position measurement based on image processing, and the measurement method by stereo camera[4] have been accomplished. However, the application of those researches to the industry is difficult. And in our research, we suppose that the crane is huge and the working radius change from several to 100m with expansion and contraction of the jib. It is difficult to measure the huge object for traditional researches. The purpose of this paper is measurement of the working radius of the crane. In order to measure the working radius of the crane, we develop an observation device (shown in Fig. 1(a)) with a CCD camera and a laser range finder. And we propose the methods of recognition of the crane hook. However, a shape of the crane hook is not unique, so it is difficult to recognize it. Then we attempt to simplify A part of this work was supported by TADANO LTD and Japan Society for the Promotion of Science. recognition by attaching a specific mark to the crane hook, and we recognize the mark and measure the position. Next, the crane hook is moved by an operator and own weight. Then to track the moving hook, we control the posture of the observation device", + " This paper is organized as follows, in Section II, we explain methods the recognition of the mark and position measurement using image information. And Section III expresses the tracking control based on image information. Then, we calculate the working radius of the crane in Section IV. In Section V we show the experimental results and verify the availability of the observation device. Finally, Section VI makes the summary. This section explains the methods of the mark recognition and position measurement. First, we introduce the mark which we attach to the crane. In this paper, we use the mark (shown in Fig. 1) which has the cross in the circle with a diameter of 30 cm. Measurement process consists of the mark recognition by the image processing and the position measurement of the mark based on the camera image, the laser range finder and the encoder information of motors. Additionally we control the zoom lens in order to keep the mark size on the image constant. Details are described below. We explain about the mark recognition based on the image processing. Today, several methods are proposed by 1841-4244-0136-4/06/$20", + " And tf means the threshold of brightness, and this value is given by histogram of input image. This paper defines that brightness value 0 is black and brightness value 255 is white. Next median filter smoothes the binary image. Then labels are attached to connecting regions which have the white pixel based on a Scan Line Seed Fill Algorithm. In the labeling process, the area S and the circumference length L of an each connecting region are taken. However the area S which is taken from connecting region does not contain the area of cross line shown in Fig. 1, so we use the area S\u0302 in consideration of the cross line area and 1.16 times the area size S in order to correct the lack of the circle area. Using parameters of the connecting region, the degree of circularity of the connecting region e(= 4\u03c0S\u0302/L2) is calculated. The degree of circularity takes the value of 0 \u2264 e \u2264 1, and a perfect circle has 1. Then we set the suitable value of the threshold and detect the connection region with the bigger degree of circularity than threshold. This paper set threshold 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002430_012006-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002430_012006-Figure5-1.png", + "caption": "Figure 5. Dynamic model of parallel mechanism", + "texts": [ + " 4) Slider\u2019s acceleration Differentiating both sides of equation (7) with respect to time leads to: 2 2 1 1 1 1 1 1 1 1 1 3 1 3 \u02c6 \u02c6( ) ( )q l l e Eu u ER R R R (11) Multiplying the two sides of equation (11) with 1u to get: 2 T 2 1 1 1 1 1 1 3T 1 1 1 = ( )q l J u R R u e (12) 5) Link\u2019s angular acceleration Multiplying the two sides of equation (11) with 1E\u0302u to get: T T T 1 1 1 1 3 1 1 3 T 1 1 2 1 1 21 1 1 1 1 \u02c6 \u02c6(( ) ) ( )\u02c6 \u02c6( ) ( )1 1 1 = ( ) ( ) 0 l l l Eu e u ER R Eu ER R Ju e D D D D (13) Where T 2 T 21 1 1 1 1 1 1 3T 1 1 \u02c6( ) = ( )l D Eu e u R R u e ; T 2 2 1 1 3 \u02c6( ) D Eu R R . 6) Centroid acceleration analysis of link Differentiating both sides of equation (10) with respect to time leads to: 2 21 1 1 1 1 1 2 1 3 1 \u02c6 = ( ) 2 2v l Eu v J R R uD D (14) Without considering the friction, the dynamic model of the mechanism is obtained as shown in Figure 5. In the figure, dm is the mass of the moving platform; zm is the mass of the adapter; en is the external moment of the moving platform; ef is the external force of the moving platform; il m is the mass of the MEMAT 2021 Journal of Physics: Conference Series 1820 (2021) 012006 IOP Publishing doi:10.1088/1742-6596/1820/1/012006 link i ; iqm is the mass of the slider i ; In view of the simple structure of the parallel mechanism, and the Newton-Euler method is intuitive and easy to understand, this article adopts the Newton-Euler method to establish the dynamic model of the parallel rotating mechanism" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001775_apccas.2008.4746130-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001775_apccas.2008.4746130-Figure2-1.png", + "caption": "Fig. 2 The Knee Surgical Robot Using a Hybrid Cartesian Parallel Manipulator", + "texts": [ + " No whole pieces of cut bone are remained. The proposed novel bone resection method is shown in Fig. 1(b). The side edge of the cutter, an 741978-1-4244-2342-2/08/$25.00 \u00a92008 IEEE. end mill, is used to cut the bone by moving the cutter along the cutting planes of the designate shape of bone preparation. Lateral milling of bone was actuated by a hybrid Cartesian parallel manipulator, which is modified from [5] to satisfy sufficient degree of freedoms for desired position and orientation, as shown in Fig. 2. The modified Cartesian parallel mechanism has three limbs with three passive prismatic joints instead of revolute ones. The problems with revolute joints of the CPM in [5] were then avoided of its singularities and limb configuration out of the fixed frame. Furthermore, the robot is kinematically decoupled for alignment purposes and cutting-bone function. As a result, only one motor is needed to be actuated at once time so that the robot mechanism has the features of control simplicity, high positioning accuracy and safety" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002755_052027-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002755_052027-Figure2-1.png", + "caption": "Fig. 2 Planetary gear system", + "texts": [ + " For the test of gear bending fatigue, the Normal distribution, Lognormal distribution and two-parameter Weibull distribution are usually used for the test of the life distribution[19]. The weighted least square method[20] is used for the distribution fitting of mR. The statistical parameters of the distribution of mR are listed in bility model of gear system considering strength degradation can be expressed as \ud835\udc45 \ud835\udc53 \ud835\udc5a \u210e \ud835\udc47 \u220f \ud835\udc54 \ud835\udeff \ud835\udeff \ud835\udc51\ud835\udeff \ud835\udeff \ud835\udc51\ud835\udc47 \ud835\udc51\ud835\udc5a (14) Taking a planetary gear system for the case study, it is shown in Fig. 2, which contains three planetary gears. The parameters of its gears are listed in Table 3. Profile shift coefficient -0.33 0.38 -0.43 0.52 0.44 0.36 -0.23 In order to validate the proposed reliability model (Eq. (14)), Monte Carlo simulation is used for the comparison with the proposed method. Because Monte Carlo simulation is inefficient, the gear system composed of the sun gear and the planetary gears is used to validate the proposed model. The procedure of the Monte Carlo simulation is shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000312_978-3-540-44410-7_8-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000312_978-3-540-44410-7_8-Figure1-1.png", + "caption": "Fig. 1. A common example of a redundant fixed-base single arm.", + "texts": [ + " Then, in Section 4, the previously obtained results are extended to the more general case of still 3D and nonholonomic multiarm mobile systems, when performing grasping and object manipulation tasks. Some conclusions and directions for future research activities are given in a final section. 2 Control of a Fixed-Base Single Arm with Singularity Avoidance Let us consider a redundant fixed-base single arm, i.e. with a number of degrees of freedom (dof\u2019s) greater than six. Without loss of generality, let us refer to the arm in Fig. 1, where the wrist be constituted by a 3-dof rotational joint, typically of Euler and/or Roll-Pitch-Yaw type. In the figure, frame < g > represents the \u201cgoal frame\u201d, which has to be reached (in position and orientation) by the \u201cend-effector frame\u201d < e > of the manipulator. Let e := [ \u03c1T dT ]T (1) be the collection of the misalignment error vector \u03c1 and distance error vector d of frame < e > with respect to , when projected on world frame < 0 >. Also, let x\u0307 := [ \u2126T \u03bdT ]T (2) be the collection of angular and linear velocities of < e >, still projected on < 0 >, where a small abuse of notation has occurred as for the use of the derivative" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000879_aupec.2007.4548084-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000879_aupec.2007.4548084-Figure3-1.png", + "caption": "Fig. 3: Block diagram of the implementation of the esd augmented NFO control on the experimental system.", + "texts": [ + " An improved version of NFO control was suggested in [4] that helps overcome the above problem in a major part of regeneration region while retaining the algorithm\u2019s inherent simplicity. It was noticed that a non-zero esd voltage in steady state would indicate misalignment of the control frame. The presence of this signal was used to provide an auxiliary feedback that can stabilise the control frame alignment by applying the following augmentation to the frame angular velocity estimation: \u03c9\u0302ms = esq \u2212 sign(esq)kesd |\u03c8\u2217 s | (2) where k > 0 is included as a tunable gain parameter. Fig. 3 illustrates implementation of the augmented NFO control algorithm on a laboratory system used in our experiments. Note that with k = 0 one has the implementation of the nonaugmented (\u201craw\u201d) NFO version. Some residual misalignment of the assumed and true positions of the stator flux vector under NFO control will always be present due to non-linear and leakage effects, as explained in [5]. If such a misalignment is small it does not compromise stable operation of the algorithm under both motoring and regeneration conditions", + " 8a that under the simulated 4\u00b5sec dead time conditions, the \u221250% error in the Rs value results in a very minor difference to the NFO algorithm performance. This confirms the algorithm robustness under such extreme conditions as those being simulated. The experiments were carried out using an IGBT based inverter connected to a 7.5kW induction machine. The machine was mounted on a dynamometer test bed with a DC load machine configured as a simple Ward-Leonard system to provide static loading capable of regeneration or motoring operation. The structure of the control system used for experiments corresponds to the block diagram of Fig. 3. Two sets of plots shown in Fig. 9 were obtained for the exact and \u221250% erroneous stator resistance values. The dead time error of the inverter was approximately 4\u00b5secs but due to the earlier explained specific implementation was not estimated during the run-time of the machine. After fluxing the machine for 50 msecs (not shown in the plots) a step up from zero to 10 rad/sec in the speed reference value was applied followed by a step down from 10 to 5 rad/secs. Thus the conditions of the experiment were very close to those in simulations of Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000879_aupec.2007.4548084-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000879_aupec.2007.4548084-Figure4-1.png", + "caption": "Fig. 4: NFO reaction to the stator resistance errors.", + "texts": [ + " Simulation and experimental results are presented to support the main ideas of the paper. These ideas are then summarised in the Conclusions. The following factors have been found to limit the slow speed performance of the NFO control algorithm: \u2022 Sensitivity to the stator resistance value; \u2022 Side effect of the auxiliary esd feedback for the aug- mented NFO; \u2022 Sensitivity to the inverter dead time error. In the following subsections these are explained in more detail. This effect can be explained with the help of Fig. 4a and Fig. 4b. First assume the estimated stator resistance value smaller than the true value, i.e. R\u0302s = Rs \u2212 \u2206Rs. This is equivalent to adding a vector with relative coordinates (\u2206Rsi\u2217sd,\u2206Rsi \u2217 sq) to the assumed flux voltage vector. To let esd become positive the control frame should turn counterclockwise with respect to the true frame as shown in Fig. 4a. In the new misaligned position e\u0302sd = esy sin \u03b8e+\u2206Rsi\u2217sd e\u0302sq = esy cos \u03b8e+\u2206Rsi\u2217sq (3) and the estimated angular velocity of the stator flux vector is: \u03c9\u0302ms = e\u0302sq |\u03c8\u2217 s | = esy cos \u03b8e + \u2206Rsi\u2217sq Lmi\u2217sd (4) The true angular velocity of the stator flux vector would change because of the changed flux: \u03c9ms = esy |\u03c8 s | = esy Lm ( i\u2217sd cos \u03b8e \u2212 i\u2217sq sin \u03b8e ) (5) If the two angular velocities are equal, the misalignment remains statically stable. This corresponds to: esy sin \u03b8e \u2206Rsi\u2217sq = i\u2217sd cos \u03b8e \u2212 i\u2217sq sin \u03b8e i\u2217sq cos \u03b8e + i\u2217sd sin \u03b8e = isx isy (6) Figure 4a shows the position of the frames corresponding to the above condition. Applying similar logic to the case when R\u0302s = Rs + \u2206Rs leads to the static misalignment of Fig. 4b. It can be easily shown that for a wide range of \u2206Rs values of both signs and a wide range of practical ratios for x = i\u2217sq/i \u2217 sd, an angle value that satisfies (6) exists. Furthermore, in most cases this angle has a low sensitivity with respect to \u2206Rs, which can be approximately found by solving the equation: Lm\u03c9ms\u03b8 2 e + (Lm\u03c9ms + \u2206Rsx)x\u03b8e \u2212\u2206Rsx = 0 (7) Equation (7) is derived from (6) by assuming cos \u03b8e \u2248 1, sin \u03b8e \u2248 \u03b8e and esy/i\u2217sd \u2248 Lm\u03c9ms. It follows from (7) that the frame misalignment resulting from the stator resistance error is dependent not only on torque (through the x parameter) but also on frame angular velocity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001496_s10773-007-9497-9-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001496_s10773-007-9497-9-Figure1-1.png", + "caption": "Fig. 1 A vertical disk rolling on a horizontal plane", + "texts": [ + " (28) A special case of (24) is when the constraints are time-independent (scleronomic) and the coordinates qs+1, . . . , qn are cyclic, i.e. Bs+\u03b2,\u03b1 = Bs+\u03b2,\u03b1(q1, . . . , qs), Bs+\u03b2,t = 0. (29) Equation (28) degenerates then to the Chaplygin equation [5, 14]: d dt \u2202T\u0303 \u2202q\u0307\u03b1 \u2212 \u2202T\u0303 \u2202q\u03b1 \u2212 k\u2211 \u03b2=1 \u2202T \u2202q\u0307s+\u03b2 ( \u2202Bs+\u03b2,\u03b1 \u2202q\u03c3 \u2212 \u2202Bs+\u03b2,\u03c3 \u2202q\u03b1 ) q\u0307\u03c3 = Q\u0303\u03b1 (\u03b1 = 1, . . . , s). (30) For a demonstration we consider deliberately a simple example, namely that of a vertical uniform disk rolling on a horizontal xy-plane without slipping (cf. Fig. 1). The radius and the mass of the disk are respectively a and m. We may choose the generalized coordinates to be the Cartesian coordinates x, y of the disk center, the angle of rotation \u03c6 with respect the disk axis, and an angle \u03b8 between the axis of disk and the x-axis. The constraints in the directions tangential and normal to the trace of the contact point read respectively We may choose q\u03071 = \u03c6\u0307, q\u03072 = \u03b8\u0307 as the independent velocities, and q\u03073 = x\u0307, q\u03074 = y\u0307 as the dependent velocities. From (31) and (32), we can solve q\u03073 = x\u0307 = a\u03c6\u0307 sin \u03b8, (33) q\u03074 = y\u0307 = \u2212a\u03c6\u0307 cos \u03b8, (34) which implies that (cf" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000582_2008-01-2631-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000582_2008-01-2631-Figure4-1.png", + "caption": "Fig. 4 Mechanical diagnostic test rig Fig. 5 Location of the accelerometer on", + "texts": [ + " The objective function or performance index which has to be minimized is set as [8]: ni i XXXXXXxF 1 321 )()()()( (21) eeeJ 32 0 1 (22) Where; XX , and x are the vibration acceleration, vibration velocity, and vibration displacement respectively, XX , and X the mean value of the vibration acceleration, vibration velocity, and vibration displacement respectively, and the 1, 2 and, 3 are weighting factors. The global block diagram shown in Fig. 3 shows the sequence of the basic operators used in the optimization. The program will terminate when the accuracy reaches to the setting value or when the maximum generation is reached. The experiment work was performed on a mechanical diagnostic test rig, as illustrated in Fig. 4, with the location of the accelerometer on the gearbox input shaft bearing is shown in Fig. 5. The primary purpose of this rig is to study the effects of gear tooth damage. The use of this rig for diagnostic studies is practical. The rig consists of a 25 horsepower (Hp), 3000 rev/min induction motor drawing power and driving a 4-stage automotive gearbox, a separate hydraulic brake that is couple to the output shaft of the gearbox. A flywheel of weight of 250 N is also connected to increase the amount of load and help for the system stability" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002760_tmag.2021.3081921-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002760_tmag.2021.3081921-Figure2-1.png", + "caption": "Fig. 2: Rotor winding configuration: (a) cage rotor (parallelconnected), (b) wound rotor (series-connected).", + "texts": [ + " \ud835\udc35 is the parameter for the UMP primarily caused by magnetizing current. Both \ud835\udc34 and \ud835\udc35 constants can be found by experimental work or Finite Element Analysis (FEA). After plotting UMP-slip curve, a curve-fitting method is used to find the constants, where \ud835\udc34 is the quadratic component and \ud835\udc35 is the constant component. Based on the rotor winding configuration, induction machines can be divided into squirrel cage induction machine (SCIM) and wound rotor induction machine (WRIM). Typically, the SCIM has a parallel-connected rotor bar (see Fig. 2 (a)), in which the current in every rotor bar can be different, while the WRIM has a series-connected rotor winding (see Fig. 2 (b)), in which the same current flows in the rotor winding of the same pole. Therefore, SCIM generally has a lower UMP than the WRIM, due to the parallel paths at the cage rotor produce the counteracting flux to damp the UMP. This scenario was extensively discussed in [24]. To consider the UMP damping effect in the UMP calculation, a UMP damping coefficient is proposed in [24] for steady state UMP calculation. UMP damping coefficient is the ratio between the resultant magnetic flux of \ud835\udc5d \u00b1 1 (\ud835\udecc\ud835\udc5b (\ud835\udc5d\u00b11)) and the \ud835\udc5d \u00b1 1 magnetic flux produced by the stator (\ud835\udecc\ud835\udc60 (\ud835\udc5d\u00b11))" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001573_s09-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001573_s09-Figure2-1.png", + "caption": "Figure 2. Analysis of experimental data for the total gas pressure over liquid UO2 at different temperatures: 1\u00d0computational fitting of the curve of driven uniform evaporation to the experimental data ( ) by varying the input thermodynamic data within the acceptable error interval, 2\u00d0the curve of driven uniform evaporation normalized to the ratio O=U 2:00, 3\u00d0the curve of driven uniform evaporation plotted using the equation of state of irradiated fuel UO2 (the equation of vapor\u00b1 liquid equilibrium).", + "texts": [ + " In the case of thermodynamic equilibrium, the law of mass action yields [2] DGT MOB \u00ff DGT MOA RT ln \u00ff pMOB aMOA p B\u00ffA =2 O2 0 or, after the substitution RT ln pO2 D GO2 , ln pMOB ln aMOA 1 RT ln 10 DGT MOA \u00ff DGT MOB B\u00ff A 2 D GO2 ; 2 where pO2 is the oxygen pressure, pMOB is the vapor pressure of the gaseous metal oxide MOB, aMOA is the activity of the (evaporating) condensed oxide MOA, DGT MOA is the free enthalpy of formation of condensed oxide MOA at a temperature T, DGT MOB is the free enthalpy of formation of condensed oxide MOB at the temperature T, D GO2 is the oxygen potential of the oxide system, and R 8:314 J mol\u00ff1 K\u00ff1 is the universal gas constant. The pressure of the gaseous component of the fuel can be calculated if we know the thermodynamic quantities in the right-hand side of Eqn (2). In the first approximation, the evaporation temperature is found by detecting the instant of emergence of a gas jet from the heated surface and by applying the laws of gas dynamics to the expansion of the gaseous fuel in the vacuum. Figure 2 [3, 4] plots ln p as a function ofT obtained in experiments with uranium dioxide UO2, where p is the gas pressure in the medium above the heated fuel. Despite the scatter of the data, mostly due to defects in the material, we note some degree of curving of the line in the region of higher pressures at increased temperatures. A high evaporation rate at extremely high temperatures alters the composition of the surface layer, which is not given enough time for recovering via diffusion. This change results in a driven uniform evaporation, at which vapor composition does not differ from the composition of the fuel. The curve of driven uniform evaporation 1 (see Fig. 2) was fit to experimental points. Then curve 1 was normalized to the ratio O=U 2:00, which gave us curve 2. In fitting the data, thermodynamic equilibrium was assumed, the Rand\u00b1Breitung method was used [3, 4], and we took into account the maximum possible error in the determination of free energies of formation of different fragmentation products suggested in [5]. After normalizing, curve 2 agrees almost completely with the initial data obtained in the temperature range 4400\u00b14700 K [6, 7]. The need to measure deviations from linearity in the dependence p T of pressure on the heating temperature stems from the need to obtain relations between the results of measurements of pressure at temperatures up to 4700 K and the range of possible values of the critical point for the heating of uranium\u00b1plutonium oxide" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002948_tfuzz.2021.3089053-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002948_tfuzz.2021.3089053-Figure6-1.png", + "caption": "Fig. 6. Pendulum physical diagram.", + "texts": [ + " Consider the dynamic model of a simple pendulum ml2\u03b8\u0308 = \u2212mgl sin(\u03b8)\u2212 fv \u03b8\u0307 +K\u03c4(t), (36) where \u03b8(t) \u2208 [\u22122\u03c0, 2\u03c0] is the angular position, \u03b8\u0307(t) \u2208 R is the angular velocity, \u03c4(t) \u2208 R is the control input, t \u2208 R\u22650 its time, m = 0.118 kg is the load mass, l = 0.193 m defines the pendulum length, fv = 2 \u00d7 10\u22122 Nm/s is the viscous friction coefficient, g = 9.81 m/s2 is the gravity constant acceleration, and K = 0.25 Nm is the torque constant. The simple pendulum scheme with their components is depicted in Fig. 6. Let x1(t) = \u03b8(t) and x2(t) = \u03b8\u0307 be the state variables and redefine u = K\u03c4(t)/ml2, \u03b3 = g/l and \u03c6 = fv/ml 2, the system (36) can therefore be represented in the state-space form as x\u0307 = f(x) + g(x)u, (37) where x = (x1, x2)T , f(x) = (x2,\u2212\u03b3 sinx1 \u2212 \u03c6x2)T and g(x) = (0, u)T . The control objective is to generate a periodic motion around the origin (x1, x2)T = (\u03b8, \u03b8\u0307)T = (0, 0)T \u2208 R2. However, the origin of the undriven system (u = 0) is an asymptotically stable equilibrium point. Therefore the natural motion of the pendulum does not keep persistent oscillations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002438_012190-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002438_012190-Figure3-1.png", + "caption": "Figure 3. Case 1 of the domain, region and sensor.", + "texts": [ + " , \ud835\udcc9)\u2016(\ud835\udc3b1(\ud835\udf14))\ud835\udc5b = 0, where { \ud835\udf15\ud835\udccc \ud835\udf15\ud835\udc61 (\ud835\udf011, \ud835\udf012, \ud835\udcc9) = \ud835\udf152\ud835\udccc \ud835\udf15\ud835\udf092 (\ud835\udf011, \ud835\udf012, \ud835\udcc9) + ((1 + \ud835\udc3b\ud835\udc34\ud835\udc45\ud835\udc3a) \ud835\udccc (\ud835\udf011, \ud835\udf01, \ud835\udcc9) +(1 \u2212 \ud835\udc3b\ud835\udc34\ud835\udc45\ud835\udc3a) \ud835\udf15\ud835\udc651 \ud835\udf15\ud835\udf092 (\ud835\udf011, \ud835\udf012, \ud835\udcc9) + (\ud835\udc3b\ud835\udc34\ud835\udc45\ud835\udc3a 2 \u2212 1) (\ud835\udf011, \ud835\udf012, \ud835\udcc9)) \ud835\udcac \ud835\udccc(\ud835\udf011, \ud835\udf012, 0) = \ud835\udccc0 (\ud835\udf011, \ud835\udf012) \u2127 \ud835\udccc (\ud835\udf021, \ud835\udf022, \ud835\udcc9) = 0 \u0398 (25) Reflect the process system well-defined by { \ud835\udf15\ud835\udccd2 \ud835\udf15\ud835\udc61 (\ud835\udf011, \ud835\udf012, \ud835\udcc9) = \ud835\udf152\ud835\udccd2 \ud835\udf15\ud835\udf092 (\ud835\udf011, \ud835\udf012, \ud835\udcc9) + \ud835\udccd2(\ud835\udf011, \ud835\udf012, \ud835\udcc9) \u2212\ud835\udccd1(\ud835\udf011, \ud835\udf012, \ud835\udcc9) \ud835\udcac \ud835\udccd2(\ud835\udf011, \ud835\udf012, 0) = \ud835\udccd02(\ud835\udf011, \ud835\udf012) \u2127 \ud835\udccd2(\ud835\udf021, \ud835\udf022, \ud835\udcc9) = 0 \u0398 (26) with \u2127 = (0,1) \u00d7 (0,1) and the output function Iraqi Academics Syndicate International Conference for Pure and Applied Sciences Journal of Physics: Conference Series 1818 (2021) 012190 IOP Publishing doi:10.1088/1742-6596/1818/1/012190 \ud835\udcb4(\ud835\udc61) = \ud835\udc9e \ud835\udccd1(. , \ud835\udcc9) (27) Let \ud835\udf14 = (\ud835\udefc1, \ud835\udefd1) \u00d7 (\ud835\udefc2, \ud835\udefd2) be studied region provided with \ud835\udf11\ud835\udc56\ud835\udc57(\ud835\udf011, \ud835\udf01) = 2 \u221a(\ud835\udefd1\u2212\ud835\udefc1)(\ud835\udefd2\u2212\ud835\udefc2) \ud835\udc46\ud835\udc56\ud835\udc5b \ud835\udc56\ud835\udf0b ( \ud835\udf011\u2212\ud835\udefc1 \ud835\udefd1\u2212\ud835\udefc1 ) \ud835\udc46\ud835\udc56\ud835\udc5b \ud835\udc57\ud835\udf0b ( \ud835\udf012\u2212\ud835\udefc2 \ud835\udefd2\u2212\ud835\udefc2 ) and (28) \ud835\udf06\ud835\udc56\ud835\udc57 = \u2212( \ud835\udc562 (\ud835\udefd1\u2212\ud835\udefc1) 2 + \ud835\udc572 (\ud835\udefd2\u2212\ud835\udefc2) 2)\ud835\udf0b 2, \ud835\udc56, \ud835\udc57 \u2265 1 (29) 5.1. Case 1 Reflect case 1 in rectangular sensor supports which is illustrated in \u201cFigure 3\u201d and characterize by equations (26)-(27). So the measuring output is specified by \ud835\udcb4(\ud835\udcc9) = \u222b\ud835\udc37 \ud835\udccd2(\ud835\udf011, \ud835\udf012, \ud835\udcc9)\ud835\udc53(\ud835\udf011, \ud835\udf012) \ud835\udc51\ud835\udf011 \ud835\udc51\ud835\udf012, (30) wherever \ud835\udc37 \u2282 \u2127, is the position of the sensor in zone type. So that the sensor (\ud835\udc37, \ud835\udc53) may be enough for deriving an \ud835\udc34\ud835\udc45\ud835\udc3a\ud835\udc45\ud835\udc42-observer, and \u2203 \ud835\udc3b\ud835\udc34\ud835\udc45\ud835\udc3a \u220b (\ud835\udc9c22 \u2212 \ud835\udc3b\ud835\udc34\ud835\udc45\ud835\udc3a\ud835\udc9c12) creates \ud835\udc34\ud835\udc45\ud835\udc3a-stable. Thus we have lim \ud835\udc61\u2192\u221e \u2016(\ud835\udccc (\ud835\udf011, \ud835\udf012, \ud835\udcc9) + \ud835\udc3b\ud835\udc34\ud835\udc45\ud835\udc3a \ud835\udc652(\ud835\udf011, \ud835\udf012, \ud835\udcc9)) \u2212 \ud835\udccd (\ud835\udf011, \ud835\udf012, \ud835\udcc9)\u2016(\ud835\udc3b1(\ud835\udf14))\ud835\udc5b = 0, where { \ud835\udf15\ud835\udccc \ud835\udf15\ud835\udc61 (\ud835\udf011, \ud835\udf012, \ud835\udcc9) = \ud835\udf152\ud835\udccc \ud835\udf15\ud835\udf012 (\ud835\udf011, \ud835\udf01, \ud835\udcc9) + ((1 + \ud835\udc3b\ud835\udc34\ud835\udc45\ud835\udc3a) \ud835\udccc(\ud835\udf011, \ud835\udf01, \ud835\udcc9) +(1 \u2212 \ud835\udc3b\ud835\udc34\ud835\udc45\ud835\udc3a) \ud835\udf15\ud835\udc652 \ud835\udf15\ud835\udf012 (\ud835\udf011, \ud835\udf01, \ud835\udcc9) + (\ud835\udc3b\ud835\udc34\ud835\udc45\ud835\udc3a 2 \u2212 1)( \ud835\udf011, \ud835\udf012, \ud835\udcc9)) \ud835\udcac \ud835\udccc (\ud835\udf011, \ud835\udf092, 0) = \ud835\udccc0(\ud835\udf011, \ud835\udf012) \u2127 \ud835\udccc (\ud835\udf011, \ud835\udf012, \ud835\udcc9) = 0 \u0398 (31) Proposition 5.1. Supposing \ud835\udc37 = [\ud835\udf0101 \u2212 \ud835\udc591, \ud835\udf0101 + \ud835\udc591] \u00d7 [\ud835\udf0102 \u2212 \ud835\udc592, \ud835\udf0102 + \ud835\udc592] \u2282 \u2127 as in Fig.3. Thus the system (31) is not \ud835\udc34\ud835\udc45\ud835\udc3a\ud835\udc45\ud835\udc42-observer for the systems (26)-(30), if for any \ud835\udc560, \ud835\udc560 (\ud835\udf0101 \u2212 \ud835\udefc1) (\ud835\udefd1 \u2212 \ud835\udefc1) and \ud835\udc560 (\ud835\udf0102 \u2212 \ud835\udefc2) (\ud835\udefd2 \u2212 \ud835\udefc2) \u2044\u2044 is rational number and \ud835\udc53 stays symmetric around the \ud835\udc6501 = \ud835\udf0901. Proof. Supposing \ud835\udc560 = 1, (\ud835\udf0101 \u2212 \ud835\udefc1) (\ud835\udefd1 \u2212 \ud835\udefc1) and \ud835\udc560(\ud835\udf0102 \u2212 \ud835\udefc2) (\ud835\udefd2 \u2212 \ud835\udefc2)\u2044\u2044 \u2208 \ud835\udc44, then there exists \ud835\udc570 \u2265 1 such that \ud835\udc46\ud835\udc56\ud835\udc5b (\ud835\udc570\ud835\udf0b\ud835\udc501 \ud835\udefd1\u2044 \ud835\udefc1) = 0. But Iraqi Academics Syndicate International Conference for Pure and Applied Sciences Journal of Physics: Conference Series 1818 (2021) 012190 IOP Publishing doi:10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003341_s11665-021-05551-4-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003341_s11665-021-05551-4-Figure7-1.png", + "caption": "Fig. 7 The testing principle and experimental scene of measuring residual stress; (a) the testing principle; (b) the experimental scene", + "texts": [ + " Before testing the residual stresses along the axial direction, the materials need to be firstly electropolished. Then, the materials surface is thinned layer by layer. Finally, the residual stresses along the axial direction can be measured. The detailed measurement method based on x-ray diffraction can be found in literature (Ref 28). Based on the above analysis, the main measurement parameters of the stress meter were given in Table 2 (Ref 29), and the main laser testing parameters were given in Table 3. Figure 7 shown the testing principle and experimental scene of measuring residual stresses. When the power density was 3.2 GW/cm2, the residual stresses along the radial direction were measured by selecting 4 points, and the residual stresses along the axial direction were also measured by selecting 8 points. Finally, the measured results of the residual stresses along the radial and axial direction were obtained, as shown in Fig. 8. According to the Fig. 8a, it was observed that the simulated values were close to the measured results, and they had similar distribution trend along the radial direction of the sample surface by comparing experimental results with computational values" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001810_iccas.2008.4694328-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001810_iccas.2008.4694328-Figure5-1.png", + "caption": "Fig. 5 Manufactured WHC", + "texts": [ + " This controller sends the optimal trajectory data received from WHC to each joint motor of the humanoid lower body at sampled time, and the sensor data (an accelerometer and a gyro sensor) of robot are transferred to LN2415SBC board. The specification of the LN2413SBC board which implements WHC is written in Table 1. - Upon a user\u2019s walking command, the information on the corresponding optimal joint motor trajectories calculated by WHC is transferred to the ATmega128 chip. - One can order the humanoid to walk or turn by using direction switches and a joystick, and command buttons. - An haptic actuator delivers walking status to the user by vibrating its motor. 4.2 Haptic Controller Fig. 5 shows manufactured WHC that implements the various functions stated above. We employ genetic algorithm (GA) as a basic optimization method for attaining thirteen optimal parameters of blending polynomials that describe the joint trajectories. The stride length is fixed as mS 1.0= and the stride cycle T is changed from 2 to 4 seconds for simulation of various-speed walking. As T increases, walking speed decreases, and vice versa. Fig. 6 illustrates the simulation result of biped walking for three representative walking cycles" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000701_1-4020-2204-2_29-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000701_1-4020-2204-2_29-Figure1-1.png", + "caption": "Figure 1. (a) 3D Biped Robot view (b) Moving Frames (Mita et al, 1984)", + "texts": [ + " Gon\u00e7alves and Zampieri (2003) used a recurrent neural network (RNN) to determine the trunk motion for a br, based on the ZMP criterion, to plan a stable gait for 10 DOF br that has a trunk like as inverted pendulum. So a RNN is trained to determine a compensative trunk motion that makes the actual ZMP get closer to the planned ZMP. Determination of kinematic and dynamic models for biped robots can be carried out in different ways. Taking the biped robot RB-1 as an example, constituted of seven bodies, so called right foot - link 1, right leg - link 2, right thigh - link 3, pelvis - link 4, left thigh - link 5, left leg - link 6 and left foot - link 7, Fig. 1(a). Li is the length of each robot link, ai is the localization of the COG of each link and, A, B, C, D, E, F, G, H, I and J are the geometric centers of each joint. All the six joints between the links have rotation of 1 DOF each, not considering the friction between them. An independent actuator moves each joint. Then, the position of point I in relation to the point A is given by (Campos et al, 2000): 4 754321643215213121 5432164321521312 )sen()cos()cos()cos( )sen()sen()sen()sen( Lz LqqqqqLqqqqLqqLqLLy qqqqqLqqqqLqqLqLx AI AI AI \u2212= \u2212\u2212\u2212++\u2212\u2212++\u2212+++= \u2212\u2212+++\u2212++++\u2212\u2212= (1) Considering that both feet remain parallel to the ground during all movement, where qi (i =1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001057_20080706-5-kr-1001.01557-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001057_20080706-5-kr-1001.01557-Figure1-1.png", + "caption": "Fig. 1. The basic principle of the proposed localization system.", + "texts": [ + " Sections 5, 6, and 7 analyze the robustness of the proposed least squares velocity estimation against measurement noise, partial malfunction, and imprecise installation. To explain the basic principle of the proposed localization method, let us consider a regular triangular array of optical mice attached at the bottom of a mobile 978-3-902661-00-5/08/$20.00 \u00a9 2008 IFAC 9209 10.3182/20080706-5-KR-1001.0769 robot. Generally, the traveling pattern of a mobile robot can be specified in terms of the location of ICR (Instantaneous Center of Rotation) on the plane. For three different traveling patterns of a mobile robot, Fig. 1 shows the linear velocities observed by a regular triangular array of optical mice. When a mobile robot is rotating with ICR apart from the center of a mobile robot as shown in Fig. 1a), three velocity vectors are different in both direction and magnitude. When a mobile robot is moving straight as shown in Fig. 1b), corresponding to the case of ICR at infinity, three velocity vectors become the same in both direction and magnitude. When a mobile robot is rotating with ICR coincident with the center of a mobile robot as shown in Fig. 1c), three velocity vectors become different in direction but the same in magnitude. Theses observations tells that a different traveling pattern of a mobile robot results in a set of different velocity readings of an array of optical mice. Reversely, it is possible to estimate the linear and angular velocities of a traveling mobile robot from the velocity readings of optical mice. Assume that N optical mice are installed at the vertices, P i, i= 1,\u2026,N , of a regular polygon that is centered at the center, O b , of a mobile robot traveling on the xy plane" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002752_012051-Figure10-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002752_012051-Figure10-1.png", + "caption": "Figure 10. Fragment of the stage of comparing the coordinates of the points of the model.", + "texts": [ + " The darker areas on the sides of the tooth space show the modified faces. The obtained models were compared with three-dimensional models obtained in the KissSoft software without modification and zero tolerances for the values of the addenum and dedendum diameters. The models were aligned along the axes of the gears, along the end face, and along the involute section of one of the faces of the teeth. The error of the models was estimated by the difference in distances between points located on the same diameter of each of the models. Figure 10 shows a snippet of measuring the distance between points. Satisfactory results were obtained (the distance between the points did not exceed 1 \u03bcm for the section of the tooth profile without modification). The accuracy of the modification assignment was verified by sketching the tooth profiles. Distance measurements on the sketch were carried out along the normal between the modified and true involute profile. The error should not exceed 10-5 mm. The International Conference on Aviation Motors (ICAM 2020) Journal of Physics: Conference Series 1891 (2021) 012051 IOP Publishing doi:10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000088_j.jsv.2006.01.044-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000088_j.jsv.2006.01.044-Figure5-1.png", + "caption": "Fig. 5. The models of the imitated crankshaft (a) The solid element model; (b) the beam element model.", + "texts": [ + " Using the principle of virtual work, the stiffness matrix of the cracked element can be written as \u00bdKec \u00bc \u00bdT \u00bdCc 1\u00bdT T, (10) where [T] is the transformation matrix \u00bdT T \u00bc 1 0 0 0 0 0 1 0 0 0 0 0 0 1 0 l 0 0 0 1 0 0 0 0 0 0 1 0 l 0 0 0 1 0 0 0 0 0 0 1 0 0 0 0 0 1 0 0 0 0 0 0 1 0 0 0 0 0 1 0 0 0 0 0 0 1 0 0 0 0 0 1 2 666666664 3 777777775 . In order to validate the current crack element and the approach for cracked crankshaft model, an imitated crankshaft (not an actual crankshaft) including a crack is applied. Fig. 5(a) shows the finite element model meshed by solid elements, and a slant crack is contained in the left crank web. The total of elements is 8646, and the total of nodes is 16,428. Fig. 5(b) shows the model meshed by the current cracked element and 3-node Timoshenko beams. The total of elements is 14, the total of nodes is 28, and the element 5 is a crack beam element. The dimensions of the beam elements are listed in Table 5. The influences of the crack depth on the natural frequencies of the two models are calculated respectively, the results of first four natural frequencies are shown in Fig. 6. From Fig. 6, it can be seen that the natural frequencies of the two models are approximate, and the changes due to the crack depth are coincident in the rough" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002545_tcst.2021.3064801-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002545_tcst.2021.3064801-Figure2-1.png", + "caption": "Fig. 2. Gradual but cumulatively significant changes in trench geometry.", + "texts": [ + " These primitive-based methods are shown to be effective with the advantage of computational simplicity and suitable for many SC applications. However, they are generally sensitive to noise, which could cause the measured signal to cross the threshold. A unified metric as opposed to thresholds can efficiently quantify command changes that indicate subgoal transitions. Subgoals for a given task can change significantly over time after many repeated cycles of the task. For example, consider again the excavator trenching task. Fig. 2 illustrates a gradual Authorized licensed use limited to: East Carolina University. Downloaded on June 29,2021 at 06:02:49 UTC from IEEE Xplore. Restrictions apply. but cumulatively significant change of the trenching area and depth. These changes typically can be adapted by making small adjustments to the existing nominal subgoals during each task cycle while having the operation patterns remain largely the same. These subgoal adjustments are essential for improving the efficiency of BSC. However, existing BSC methods consider only nominal subgoals without making any adjustments resulting from task operation and operator command changes over time" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002118_s11071-020-06158-5-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002118_s11071-020-06158-5-Figure4-1.png", + "caption": "Fig. 4 The region of sliding As (shaded grey) for different values of s = h/R and wave frequency \u03a9(1,1) (first row) and \u03a9(2,2) (second row). The region of jumping and sticking is shaded green and white, respectively", + "texts": [ + " These annular regions will be the focus of the Gedanken experiments and numerical simulations discussed later in this paper. It is to be noted that the motion regions for wave frequencies \u03a9(2,1) and \u03a9(3,1) are shown in Fig. 2. Such a family of wave frequencies exciting only those eigenmodes of the plate where the nodal circle is absent, and is referred as wave frequency of type \u201cA\u201d hereafter. The motion regions as obtained for wave frequencies \u03a9(1,1) and \u03a9(2,2), which excite eigenmodes having nodal circles, are shown in Fig. 4. Such a family of wave frequencies is referred as wave frequency of type \u201cB\u201d hereafter. Also, the variation of As with the nondimensional parameter s is observed to be the same as shown in Fig. 3 for such frequencies. However, as observed fromFig. 4, the jumping region is disjoint and separated by the nodal circle. Therefore, the jumping mode may occur even when the particle is placed close to the plate centre, unlike that observed in Fig. 2 where it occurred along the plate periphery. Also, it is impos- sible to define the concentric circles with radii r j , rs1 and rs2 as done previously. With increasing s, the similar phenomena of necking (Fig. 4b, f), splitting (Fig. 4c, g) and subsequent collapse or shrinkage (Fig. 4d, h) of the multiple sticking regions are observed. In Sect. 2, it is observed that if the particle is placed with a zero initial velocity at a radial location r \u2264 r j , the jumpingmode is eliminated when excited at certain wave frequencies. Also, the annular regions identified on the plate surface showed that the particle will exhibit either sliding or sticking\u2013sliding. In this section, the dynamics of particle placed in such annular regions is studied. The in-plane velocity of the material point coincident with the particle position coordinates rp and \u03c6p is denoted as vb and defined as, vb(rp, \u03c6p, t) = vbrer + vb\u03c6e\u03c6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000653_20070829-3-ru-4911.00040-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000653_20070829-3-ru-4911.00040-Figure1-1.png", + "caption": "Fig. 1. Lay-out of the \u201cHelicopter\u201d.", + "texts": [ + " Two motors with propellers mounted on the helicopter body can generate a force proportional to the voltage applied to them. The force, generated by the propellers, causes the helicopter body to lift off the ground and/or to rotate about the pitch axis. All electrical signals to and from the arm are transmitted via a slipping with eight contacts. The system is also equipped with a motorized lead screw that can drive a mass along the main arm in order to impose known controllable disturbances (the so-called Active Disturbance Option, ADO). Following notation is used through the paper (see Figure 1): \u03b8(t) \u2013 pitch angle 1 ; \u03b5(t) \u2013 elevation angle; \u03bb(t) \u2013 travel angle; vf (t), vr(t) \u2013 control voltages of the \u201cfront\u201d (conditionally) and the \u201crear\u201d motors; u(t) \u2013 pitch torque command signal w(t) \u2013 normal force command signal (used for elevation/travel control); ff (t), fr(t) \u2013 tractive forces of the \u201cfront\u201d and \u201crear\u201d propellers. These forces are produced respectively by the control voltages vf (t) and vr(t), applied to the front and rear motors. The motor control voltages have saturation level 5 V on magnitude" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000551_2007-01-1309-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000551_2007-01-1309-Figure7-1.png", + "caption": "Figure 7: FEM analysis of shift lever socket", + "texts": [], + "surrounding_texts": [ + "The control electronics for the actuator comprises a circuit board on which nearly all required components for the electrical functions are mounted. The actuator has five electrical external interfaces: Kl30, Kl31, CAN-High, CAN-Low, Aux Supply for powering the auxiliary motor. A 16-bit microcontroller is used. In addition to the internal watchdog of the microcontroller, an external watchdog is also employed for monitoring; no second microcontroller is used for monitoring purposes. The safety function is guaranteed through the interaction with other control devices at system level. An H-bridge is used to control the main motor. The Hbridge is supplemented with a measuring circuit for recording the H-bridge current. In a failure incident case, the auxiliary motor is controlled externally. After a failure incident, or for diagnostic purposes, the actuator also has to be able to control the auxiliary motor. In this case it will suffice if the auxiliary motor turns in one direction. The heat load resulting from electrical losses in a Shiftby-Wire application is rather small, since driving mode changes only rarely occur. Most losses occur in the voltage regulator unit and in the microcontroller. This aspect, in addition to the size of the actuator, allow operation at ambient temperatures of 120 \u00b0C , and even at 140 \u00b0C for a short time period; specific heat dissipation measures for the electronics are not necessary. But due to the temperature requirements, circuit board materials with augmented glass transition point are used. Moreover, through-hole solder connections are not used. Electrical contacting to the in-molded lead frame elements in the housing is done via press-fit connections. The entire circuit board is conformal coated as protection against the expected build up of grit and residue in the actuator as a result of the mechanic wear." + ] + }, + { + "image_filename": "designv11_83_0000061_iros.2004.1389871-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000061_iros.2004.1389871-Figure5-1.png", + "caption": "Figure 5. Eight-path off-line task.", + "texts": [ + "bus, for motion at the height of the top location, (10) can he M e r simplified and rewritten as (U), where the superscript tp of c: stands for \u201ctop location\u201d. For the movement at the height of the target location, (IO) can be further simplified and rewritten as (13), where the superscript rg of C: represents \u2018 k g e t location\u201d. (12) SX, -c:(f;) and SA\u2019, -G:(F,) cTX>=G:(f;) and SX, -G:(F,) (13) Therefore, eight relationships hold hetween the image feature data and differential location data relative to the manipulator\u2019s end-effector. Eight paths along which the end-effector can move are proposed, as illustrated in Fig. 5, to determine these relationships. As shown in Fig. 5, the end-effector is individually driven along each path, stating at the dot and following the arrow that indicates the direction of motion; the end-effector stops at every specified distance h m the dot. Paths 1 and 2 represent translations along the \u2018x and \u2018Y axes of the end-effector frame, respectively, both at the height of the top location. Path 3 represents the translation along the \u2018z axis of the end-effector tame. Path 4 represents rotation about the b~ axis of the end-effector frame at the top location", + " (19) (20) Then, the end-effector is commanded to move a short distance (in this study, 10 cm along the \u2018z axis of the endeffector frame) to grasp the workpiece. V. EXPERIMENTATION The workpiece to be picked up by the manipulator\u2019s end-effector is a rectangular parallelepiped with a volume of 68mmx30mmxsSmm. The zoom of the camera is adjusted to the wide-angle view, with a horizontal angle of view of around 46\u2018. However, the gripper\u2019s fingers are still not in the field of vision. An auto-focus function is adopted to maintain the sharpness of the workpiece image. As illustrated in Fig. 5 , the eight-path off-line task must be performed according to the LSE algorithm lo approximate the mappings between image feature space and Cartesian space. Equations (21) through (28) specify the resulting relationship between AX, and F; expressed in the form of (17) for each path. AX, % GY(F;) = 292.733137-0.755345F; (21) AX, s Cy(<) =-133.760361+0.725321F, (22) P- = Jsx: + sx: + sx: 0- = Jsx: + sx: + sx: AX, P G,\u2019(FJ = + .,F, + + $8, + to,<$ (23) where ma =-1112.694699986178, 0, =O,OOCQ2197638464529337 s, =7" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001904_2008-36-0289-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001904_2008-36-0289-Figure3-1.png", + "caption": "Figure 3 \u2013 Oil film thickness for (3a) flat bearing and (3b) axial profiled bearing.", + "texts": [ + " The transition from the edges to central region was considered a flat ramp with length of 5 mm. Flat bearing (figure 2a) presented maximum oil film pressure of 176 MPa on the bearing edges with contact pressure of 74 MPa while the axial profiled bearing showed completely different shape for oil film pressure distribution with maximum value of 193 MPa localized on the central region of the bearing length and minimizing contact pressure. Another key feature considered for bearing performance is related to the oil film thickness presented on the figure 3. Once more is clear the advantage of the axial profiled bearing, even though it (3b) presented lower film thickness on central region when compared with regular flat bearing (3a) the axial profiled bearing (3b) showed minimum oil film thickness of 0.4 \u03bcm on the edges while the regular flat bearing showed pronounced metal-to-metal contact on the edges (3a). Even though it is not explored in details here the results presented on figures 2b and 3b are the best option among four different versions assessed by numerical simulation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002352_j.matpr.2021.01.408-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002352_j.matpr.2021.01.408-Figure1-1.png", + "caption": "Fig. 1. Block diagram of the experimental setup.", + "texts": [ + " From the literature it was found that very limited study have been done to co-relate the traction behaviour and vibrational signal during the running mode of roller bearing. The present study aims to develop the relationship between the traction behaviour and vibrational signal of the roller bearing. The temperature monitoring with K-type thermocouple and wear surface of the bearing was characterized with the help of scanning electron microscope to understand the bearing behaviour under loading conditions. The objective of the experiment is to detect the surface fatigue wear defects and their effects on the performance of roller bearing. Fig. 1 shows the experimental setup which is consists of a 5HP three phase induction motor used to drive the shaft through direct coupling, the shaft is supported by support bearing and test bearing. A radial load of 1.5 kN is applied on the test bearing through load bearing which is connected to manual loading via rope and pully arrangement. The experimental setup is mounted on a massive concrete block anchored by I cross section beams. A KS78.100 piezoelectric uniaxial accelerometer and DEWE-43A versatile USB data acquisition system is used to acquire the vibration signals; the accelerometer is mounted on the top of test bearing housing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003132_978-981-16-0119-4_20-Figure19.9-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003132_978-981-16-0119-4_20-Figure19.9-1.png", + "caption": "Fig. 19.9 Stiffness behavior", + "texts": [], + "surrounding_texts": [ + "A topology optimization result provides the guidance for developing a frame. A frame representing \u2018material, geometry, and manufacturing\u2019 feasibility needs to be created (see Fig. 19.7). The material, for example, is a standard material normally used on bicycles (\u2018IS 3074:2005\u20196). Certain additional members are added to connect members laterally. These are positioned in a manner that they do not interfere with other aggregates in the product.\nFor virtual verification, see Figs. 19.8, 19.9, and 19.10. The FE workbench within \u2018FreeCAD\u2019 is used. For meshing \u2018Netgen\u2019 and for solving \u2018CalculiX\u2019 are used (embedded inside the workbench). Two aspects Viz. \u2018stiffness\u2019 and \u2018propensity to yield\u2019 are assessed. The results are reasonable for the given conditions.\nThe \u2018VonMises\u2019 stresses provide a \u2018better correlation with experimental behavior than \u2018Tresca\u2019 yield criterion\u20197 and has been used to assess the developed stresses.\n6Bureau of Indian Standards. See: https://bis.gov.in. 7Teaching and learning packages of the \u2018University of Cambridge\u2019. See: https://www.doitpoms.ac. uk/tlplib/metal-forming-1/yield_criteria.php.", + "Amanufacturing team needs a drawing and specifications communicating the design (to the extent it impacts\u2014functional performance) (see Figs. 19.11 and 19.12).\nThe framework described above leverages technology to drive efficient design.", + "238 S. K. Mukherjee" + ] + }, + { + "image_filename": "designv11_83_0001278_s1068798x08110087-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001278_s1068798x08110087-Figure5-1.png", + "caption": "Fig. 5. Screw with clock indicator.", + "texts": [ + "= RUSSIAN ENGINEERING RESEARCH Vol. 28 No. 11 2008 INFLUENCE OF CONTACT PLIABILITY ON THE OPERATION 1065 The decrease in the distance between the left (fL) and right (fR) edges of the plate in the joint after applying the external loads is measured (Fig. 3) by two Gugenberger mechanical tensometers 1 (scale division 1 \u00b5m; base 20 mm); the change in size of the loaded screw section is recorded by two clock-type indicators 2 (scale division 1 \u00b5m). The attachment of the clock indicator to the screw is illustrated in Fig. 5. The joint is investigated with a force Fle = 0.5 kN applied to the end of the lever (Fig. 3). By rearranging two pins in the four holes of the lever, two loading configurations are obtained: (1) by a central force FA = Fle(L0 + L)/L (Fig. 6a); (2) by a central force FB = Fle and a tipping moment MB = L0Fle (Figs. 3 and 6b). The lever dimensions are as follows: L0 = 1000 mm; L = 100 mm. In the experiment, Fti = 5 and 10 kN. The screws are tightened by a dynamometric switch. After tightening the screws, the readings of the mechanical tensometers and the clock indicators are set to zero" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001457_med.2007.4433930-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001457_med.2007.4433930-Figure1-1.png", + "caption": "Fig. 1. Photo of a harmonic drive.", + "texts": [ + " This new gear concept used to be applied in aerospace and other specific applications but today it is used in many fields such as robotics, machine tools, medical equipment and automotive industry. Harmonic drives have high speed reduction and torque multiplication ratios using single stage and coaxial configuration of shafts. Other benefits include nearly zero backlash, small size and high torque transmission capacity due to the high number of teethes in contact. The harmonic drive is made up of three basic components: the wave generator, the flexspline and the circular spline as depicted in Fig. 1. The wave generator is an elliptical cam enclosed in an antifriction ballbearing assembly. It is inserted into the bore of the externally toothed flexspline. There are 2 less teethes on the flexspline than on the internally toothed circular spline. The flexspline is deformable and takes on the elliptical shape of the wave generator causing its external teeth to engage with the internal teeth of the circular spline at two opposite points. Thanks to this mechanism, the number of contacted teethes in harmonic drives is greater than in traditional gears" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000551_2007-01-1309-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000551_2007-01-1309-Figure2-1.png", + "caption": "Figure 2: Cam lock forces", + "texts": [], + "surrounding_texts": [ + "Copyright \u00a9 2007 SAE International\nIncreasing demands for design freedom in modern automotive cockpits stand in contrast to the still predominant mechanical coupling of the driving mode selector to the automatic transmission. Furthermore, tendencies to enhance comfort, such as \"keyless go\", are becoming ever more prevalent. A \"Shift by wire\" system is a viable solution fulfilling these requisites. A mechatronic module, optionally mountable to an existing transmission, incorporating the inter-functionality of actuation, sensing and control electronics will be presented. The mechatronic module must supply high torques within short switching times for the entire voltage range of the battery. Legal guidelines prescribe a redundant feature assuring that the vehicle can lock into the park position at all times. The mounting location on the transmission was decisive for the module design and choice of materials because it dictated severe temperature, vibration and sealing requirements. It will be shown how these requirements were satisfied within the extremely confined space envelope and with inherently high system power density.\nThis presentation examines Shift-by-Wire for torqueconverter transmissions in which the P, R, N and D drive modes are selected electrically rather than mechanically.\nA mechatronic series solution for an actuator in a Shiftby-Wire system will be described. In addition, the following topics will also be discussed in more detail:\n- Advantages of Shift-by-Wire systems - Comparison of a transmission add-on solution and a\nsolution integrated into the transmission - System, module envelope, environmental and\nstatutory requirements\nWith the current mechanical coupling between the transmission and selector lever, there is little freedom with respect to the position and size of the selector lever mechanics. With Shift-by-Wire, the driver interface can be fitted at almost any location or position, thus providing a new level of design freedom and flexibility. For\nexample, buttons, joystick and the like can be used. The space gained can be used for control elements, storage compartments, etc.\nWith Shift-by-Wire, however, new comfort and convenience functions can also be realized, which utilize the automated drive mode selection. For example, it is conceivable to have position P automatically selected when removing the key from the ignition instead of key release via the key lock function. Evidently this offers considerable benefits, particularly in conjunction with Keyless-Go. A number of other functions, such as support from a fully automatic parking assistant, are also possible.\nIn principle, two solutions are possible for a Shift-by-Wire actuator. As one option, the actuation can be integrated into the transmission (as an integrated solution), or alternatively, the actuation modules can be mounted outside on the transmission (as an add-on solution). Some topics which have to be considered before deciding between an integrated or add-on solution are discussed in the following.\n2007-01-1309\nMechatronics for \"Shift by Wire\" - A Technical Challenge\nKarl Smirra, Michael Ferstl and Thomas Eiting Siemens AG, Siemens VDO Automotive", + "Shift-by-Wire for existing transmissions: When implementing Shift-by-Wire for an existing transmission, extensive effort is required to fit the necessary actuation modules into the transmission. Accordingly, for existing transmissions, there is the add-on solution, where the actuator uses the same interface on the transmission, as for the mechanic transmission/selector lever coupling. Only the transmission housing has to be adapted in order to create a potential attachment point for the actuator.\nModule envelope and options: an integrated solution requires space inside the transmission, while an add-on solution needs the corresponding space on the outside. A check is necessary to determine which required space is more feasible. It has to be assumed that a transmission is not used only in Shift-by-Wire systems. The add-on solution ensures that no additional costs are incurred for the variants without Shift-by-Wire. This is not necessarily the case with the integrated solution\nHydraulic or electromotive: Hydraulic actuation modules are well adaptable for an integrated solution, while electromotive actuation modules are better suited for an add-on solution. Certain requirements can be easily resolved with a hydraulic actuator and others with an electromotive actuator. One example is being able to shift the driving mode without having the engine running, which can be performed more easily with an electromotive actuator.\nThe mechatronic series solution presented here is a transmission add-on solution with electromotive actuation.\nFor automatic transmissions with park position option, it is generally applicable that the ignition key can only be removed if the gear is locked in P (\"Key lock\") or if the transmission is locked in P as a direct result of the key being removed (\"Key P\").\nFor the series solution described here, the legal situation prescribes that even if the electronics or the electromotor fails (fault case) P must be securely set. In a fault case an independent control circuit, powered by a separate battery source, would deliver respective power for a predetermined duration. Due to the marginal power reserve of this auxiliary battery an energy storage function must be incorporated in the Module itself.\nFurthermore, even in a fault case a shift into P must be executable by driver command, e.g. as a result of the key being removed. An automatic shift into P is not allowable.\nIn order to guarantee a high level of availability, a driving mode change should also be possible with a battery\nvoltage of down to 6.5V; e.g. in the case when the engine is not running.\nIt is disturbingly noticeable for the driver if the time between selection of the driving mode (driver action) and the transmission traction point (transmission reaction) is too long. Therefore, the allowed shifting time between positions P and D was set to 0.3s.\nApart from the shifting time requirements (rpm of the actuator motor), the necessary shifting torque in the transmission is crucial to the design of the electromotor. A transmission driving mode change is associated with a rotation of the shift cam, resulting in a shift torque due to the cam lock spring.\nIn the P-mode the park lock engages the park lock disc, thereby immobilizing the vehicle. The mass of the vehicle as well as its pitch angle when parked will invariably increase the P-disengagement torque. The maximum shift torque encountered for transmission mode switching sums up to 8 Nm.\nThe driving mode of the transmission is determined by the position of the \u201cselector poppet\u201d in the hydraulic plate. This element must be very precisely positioned in order to set the hydraulically correct driving mode. Correspondingly, there are strict requirements for the positioning accuracy of the actuator.\nTheft protection is an important aspect of the described solution. Unauthorized persons should be prevented from shifting out of P; the vehicle must stay immobilized. A driving mode change of the transmission by part of the \"Shift-by-Wire\" module must in this case be rendered as difficult as possible. An \"override\" release from P in a fault case (e.g. by means of Bowden cable) is therefore not acceptable, even if it is technically feasible.", + "The actuator installation location in the vehicle justifies the major environmental requirements. The \"Shift-byWire\" Module will also be introduced into all-terrain vehicles.\nTemperature Range: the specified ambient air temperature is given from \u201340\u00b0C to 120 \u00b0C (shortduration up to 140 \u00b0C)\nVibration: maximum acceleration under temperature is 10 g (Sine wave)\nSealing classification: IP6K9K and immersion tightness in ice water\nDurability: 6000 hrs, or 300,000 switching cycles\nThe transmission/chassis periphery prescribes the envelope of the module. The actuator must fit to various transmissions in different vehicle platforms. The smallest resulting envelope defines the concept for the actuator and is determinant for the mechanical design and implementation. The envelope at the dynamic interface, especially in Y direction, proved to be exceptionally tight.\nIn normal operation, the lever is actuated according to the following functional sequence: the control electronics receives the driver command from the driver interface\ndevice. The electronics then powers the main motor, which drives the spindle via the gear train. A two-part carriage unit--comprising a driver nut carrier and a latching carriage--runs on the spindle. The carriage pin is engaged with the lever which is thereby displaced concurrent with the carriage unit translation. The lever position is constantly monitored and evaluated by a sensor on the circuit board. When the target position is reached, the functional sequence is terminated.\nIn a fault case incident, e.g. failure of the vehicle electrical system supply voltage, the failure operating mode is activated. In this case the module at first retains its present position until the driving position P is selected. Hereby, an external control unit can activate the \"auxiliary motor\" using an independent power source. This \"auxiliary motor\" drives an eccentric shaft via a reduction transmission which unlatches the safety catch of the two-part carriage unit. The compressed spring between the driver nut carrier and carriage is now free to push the lever to the \"P\" position (energy storage function). The module remains in this state until the failure is remediated. After failure rectification and recognition, the control electronics return the eccentric shaft back to the starting position, which is recorded by a Hall effect switch. The driver nut carrier is then driven towards the carriage, to which it then relatches. In conjunction, this operation recompresses the spring between carriage nut and carriage. The Module is now \"reset\" and ready for the normal operation described above.\nIn view of the complex envelope geometry it was evident a plastic housing with an in-molded lead frame would be considered the concept of choice for module encasement." + ] + }, + { + "image_filename": "designv11_83_0003376_icuas51884.2021.9476688-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003376_icuas51884.2021.9476688-Figure2-1.png", + "caption": "Fig. 2: Ryze Tello EDU UAV with OptiTrack markers connected to provide localization.", + "texts": [ + " These numbers of nodes are chosen to allow each algorithm to have the same number of states embedded in their nodes. A node for the simultaneous method contains a state for all of the robots. Each node in the sequential method only contains the state of a single robot. 871 Authorized licensed use limited to: University of Glasgow. Downloaded on August 18,2021 at 05:21:14 UTC from IEEE Xplore. Restrictions apply. For the experimental validation of both MK-RRT* variants, the palm-sized Ryze Tello EDU UAV, shown in Figure 2, was used. For the planners, the Tellos are approximated as spheres with a radius of 15cm, allowing for a small safety region around each platform. An OptiTrack Motion Capture (MoCap) system was used to provide real-time, high-precision localization, while a Kalman filter was used as a full state observer. Finally, a Model Predictive Controller (MPC) was used to drive the Tellos to the generated trajectories. The Tello allows reference velocity inputs on all three axes for its control. A linear model that describes the internal system dynamics of the Tello must be derived, since it is needed for both the planners and the controller" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002234_s00289-021-03570-8-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002234_s00289-021-03570-8-Figure3-1.png", + "caption": "Fig. 3 Die designed for radial orientation of short fiber", + "texts": [ + " While in this paper, a way to make short fiber get radial orientation in rubber matrix of SFRC has been studied and worked out. During the extrusion process, there are two important factors to make short fiber get radial orientation. One is velocity-gradient which makes the rubber material suffer shearing-field force and stretching-field force in die channel, as shown in Fig.\u00a02. Another is extrusion swelling which is a special character of polymer such as rubber. Therefore, considering both velocity-gradient and extrusion swelling, a type of damexpanding die has been designed, as shown in Fig.\u00a03. According to this structure, the extrusion pressure would be increased by the dam during extrusion process (C area in Fig.\u00a03a). So, the material would get high pressure before the dam (i.e., A area) and get high internal energy for expansion. Then, due to the channel space after dam (i.e., B area) becoming big suddenly, the material would expand after passing the dam even bigger than extrusion swelling. As a result, short fibers in rubber matrix would turn with the expansion process and form an angle with X axis, as shown in Fig.\u00a02. In addition, with the extrusion process going on, the angle would get larger by action of shearing-field force and stretching-field 1 3 force", + " This is the reason for designing the die as well. Furthermore, it can be concluded that the most important factors that affect short fiber radial orientation are half expansion angle \u03b1 and expansion ratio K from Formula (3). But the two factors should have a proper value. If the angle is smaller, the expansion process will be gradual, which will cut down the internal energy of material for expansion. If the angle is bigger, e.g., 90, there would be dead zones near the dam (i.e., the junction area of B and C, shown in Fig.\u00a03a), which has bad effects (1) d dK = ( n+1 n )||| z H ||| 1 n 1 \u2212 ||| z H ||| 1+ 1 n (cos 2 \u2212 1) tan + sin 2 K tan2 (2) ( n+1 n )||| z H ||| 1 n 1 \u2212 ||| z H ||| 1+ 1 n = C (3) d dK = C (cos 2 \u2212 1) tan + sin 2 K tan2 1 3 on extrusion and should be avoided. While for the expansion ratio, if it is smaller, which means expansion channel height is almost equal with dam clearance, there would not be enough space for the material to expand. As a result,\u00a0 the rotation\u00a0of short fibers in rubber matrix will be decreased\u00a0 as\u00a0 lack of expansion\u00a0 which leads to bad radial orientation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000551_2007-01-1309-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000551_2007-01-1309-Figure3-1.png", + "caption": "Figure 3: Packaging envelope in vehicle", + "texts": [], + "surrounding_texts": [ + "The actuator installation location in the vehicle justifies the major environmental requirements. The \"Shift-byWire\" Module will also be introduced into all-terrain vehicles. Temperature Range: the specified ambient air temperature is given from \u201340\u00b0C to 120 \u00b0C (shortduration up to 140 \u00b0C) Vibration: maximum acceleration under temperature is 10 g (Sine wave) Sealing classification: IP6K9K and immersion tightness in ice water Durability: 6000 hrs, or 300,000 switching cycles" + ] + }, + { + "image_filename": "designv11_83_0000610_bfb0119393-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000610_bfb0119393-Figure1-1.png", + "caption": "Figure 1. Mobile manipulator when n -- 6", + "texts": [ + " The point On+l, which is the center of this end-effector, is fixed relative to the frame ~n . 2.2. K inema t i c s equa t ions of t he mobi le m a n i p u l a t o r If we leave the wheels of the platform out of account we can identify the three generalized coordinates of the platform- defining its configuration - with the 1 Remark that the moving frames R0 and R I are parallel. three operational coordinates of this platform- defining its location. If we choose the three operational coordinates as: ~v = (~vt ~p2 ~p3) t = (z y #)t (see Fig. 1), the three generalized coordinates are defined by: qv = (qpl qp~ qp3) t = (z y #)t . With the previous abstraction it is necessary to suppose that the control of this platform is given by: u = (ul u2) t = (v w) t, in which v represents the linear velocity of the point O', on its trajectory in the plane (O, ~, y-'), and w the angular velocity of the platform, around the vector ~'. v is such that: v = b, where ~ is the curvilinear abscissa of the point O' on its path in the previous frame, and w = ~" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003132_978-981-16-0119-4_20-Figure19.2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003132_978-981-16-0119-4_20-Figure19.2-1.png", + "caption": "Fig. 19.2 Estimation of loads", + "texts": [ + " followed by \u2018interpretation of the topology into a manufacturable part\u2019 and ends with \u2018specifying the proposed design in the form of a drawing and specification sheet.\u2019 The final purpose of a \u2018product design\u2019 is practical usage. Manufacturing a product requires sharing a drawing that the manufacturer\u2019s team can use. For this study to be effectively used, it is assumed that a student has an understanding of introductory courses in \u2018Applied Mechanics\u2019 and \u2018Strength of Materials.\u2019 From a \u2018system level\u2019 standpoint, tractive effort is an important parameter needed for designing any vehicle (see Fig. 19.2). This effort is an input to estimate the pedaling force (see Fig. 19.3). As the handlebar controls the steering and supports the inertia of the rider, estimating the load on the handlebar is important (see Fig. 19.4). The saddle supports some rider inertia (see Fig. 19.5). To simplify the study, loads along the lateral direction of the bicycle are not included in this case study. For finding the tractive effort, a process explained by Krishnakumar4 has been adapted. Using mechanics, one can estimate the forces required on the pedal, once a value for the tractive effort on the rear wheel is known" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002360_012005-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002360_012005-Figure1-1.png", + "caption": "Figure 1. Fishtail open turn with rectilinear reverse movement, which is not parallel to the field boundary - movement from left to right.", + "texts": [ + " The purpose of the present work is to derive analytical dependences for determining the length of fishtail turns with a rectilinear reverse move that is not parallel to the field boundary, as well as for determining the width of the headland required to perform them in an irregularly shaped field and to analyze the influence of the angle between the direction of movement and the field boundary on the length of the non-working move and the width of the headland at different direction of movement of the unit in the field. A fishtail open turn and a fishtail closed turn with a rectilinear reverse movement in different directions of their performance in a field with an irregular shape are considered. To determine the length of the turn, a geometric method is used, in which the turn is represented by straight lines and arcs of a circle of equal radius. The different types of turns are presented in Figure 1, Figure 2, Figure 3 and Figure 4. The length of the turn is defined as the sum of the lengths of its geometric elements. The width of the headland required to perform the turn is defined as the sum of the segments perpendicular to the field boundary and depending on the elements of the turn. The symbols used in the figures are as follows: \u03b1 is the angle between the direction of working move of the unit and the boundary of the field; p. A \u2013 the beginning of the turn; p. B \u2013 the end of turn; p. O1, p", + " This can be avoided by using modern navigation systems. To perform the study, it is assumed that the angle between the direction of working move and the field boundary varies from 10\u00b0 to 90\u00b0. When machine-tractor unit crossed the boundary of the headland with the entire working width of the machine, the unit is brought to the transport position, a curvilinear movement is carried out around point O1, followed by reverse movement of the tractor to reaching the field boundary and again curvilinear forward movement around point O2 (Figure 1, (a)). Immediately afterwards, the machine is brought into working position at the boundary of the headland and the next working move begins. The length of the turn is measured from point A to point B and is determined by the dependence \ud835\udc59\ud835\udc47 = \ud835\udf0b\ud835\udc45 + \ud835\udc421\ud835\udc422 \u0305\u0305 \u0305\u0305 \u0305\u0305 \u0305 = \ud835\udf0b\ud835\udc45 + 2\ud835\udc45+\ud835\udc35 sin \ud835\udefd1 (1) The central angle \u03b21 is determined by the dependence \ud835\udefd1 = tan\u22121 ( 2\ud835\udc45+\ud835\udc35 \ud835\udc35 tan \ud835\udefc +2\ud835\udc59\ud835\udc4e ) (2) After substitution in equation (1) for lt is obtained \ud835\udc59\ud835\udc47 = \ud835\udf0b\ud835\udc45 + [(2\ud835\udc45 + \ud835\udc35)2 + ( \ud835\udc35 tan \ud835\udefc + 2\ud835\udc59\ud835\udc4e) 2 ] 1 2\u2044 (3) To determine the width of the headland, segments a, b and c are used, the lengths of which are respectively ICTTE 2020 IOP Conf. Series: Materials Science and Engineering 1031 (2021) 012005 IOP Publishing doi:10.1088/1757-899X/1031/1/012005 \ud835\udc4e = \ud835\udc59\ud835\udc4e . sin \ud835\udefc \u2212 0,5\ud835\udc35. cos \ud835\udefc (4) \ud835\udc4f = \ud835\udc45. cos \ud835\udefc (5) \ud835\udc50 = (\ud835\udc45 + 0,5\ud835\udc40) cos(\ud835\udefd1 \u2212 \ud835\udefc) (6) The width of the headland is \ud835\udc38 = \ud835\udc4f + \ud835\udc50 \u2212 \ud835\udc4e = (\ud835\udc45 + 0,5\ud835\udc35) cos \ud835\udefc + (\ud835\udc45 + 0,5\ud835\udc40) cos(\ud835\udefd1 \u2212 \ud835\udefc) \u2212 \ud835\udc59\ud835\udc4e . sin \ud835\udefc (7) field boundary at the end of the first curvilinear movement (Figure 1, (b)). The following dependencies are used to determine the width of the headland: \ud835\udc4e = 0,5\ud835\udc40. cos(\ud835\udefc \u2212 \ud835\udefd1) (8) \u210e = \ud835\udc3b. sin(\ud835\udefc \u2212 \ud835\udefd1) (9) \ud835\udc50 = \ud835\udc45. cos(\ud835\udefc \u2212 \ud835\udefd1) (10) \ud835\udc4f = \ud835\udc45. cos \ud835\udefc (5) \ud835\udc5d = 0,5\ud835\udc35. cos \ud835\udefc + \ud835\udc59\ud835\udc4e . sin \ud835\udefc (11) \ud835\udc38 = \ud835\udc4e + \u210e + \ud835\udc50 \u2212 \ud835\udc4f + \ud835\udc5d = = (0,5\ud835\udc40 + \ud835\udc45) cos(\ud835\udefc \u2212 \ud835\udefd1) + \ud835\udc3b. sin(\ud835\udefc \u2212 \ud835\udefd1) + (0,5\ud835\udc35 \u2212 \ud835\udc45) cos \ud835\udefc + \ud835\udc59\ud835\udc4e . sin \ud835\udefc (12) The transition between the modes of movement shown in Figure 1(a) and Figure 1(b) is performed at an angle \ud835\udefc = tan\u22121 ( 2\ud835\udc45+\ud835\udc3b.sin \ud835\udefd1 \ud835\udc3b .cos \ud835\udefd1+2\ud835\udc59\ud835\udc4e ), (13) in which the width of the headland in both variants is the same. In this way of movement after reaching the boundary of the headland, a curvilinear movement is performed until reaching the boundary of the field, after which a rectilinear reverse movement is performed and again a curvilinear movement forward to the boundary of the headland (Figure 2). At the end of the reverse movement, the tractor wheels may be outside the headland with width E, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000152_bf03227869-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000152_bf03227869-Figure2-1.png", + "caption": "Figure 2: Cut-away view of a truck- engine including turbo compound [2]", + "texts": [ + " The adjustability of the VNT allows the turbocharger to optimize itself on the fly to allow the engine to produce maximum torque with maximum fuel economy. 1.2 Turbocompound \u2013 Functional Description In the case of automotive or truck Diesel engines, \"Turbocompound\" means the introduction of an additional power turbine downstream or after the turbocharger. This second exhaust turbine is used to recovers a percentage of the thermical energy that would normally be lost through the engine\u2019s exhaust. The mechanical energy generated by the power turbine fed back to the engine\u2019s crankshaft via a sophisticated mechanical transmission, Figure 1 and Figure 2. The exhaust turbine is different from a standard turbocharger via the absense of a compressor stage. A Mechanically coupled system that is driven directly from a standard turbo charger might be also possible. Energy recovery using a secondary exhaust turbine and transmission can increase engine efficiency from 42% to as much as 46%. 2.1 Turbocompound \u2013 Rotor 2.1.1 Functional Description The turbine runs with max. speed up to 70,000 rpm. The gear ratio from the exhaust turbine to the Turbocompound intermediate shaft is 6:1, and the gear ratio from the intermediate shaft to the crankshaft is 5:1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002789_s00170-021-07207-y-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002789_s00170-021-07207-y-Figure5-1.png", + "caption": "Fig. 5 Three groups of the contact areas between silicon rubber heating sheets and base: contact area 1 between red part and base, contact area 2", + "texts": [ + " And Ar A \u00bc P D \u00f020\u00de where P is the contact pressure placed on two surfaces and D is the microhardness of the softer of the two materials. Therefore, the thermal contact conductance Tcc between the silicone rubber heating sheets and base, and the base and the natural granite are calculated based on the above method. In particular, the standard deviation of profile heights of base surfaces varies because of the different processing. Therefore, the contact areas between silicon rubber heating sheets and base are categorized into three groups: contact area 1, contact area 2, and contact area 3, as shown in Fig. 5. As shown in Fig. 6, the thermal model of the guideway base is developed using ANSYS Mechanical. Thermal conduction 20 22 24 26 28 30 0 1 2 3 4 5 6 7 h m / W 2 \u00b7 \u00b0C Temperature (\u00b0C) ha hb hc hd he hn 0 1 2 3 4 5 6 7 8 h m/ W 2 \u00b7 \u00b0C Temperature (\u00b0C) hf hg hh hi hj hk/hl hm 22 24 26 28 30 32 (a) (b) Fig. 4 The temperaturedependent heat transfer coefficients: a heat transfer coefficients of horizontal plates and horizontal cylinders; b heat transfer coefficients of vertical plates. between blue parts and base, contact area 3 between yellow parts and base Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002968_s00170-021-07401-y-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002968_s00170-021-07401-y-Figure1-1.png", + "caption": "Fig. 1 Forced cooling device for arc constricting from a front view and b oblique view", + "texts": [ + " This work aims at investigating the influences of constricted cold metal transfer (CMT) arc on the shaping, especially the size of deposition layers, and heat input in the WAAM and possible application of such arcs, with the expectation to improve the accuracy and to enhance the controllability of size for the additively manufactured parts. The experimental system consisted of a Motoman robot with the model of HY-HP6-A00, a Fronius CMTAdvanced 4000R power supply, an AcutEye high speed camera (Motion Pro Series), an electrical signal acquisition system, and a selfmade forced cooling device for arc constricting (FCDAC). The FCDAC, which is fixed on the torch (as shown in Fig. 1), was designed to constrict the arc and further to improve the controllability of size and surface shaping in CMT-based additive manufacturing (CMTAM). The main part of the FCDAC close to the arc is two hollow copper blocks with coolant passed through. The gap between the blocks and the height of the FCDAC can be modified to constrict the arc to different extent, thereby obtaining different arc shapes. The deposition material was the wire of ER304 stainless steel with a diameter of 1.2 mm. The nominal chemical composition of the filler material and the substrate is given in Table 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002657_tmag.2021.3073155-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002657_tmag.2021.3073155-Figure5-1.png", + "caption": "Fig. 5. Magnetic flux distributions of the C-MFSPM and FP-MFSPM machines.", + "texts": [ + " In order to evaluate the electromagnetic performance of the MFSPM machines with different winding configurations, a series of comparisons, in terms of no-load back-EMF, air-gap flux density, cogging torque, electromagnetic torque, and torque ripple, among the proposed FP-MFSPM machines and the existing C-MFSPM machines with 6/17 and 6/19 structures are conducted. Four MFSPM machines, namely 6/17 and 6/19 FP-MFSPM machines, and 6/17 and 6/19 C-MFSPM machines, are marked as \"6/17-FP\", \"6/19-FP\", \"6/17-C\", and \"6/19-C\". A. No-Load Performance Fig. 5 shows the no-load flux distribution of the 6/17 CMFSPM and FP-MFSPM machines, where the rotor position is that the fluxes flowing through phase A reach the maximum value, respectively. For the FP-MFSPM machine, the magnetic fluxes flowing into coil A1 are 0 ~ 180\u00b0, while that of the CMFSPM machine are 0 ~ 60\u00b0. Hence, the fluxes flowing into phase windings of 6/17 FP-MFSPM machines are much more than that of the C-MFSPM one. Fig. 6 presents the radial air-gap flux density (Br\u03b4) distribution of the 6/17 C-MFSPM and FP-MFSPM machines in the region of coils A1 and A2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002838_012021-Figure24-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002838_012021-Figure24-1.png", + "caption": "Figure 24. Total Deformation in Carbon Fiber", + "texts": [ + "804 MPa and 0.37819 MPa respectively shown in Figure 22. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.6.3. Strain Distribution The Max. And Min. Strain Distribution in Basalt Fiber is 0.00074223 and 0.0000057032 respectively shown in Figure 23. 3.7. Analysing Testing Result of Carbon Fiber 3.7.1. Total Deformation The Max. And Min. Total Deformation in Carbon Fiber is 0.68523 mm and 0 mm respectively shown in Figure 24. The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.7.2. Stress Distribution The Max. And Min. Stress Distribution in Carbon Fiber is 97.378 MPa and 0.28227 MPa respectively shown in Figure 25. 3.7.3. Strain Distribution The Max. And Min. Strain Distribution in Carbon Fiber is 0.0024296 and 0.000021423 respectively shown in Figure 26" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000237_pedes.2006.344346-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000237_pedes.2006.344346-Figure4-1.png", + "caption": "Fig. 4. The flux distribution in a 15 hp motor.", + "texts": [ + " AC magnetization curves for the teeth and core yoke and the DC magnetization curve. Two components were created in the stator core, first one was for the core yoke and material property was set to be as core yoke saturation cure, and for the second one material property was set to be as teeth saturation curve [3] as shown in Fig. 2. 1.60 1.40 1.20- _ 1.00 a. m 0.80 Xm0.80 0.60 FEM analyses have been done for the two three phase, induction motors. Specifications of the analyzed motors are given in Table I and Table II. Fig. 4 shows the flux distribution in the mentioned 15 hp motor. Table IV shows the computational and tested line currents (A) for 5hp and 15hp 4P three phase induction motors at fullload. It is observed from the table that the percentage error between the computational and tested results reduces considerably in this case also if separate saturation curves for teeth and core yoke are used instead of one ac saturation curve for both the teeth and core yoke. V. CONCLUSION The percentage error between the computational and tested results reduces, when separate saturation curves for teeth and core yoke derived from the normal dc saturation curve are used, instead of one ac saturation curve for both the teeth and core yoke" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000656_1.2953059-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000656_1.2953059-Figure4-1.png", + "caption": "Fig. 4. General diagram for simple pendulum", + "texts": [ + "1 Model of a simple pendulum It is about an unstable system described perfectly by a nonlinear model obtained by applying the laws of physics. The simple pendulum is a traditional example [4]. Observer Estimate state r cess Control Output Control signal Input Output Control Observer Adaptation of parameters Estimate st t Input Estimate parameter 1 2 2 1 Process Control Nonlinear part Linear part 86 The interest of design is to stabilize the pendulum in its position of unstable balance vertical. This system is treated like the model ( qq &, ), where q is the rod angle from the vertical position as shows in the figure 4. The Lagrangian is given by: ( )qcos.l.g.mE q.l.mE EEL P C PC \u2212= = \u2212= 1 2 1 22 & (3) what implies: ( )qcos.l.g.mq.l.mL \u2212\u2212= 1 2 1 22 & (4) The differential equations are: qsin.l.g.m q L q.l.m q L dt d q.l.m q L \u2212= \u2202 \u2202 = \u2202 \u2202 = \u2202 \u2202 && & & & 2 2 (5) According to the expression of Lagrange, the equation of the system will be expressed by: usinq.l.g.mq.l.m =+&&2 (6) Where u is the controller applied to the pendulum rod, m is the mass located at the end of the rod, g is the gravitational constant and l is the half length of the rod" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000360_4-431-27901-6_2-Figure2.9-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000360_4-431-27901-6_2-Figure2.9-1.png", + "caption": "Fig. 2.9a-c. Inclined pad and journal bearings (1). a inclined pad bearing, b journal bearing, c simplified journal bearing", + "texts": [ + " In these cases, since the surface is usually stretched, the generated pressure is negative and so the formation of the lubricant film is hard to explain hydrodynamically. Now, let us consider Eq. 2.20. It is written again here for the sake of convenience: \u2202 \u2202x ( h3 \u2202p \u2202x ) + \u2202 \u2202z ( h3 \u2202p \u2202z ) = 6\u00b5 [ (U1 \u2212 U2) \u2202h \u2202x + 2V2 ] (2.23) It is interesting to compare the meaning of the right-hand side of this equation in the case of an inclined pad bearing and in the case of a journal bearing. Consider the following conditions in an inclined pad bearing (V2 = 0) as shown in Fig. 2.9a. RHS stands for \u201cright-hand side.\u201d 1. If U1 = U and U2 = U, RHS = 6\u00b5 ( 0 + 0 ) = 0 2. If U1 = U and U2 = 0, RHS = 6\u00b5 ( U \u2202h \u2202x + 0 ) = 6\u00b5U \u2202h \u2202x 3. If U1 = U and U2 = \u2212U, RHS = 6\u00b5 ( 2U \u2202h \u2202x + 0 ) = 12\u00b5U \u2202h \u2202x In case 1, the relative velocity of the two surfaces is 0 and so no pressure is generated. If cases 2 and 3 are compared, the relative velocity in case 3 is twice that in case 2, and accordingly the generated pressure is twice that of case 2. Turning to the journal bearing shown in Fig. 2.9b, let us expand the cylindrical surface of the journal to a plane as shown in Fig. 2.9c. This is different from Fig. 2.9a in the important point that the circumferential velocity U j of the journal is along the slope, and this is equivalent to a surface having the velocity of U2 \u2248 U j in the x direction and that of V2 \u2248 U j(\u2202h/\u2202x) in the y direction. Therefore, the following relations are obtained: 1. If U1 = U and U j = U, RHS = 6\u00b5 ( 0 + 2U \u2202h \u2202x ) = 12\u00b5U \u2202h \u2202x 2. If U1 = 0 and U j = U, RHS = 6\u00b5 ( \u2212U \u2202h \u2202x + 2U \u2202h \u2202x ) = 6\u00b5U \u2202h \u2202x 3. If U1 = \u2212U and U j = U, RHS = 6\u00b5 ( \u22122U \u2202h \u2202x + 2U \u2202h \u2202x ) = 0 It is interesting to compare these results with those for the case of Fig. 2.9a. For example, if they are compared for case 2, the result, 6U(\u2202h/\u2202x), is the same, but the factors contributing to this result are different. For the inclined pad bearing, 6U(\u2202h/\u2202x) is directly obtained from the wedge effect. For the journal bearing, 6U(\u2202h/\u2202x) is interpreted as the sum of the wedge effect, which is negative, and the squeeze effect (the effect of V2), which is positive with a magnitude twice that of the wedge effect. Although it can be said from this analysis that the mechanisms of pressure generation in an inclined pad bearing and a journal bearing are different, this is not necessarily true. Figure 2.9a and 2.9c can be redrawn as Fig. 2.10a,ba and 2.10a,bb; considering case 2, it is clear that Fig. 2.10a,bb is equivalent to 2.10a,ba if viewed upside down. Therefore, the difference described above is only an apparent one. In the case of journal bearings, the circumferential speed of the journal U2 is along the inclined surface, and Eq. 2.20 becomes: \u2202 \u2202x ( h3 \u2202p \u2202x ) + \u2202 \u2202z ( h3 \u2202p \u2202z ) = 6\u00b5 [ (U1 + U2) \u2202h \u2202x + 2V2 ] (2.24) It is this form of the equation that appears in Reynolds\u2019 original paper [3]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000139_4-431-31381-8_17-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000139_4-431-31381-8_17-Figure6-1.png", + "caption": "Fig. 6. Time-to-foot information [17].", + "texts": [ + " Time to contact an obstacle or a target, which information can be obtained directly from a set of invariants present in the optical flow, has been studied as a key element in the visual control of locomotion [30]. We assumed that the step length modulation command, which was modelled in [a] previous study, was continuously related to optical information about the time remaining before one reached the target with the current eyefoot axis [17]. This optical variable in relation to the subject\u2019s own movement was labelled as time-to-foot (TTF) as shown in Fig. 6. We further assumed that the current step period was available and that it could be used with TTF to determine whether the step length must be shortened or lengthened to position the foot on the target. Results of computer simulation gave rise to successful pointing behavior as shown in Fig. 7. The generated behaviors for regulating step length were similar to those observed in human subjects performing a locomotor pointing task: namely, the time course of the inter-trial variability of the toe-target distance and the relationship between the step number at which the regulation is initiated and the total amount of adjustments involved" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002968_s00170-021-07401-y-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002968_s00170-021-07401-y-Figure3-1.png", + "caption": "Fig. 3 Application of the constricted arc a in the FPDL direction and b in the FADL direction", + "texts": [ + " L is the distance from the lower end of the electrode to the lower edge of the FCDAC, which is called the neck-in distance. h is the height from the lower edge of the FCDAC to the surface of the base metal, which is called the constricting height of the FCDAC. d is the width of the arc near the surface of the base metal, which is called the arc spreading width. Such constricted arc could be applied in two directions: the direction with the FCDAC perpendicular to the deposition layer (FPDL) and the direction with the FCDAC along the deposition layer (FADL), as shown in Fig. 3. Therefore, the shape of constricted arc was observed in two directions by the high speed camera. To ease the observation of arc, the welding torch stayed stationary and the workpiece moved. The observation direction of camera did not change in all the subsequent experiments. The real-time arc voltage and current signals were collected by a data acquisition system to characterize the heat input. To study constriction effects of arc by the FCDAC and its resultant shaping of depositions, the experiments in Table 2 were conducted" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001455_iros.2008.4650859-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001455_iros.2008.4650859-Figure3-1.png", + "caption": "Fig. 3. Four-link robot model", + "texts": [ + " xC, x\u0307C are the position and velocity of COM on the horizontal-axis. xd C, x\u0307d C are the desired of those. Superadding the GRF which follows this fxP2 to f\u0304d gr f in (7) result in \u03c4\u0303F . Therefore, the total joint torque is given as: \u03c42 = \u03c4\u0303F + \u03c4J (11) This is total joint torque of COM compensation in the horizontal-direction. This feedback control in (10) may interfere with the controller in (9), and tracking errors may increase. We argue this issue in next section and show results in Table III We apply our proposed method to a simulated four-link robot (see Fig. 3). Physical parameters of the robot model are described in Table I. Two kinds of squat motions with knee amplitude of 45 and 90 degrees were measured using motion capture system and force plates. Using this measurement system, joint angles, joint angle velocities, and the GRF were measured. We measured five squat cycles for each squat motion. Average squat frequency was about 0.5 Hz. Three of the five cycles are used as training data to determine the parameters of the linear models. The other two cycles are used as test data for the proposed imitation learning method", + " In this paper, we only consider the generation of motions in a two-dimensional sagittal plane. Since, in this study, we only consider squat motions, we assume that movements of the left and the right legs are the same and we used average joint angles and joint angle velocities of the left and right legs for simplicity. However, in general, we should treat the left and right legs independently. Latent continuous state variable X in (1) is defined as GRF fx in the horizontal and GRF fz in the vertical dimension, joint angles and joint angle velocities (see Fig. 3): X = (\u03b8ankle,\u03b8knee,\u03b8hip, \u03b8\u0307ankle, \u03b8\u0307knee, \u03b8\u0307hip, fx, fz)T , (12) We assume that only the joint angles and the joint angle velocities can be observed. Therefore, output matrix C is a 6\u00d7 8 as matrix. The element of this matrix is Ci j = \u03b4i j, where \u03b4i j is the Kronecker delta. We use two linear models (M = 2) to represent the two different squat behaviors with knee amplitudes of 45 degree and 90 degree. As discussed in Section II-A.3, the parameters of the linear models A(1,2),Q(1,2), and R(1,2) were determined from the training squat motions with knee amplitude of 45 and 90 degrees by using the least squares estimation as Fig. 4. In SSSM, we estimate X and S while parameter (\u03c0,\u03a6 and \u03bc (1,2) X1 and Q(1,2) 1 ) are optimized. Then, we derive joint torques using (8) based on the estimated joint angles, joint angle velocities, and GRF of the four-link robot model (see Fig. 3). The gains described in (9) are set as (Kp,Kd) = (100,70). The desired position xd C and velocity x\u0307d C in (10) are set to initial position and 0, respectively. The gains in (10) were set to (Kp2,Kd2) = (150,100). The computation time required to perform the learning of SSSM was 420 s for 5.6 s squat data by using a standard PC (Pentium4 at 2.80 GHz with 1 GB MEMORY). The algorithm converged for all 16 squat movements. Fig. 5 shows the estimated trajectories of the continuous latent variable X of (12): [(1)] GRF ( fx), [(2)] GRF ( fz), [(3)] joint angle, [(4)] joint angle velocity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001413_j.jsv.2008.02.056-Figure10-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001413_j.jsv.2008.02.056-Figure10-1.png", + "caption": "Fig. 10. Loading and meridional stress assumptions with axial restraint at R2.", + "texts": [ + " Our displacements in the coordinate directions 1, 2, 3, and 4 and in the axial direction for any particular ring element will be given by d1T \u00bc bQR1, (61a) d2T \u00bc bQR2, (61b) d3T \u00bc bQ\u00f0R2 R1\u00decos a \u2018 , (61c) d4T \u00bc bQ\u00f0R2 R1\u00decos a \u2018 , (61d) dVT \u00bc bQh, (61e) where b is the coefficient of thermal expansion, and where Q is the ring temperature rise. T.A. Smith / Journal of Sound and Vibration 318 (2008) 428\u2013460 447 Our displacements for the coordinate directions 1, 2, 3, and 4 and in the axial direction for a typical free ring element due to pressure loadings and known axial loadings will be obtained by using the principle of virtual work as used previously to obtain the ring element influence coefficients. The loadings and meridional stress diagrams associated with these loadings are shown in Fig. 10. Our stresses sx under these loadings are found by use of Eqs. (1), (8), and (10) to be sx \u00bc T02R Iz0 \u00f0y cos a z sin a\u00de HpR3 A . (62a) It may be seen from Fig. 10 that under the loadings L the stresses sy vary only slightly with y. We therefore assume them to be the average value at the centroid of the ring element cross section. In the case of the boundary loading Vp2, which is a triangular shaped loading, it is assumed, in accordance with the development and discussions relative to Eqs. (13)\u2013(16) hereinbefore, that the values of sy will be found closely on the basis of the value of R at the centroid of the ring element cross-section, with R assumed to be constant between the ARTICLE IN PRESS T" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001623_sice.2008.4654762-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001623_sice.2008.4654762-Figure4-1.png", + "caption": "Fig. 4. A free body diagram to distribute normal force algorithm: (a) forces on the plat surface, (b) heave force (c) rolling moment, and (d) pitch moment", + "texts": [ + " NrTrFTJ wmwxmww )(\u03bb\u00b5\u03c9 \u2212=\u2212= (6) Finally, friction estimation model for a tire can be obtained by (7). Nr JT w wwm \u03c9 \u03bb\u00b5 \u2212 =)( (7) where, mmm IKT = is a wheel driving torque that is obtained by input current to the driving motor. The difficulty to solve (4) or (7) is occurred to get normal force. For the commercial vehicle system, normal force estimation is other topic to conquered. The wheel-driven robot system of this paper has a 6-DOF sensor to monitor robot body states. Therefore, we suggest vehicle force distribution algorithm using this sensor and 6 wheeled body dynamics. Fig. 4 shows a free body diagram for our system. The main idea is that the forces acting on the tire\u2019s normal direction are occurred only by gravity, heave force, roll moment, and pitch moment, independently. So, while driving, the normal force is the sum of these forces altogether. If the robot body is on the plat surface, then only the normal force from gravity is acting on tire. If only heave force occurred, then, the normal force will be zMMg 6 1 6 1 + , the sum of gravity and have force in Fig. 4 (b). And if there\u2019s a positive direction roll moment in Fig. 4 (c), then force, r rollroll d J \u03b8 , are added to right side of robot body and subtracted to left side of robot body. The separated forces acting on left and right side are equivalently distributed 3 wheels on each side of which amount is r rollroll d J \u03b8 3 1 . In the same way, if there\u2019s a pitch moment, additional force, f pitchpitch l J \u03b8 2 1 , of pitch moment is added to front side of robot body and subtracted to rear side of robot body. Regarding the kinematics of robot body, mrzF _ and mlzF _ are attached to the center of robot body\u2019s longitudinal x -direction that is the same position of the 6-DOF sensor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003042_s11668-021-01191-x-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003042_s11668-021-01191-x-Figure1-1.png", + "caption": "Fig. 1 Basic components of EGR", + "texts": [ + " In petrol/gasoline and diesel internal combustion engines, a nitrogen oxide (NOx) emissions reduction method has been utilized called exhaust gas recirculation (EGR). The working principle of the EGR is to recirculate a portion of the exhaust gas back into the combustion chambers. This dilutes the O2 in the air stream, in turn reduces the mixing of nitrogen and oxygen that occur in the combustion cylinder and reduces the cylinder temperatures. The main components of the EGR system are shown in Fig. 1. The housing is the main component and removes the heat from the gas occupied tubes held by the header. The exhaust gases enter and exit through the diffuser, which also functions as the guide for smooth flow of gas at inlet and outlet of the cooler. EPDM seals are used to avoid the coolant side leakages. Winglet tubes create turbulence and provide more area for heat transfer. An EGR housing is subjected to high exhaust temperatures. But the material used in the EGR housing will withstand the high exhaust temperatures" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001571_14484846.2008.11464571-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001571_14484846.2008.11464571-Figure1-1.png", + "caption": "Figure 1: Dragline diagram, coordinate frames and strain gauge locations.", + "texts": [ + " Preliminary results from an investigation into reducing fatigue duty via improved slew torque control are also presented. The dynamics of the dragline bucket swing motion during house slewing (rotation) are of particular importance for both structural loading and effi cient operation. A simple, realistic slew torque profi le is fi rst assumed to give an indication of what could be achieved in practice. The model is then optimised allowing for more extreme variations in slew torque in order to demonstrate the contribution of out-of-plane bucket motion to duty. Figure 1 shows a diagram of the dragline model in the inertial reference frame XYZ. The origin O is at the intersection of the boom axis and the vertical slew rotation axis Y. The house and boom structure is modelled as a rigid body that rotates (slews) about the vertical axis with moment of inertia Ih. The bucket is modelled as a point of mass m suspended by the massless drag and hoist ropes d and h, which control the motion of the bucket within the bucket plane. The bucket position is defi ned by the four independent degrees of freedom: \u03c6, \u03b8, P and B" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003433_978-3-030-40667-7_6-Figure3.6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003433_978-3-030-40667-7_6-Figure3.6-1.png", + "caption": "Fig. 3.6 Different types of struts: (a) simple cardan strut joint; (b) modified cardan strut joint; (c) \u201cnear the top\u201d cardan strut joint", + "texts": [ + " The software, consequently, identifies the strut/platform connection point according to the type of support ring used or according to the input data. In hexapod systems where all kinds of supports can be used\u00a0 (SUV) with no need for a specific ring, struts can connect to the support platform everywhere, with no site restrictions. Consequently, in this scenario, it is necessary to input data regarding the distance of the strut connection point on the platform, both proximally and distally obtained by manual measurements. The strut is the mobile and adjustable element of the system (Fig.\u00a03.6). It can either consist of one threaded rod linked to mobile joints and adjusted by blocking nuts or made up of coaxial telescopic rods forming the real strut and cardan joints at each extremity. The strut length can be read on a graduated scale applied on the strut itself, otherwise, the strut length can be directly measured by a meter applied between the two strut connection points on the supporting platforms. Strut length parameters must be inserted in the software. Struts are always arranged counterclockwise for the support platform and numbered from 1 to 6 and marked with different colors" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002772_978-3-030-67750-3_14-Figure17.5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002772_978-3-030-67750-3_14-Figure17.5-1.png", + "caption": "Fig. 17.5 Ideal gearbox: steady-state model", + "texts": [ + " The vehicle is not accelerating, dv/dt is 0, and FTR is the maximum tractive force delivered by the motor at or near zero speed. With the assumptions, at near stall conditions, \u2211 F = 0 \u2192 FTR \u2212 FgxT = 0 \u2192 FTR = mg sin \u03b2 (17.8) If required mass (m) = 1300 kg, (full load) FTR = 1300 \u00d7 9.81 sin 18 = 3940N the maximum percent grade is max % grade = 100 tan \u03b2, max% grade = 100FTR \u221a (mg)2 \u2212 FTR2 = 100 \u00d7 3940 \u221a (1300 \u00d7 9.81)2 \u2212 (3940)2 = 32.48% TheEV transmission equation has established by assuming an ideal gearbox as shown in Fig. 17.5 with Plosses is 0, and the efficiency is 100%, perfectly rigid gears, and no gear backlash. For a tire wheel with radius r, the tangential and the angular velocity are related by: \u03c9r = v, \u03c9 = v r (17.9) The tangential velocity at the gear teeth contact point is the same for the two gears with different radius. rin\u03c9in = v = rout\u03c9out (17.10) The gear ratio has defined in terms of speed transformation between the input shaft and the output shaft. GR = \u03c9in \u03c9out = rout rin (17.11) Assuming 100% efficiency of the gear train: Pout = Pin,\u21d2 Tout\u03c9out = Tin\u03c9in (17" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003062_14644207211026696-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003062_14644207211026696-Figure1-1.png", + "caption": "Figure 1. Application of the circular fillet profile in the tooth root area.16 Source: reproduced with permission from Spitas et al., 2019.16", + "texts": [ + " In the test results for symmetrical and asymmetrical gears, it was observed that the increase in the pressure angle of asymmetric gears changed the way the tooth took damage. The equivalent working surface of symmetrical and asymmetrical gears used in the experimental study was designed with an involute profile curve. The root of the tooth was designed with the circular fillet profile. As suggested by Spitas et al.13,15,24,25 and stated by Kapelevich,12 to exhibit superior tooth root resistance with the circular fillet compared to the trochoid curve, obtaining the tooth root curve is shown in Figure 1. Gears tested were designated as a symmetrical spur gear with a 20 /20 pressure angle, an asymmetrical spur gear with a 20 /22 pressure angle, and an asymmetrical spur gear with a 20 /25 pressure angle. The designed symmetrical and asymmetrical gears are comparatively presented in Figure 2. As can be seen from Figure 2(b), as the pressure angle of the asymmetric gear increases, the tooth root area expands. In the experimental studies on single tooth bending fatigue test (STBFT) in the literature, test gears are determined as one of three or four adjacent teeth" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001889_2007-01-2368-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001889_2007-01-2368-Figure2-1.png", + "caption": "Fig. 2 Vehicle model: (a) full vehicle model; (b) normal ; (c) with faulty USM", + "texts": [ + " Assuming all the damped frequencies, di , are well separated in frequency, in comparison with 0 , the modal contributions in the CWT will be distinct and only the i th term in the summation in (13) will make a significant contribution. In this case for a fixed value of frequency, di , the logarithm of the CWT can be written as: (14) i ii d inidx XttW 02ln),(ln This shows that for a linear system the damping ratio, i , can be estimated from the slope of the line obtained by plotting the cross-section ),(ln idx tW against time. A full model of a vehicle system, with seven degrees-offreedom with USM effect, is shown in Fig.2(a). when considering the vertical displacement along the z axis, the equations of motions of the system are derived as: ).5.0.( 51111 ffsu tlzzczm (15) ).5.0.()( 511111 ffeu tlzzkuzk ).5.0.( 52222 rrsu tlzzczm (16) ).5.0.()( 522222 rreu tlzzkuzk ).5.0.( 53333 ffsu tlzzczm (17) ).5.0.()( 533333 ffeu tlzzkuzk ).5.0.( 54444 rrsu tlzzczm (18) ).5.0.()( 544444 rreu tlzzkuzk 4 1 5 ) 3 ( 4 1 55 [].)()1( .)()1([ i ei i INT i i isis zkiT iLzzczM (19) ].)()1(.)()1( ) 3 ( iTiLz i INT i i 4 1 5 )( 4 1 5 )1([)(", + " The roadway input motion )4,,1(iui is transferred to the unsprung mass uim through tire stiffness uik . I and I are the pitch and roll moments of inertia of the car body, respectively. Vehicle geometry parameters are frf tll ,, and rt where fl and rl are the distance from the front and rear axle to car body center of gravity and ft and rt are the front and rear wheel tracks. The vertical displacement of masses )4,,1(imui and sM are defined as iz and 5z , respectively. is the car body's pitch angle and is the car body's roll angle. In Fig. 2(a) we can see that the USM are supposed to be without mass and they have a stiffness of bk , and the cascade form is employed for the connection of the springs of the suspension system as the USM are modeled in ref. [12]. As mentioned in section 1, the bushing fault is modeled as a gap which in practice can appear as clearance in connections. This type of fault may be occurred due to wearing, rupturing or crushing. In this paper, the fault of bushing or damper is occurred in front-left part of VSS and the range of the gap, d , is assumed to be about 2 mm, Fig. 2(c), therefore in the case of faulty bushing, the 1bk parameter is defined as: If 025.0 15 bff kmmltz Else , mNkb /105 5 1 These conditions suggest that if the absolute displacement of the car body at the front-left corner is larger than 2mm, the car body and the USM will connect and no clearance would be between them. Otherwise, there would be no connection between the body and the suspension system in that section and the bushing and spring stiffness of the front-left part are neglected. As a result, the system equations would change to nonlinear equations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003132_978-981-16-0119-4_20-Figure19.11-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003132_978-981-16-0119-4_20-Figure19.11-1.png", + "caption": "Fig. 19.11 Frame drawing", + "texts": [], + "surrounding_texts": [ + "Amanufacturing team needs a drawing and specifications communicating the design (to the extent it impacts\u2014functional performance) (see Figs. 19.11 and 19.12).\nThe framework described above leverages technology to drive efficient design.", + "238 S. K. Mukherjee", + "It provides an example of proposing concepts that are a suitable option from strength and stiffness standpoint, while meeting form and functional requirements.\nStudents love to design products which consider human factors and are pleasant to use. The framework embodied in this example could be applied to other products and/or aggregates as an integral attempt to propose a design that is \u2018desired by customer\u2019 and which is appreciated by engineers and manufacturers. For those interested in designing bicycles for design competitions, this case study could be adapted to comply with the \u2018approval protocol\u2019 of UCI.8\nOptimization as a process has been used in industry for over a decade [14]. Virtual testing has been used in automotive industry to estimate \u2018fatigue life\u2019 [15]. The hypothesis that \u2018drawings are important in mechanical design process\u2019 has been evaluated, and support for the hypothesis has been reported [16].\nThere may be other case studies that already provide frameworks which leverage technology to incubate aspects of the \u2018applied sciences\u2019 in design of products. Considering the existing focus on inter-disciplinary studies in product design among researchers, this case study is an addition to these efforts.\n8UCI is an abbreviation of \u2018Union Cycliste Internationale\u2019\u2014Document: \u2018Approval Protocol for Frames and Forks\u2019. See: https://www.uci.org." + ] + }, + { + "image_filename": "designv11_83_0002801_s00006-021-01119-6-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002801_s00006-021-01119-6-Figure5-1.png", + "caption": "Figure 5. A representation of the Delta manipulator (adapted from [60] where B is the reference frame at the origin (centre of the base plate) and E is the reference frame of the end effector (centre of the end effector plate)", + "texts": [ + " Note that we have developed Eqs. (4)\u2013(14) only as a foundation for the difference maps and our objective is not to present a superior method to solve position problems which have been well researched and studied by experts in the field of robotics, such as methods that employ standard Denavit\u2013Hartenberg parameters [50], motor algebra G+ 3,0,1 [6,9] or projective geometric algebra [21]. As an example, we demonstrate how the quaternion rotor method can be used to solve the inverse position problem for a Delta manipulator. Figure 5 shows a 3-RUU Delta manipulator, which has three parallel limbs, five kinematic pairs per limb and three degrees of freedom, where the first joint of each limb is active and the other joints are passive. Each limb of the Delta manipulator can be deconstructed into five nodes, where each node defines a specific 1-DOF kinematic pair, resulting in a total of 15 nodes for the entire manipulator (Fig. 5). Table 1 presents the kinematic pair parameters, with reference to Figs. 5 and 6, where Hi is the initial vector of the kinematic pair from joint i to i + 1 (except for H5,H10 and H15, which are connected to the centre of the end effector platform) and si is the initial normalised unit vector along the joint axis. Additionally, |H0| = r1 and |H5| = |H10| = |H15| = r2 (with reference to Fig. 5). For parallel manipulators with M number of symmetrical limbs and b number of kinematic pairs per symmetrical limb, the following position element conversion rule applies for the symmetrical limbs: X l i j = Xi+(l\u22121)b j+(l\u22121)b, where l = 1, 2, . . . ,M . As such, for the sake of brevity and notional consistency, when referring to the Delta manipulator (M = 3), X l i j = Xi+(l\u22121)5 j+(l\u22121)5, such that X1 0 1 = X0 1, X2 1 2 = X6 7 and X3 2 3 = X12 13, for example. We construct a quaternion rotor inverse position equation for each lth limb of the parallel Delta manipulator (where l = 1, 2, 3), as shown in Eq", + " Since only the active joints (\u03b8l,1), participate in the computation of the inverse kinematics and dynamics of the manipulator, it is necessary to remove passive joints because they are unknown. As such, the computation of the inverse joint angles can be simplified by adapting the method presented in [53], such that the passive joint angles are removed. We square both sides of Eq. (15), giving the following equation for each limb:( X l 0 1 )2 + ( X l 1 2 )2 + (X l 3 4) 2 + ( X l 5 E )2 = (XE)2 (16) where \u03b4Hi = 0 for i = 1, 2, 3, 4, 5, for each limb, since none of the kinematic pairs in a Delta manipulator are prismatic joints (Eq. 7 and Fig. 5). Expanding Eq. (16) for limbs 1, 2 and 3, we get Eqs. (17a), (17b) and (17c), respectively. For the sake of brevity, we set XE = [0, x, y, z]. (2|X1 2||X0 1| \u2212 2|X1 2||X5 E | \u2212 2|X1 2|x) cos(\u03b81) + (2|X1 2|z) sin(\u03b81) +|X0 1|2 \u2212 2|X0 1||X5 E | \u2212 2|X0 1|x + |X5 E |2 +2|X5 E |x + x2 + y2 + z2 + |X1 2|2 \u2212 |X3 4|2 = 0, (17a) (2|X6 7||X0 7| \u2212 2|X6 7||X10 E | + |X6 7|x \u2212 \u221a 3|X6 7|y) cos(\u03b86) +(2|X6 7|z) sin(\u03b86) + |X0 6|2 \u2212 2|X0 6||X10 E | + |X0 6|x \u2212 \u221a 3|X0 6|y + |X10 E |2 \u2212 |X10 E |x + \u221a 3|X10 E |y +x2 + y2 + z2 + |X6 7|2 \u2212 |X8 9|2 = 0, (17b) (2|X11 12||X0 10| \u2212 2|X11 12||X15 E | + |X11 12|x + \u221a 3|X11 12|y) cos(\u03b811) +(2|X11 12|z) sin(\u03b811) + |X0 10|2 \u2212 2|X0 10||X15 E | + |X0 10|x + \u221a 3|X0 10|y + |X15 E |2 \u2212 |X15 E |x \u2212 \u221a 3|X15 E |y +x2 + y2 + z2 + |X11 12|2 \u2212 |X13 14|2 = 0, (17c) Eqs", + " Continuing with the Delta manipulator analysis presented in Sect. 2.1, we use the values for the joint axes (si) and link vectors (Hi) as shown in Table 1. In this example, we perform a P 1a {i}(\u0394Hi) transformation operation on kinematic pair i = 1, while fixing the end-effector position (XE), to produce the difference map P 1a {i}(\u0394Hi) (Eq. 34) for limb 1, as presented below P 1a {1}(\u0394H1) = UG1U1\u0394H1U \u22121 1 U\u22121 G1 (46) since \u03b4Hi = 0 for i = 1, 2, 3, 4, 5, none of the i-th kinematic pairs in a Delta manipulator are prismatic joints (Fig. 5). We perform a P 1b1 {i} (U\u03b4i) transformation operation on kinematic pair i = 1, while fixing the end-effector position (XE), to produce the difference map P 1b1 {i} (U\u03b4i) (Eq. 42) for limb 1, as presented below P 1b1 {1,1,n+1} (U\u03b41) = (R\u03b41 \u2212 1)x1 + 5\u2211 k=2 \u0394RGkxk = (R\u03b41 \u2212 1)(R1H1) + 5\u2211 k=2 (R \u2032 GkR\u22121 Gk \u2212 1)(RGkRkHk) = R\u03b41R1H1 + R \u2032 G3R3H3 + R \u2032 G5R5H5 \u2212R1H1 \u2212 RG3R3H3 \u2212 RG5R5H5. (47) Since H2 and H4 are zero vectors for the Delta manipulator where R \u2032 G3 = R\u03b41R1R2, R \u2032 G5 = R\u03b41R1R2R3R4, RG3 = R1R2, RG5 = R1R2R3R4 and \u03b4Hi = 0", + " Then, we presented the Transformation Cases and the objective indices for Transformation Cases 1 and 2, which allows us to find the transformation variables required to achieve the objectives of the Transformation Cases 1 and 2 (Eqs. (61)\u2013(63)). The Lagrangianbased characterisation of robot dynamics, as presented in this work, is applicable to serial, parallel, hybrid serial-parallel and non-standard robot architectures and morphologies. The Transformed Lagrangian and the Difference Lagrangian Functions are presented in the case study section (Sects. 3.2\u20133.3). Continuing with the Delta manipulator analysis (shown in Fig. 5), if we select p number of possible type transformation variables per kinematic pair (and set for example p = 5) and k number of reassembling transformations (and set for example k = 3) to be performed for n number of kinematic pairs (where n = 15 for the Delta manipulator), we have a total of 67525 (since pnCk = pn! k!(pn\u2212k)! = (5\u221715)! 3!(5\u221715\u22123)! = 67525) possible combinations of reassembling transformations. This prohibitively large number of combinations produces a computationally challenging and unfeasible number of potential permutations in the number of generic type transformations potentially applicable to each kinematic pair", + " (109i) We present expressions for the quaternion rotor based Lagrangian dynamics of a Delta manipulator, where the expressions for the inverse position solutions to find \u03b8l,1 are given in Eq. (22), while the expression for the inverse kinematics solutions, \u03b8\u0307l,1 and \u03b8\u0308l,1 are given in Eqs. (27a) and (27b), respectively. In this example, we set the six generalised coordinates to be c1, c2, . . . , c6 = x, y, z, \u03b81,1, \u03b82,1, \u03b83,1, where x, y, z are redundant coordinates [50]. Since the distance between kinematic pairs 3 and 4 ((X3 4) 2) is always equal to the length of the connecting rod of the upper arm a2 = |H3| = |X3 4| (based on Fig. 5), we define the constraint function \u0393l as follows [50] \u0393l = ( X l 3 4 )2 \u2212 (|X l 3 4| )2 = ( x + |X l 5 E |cos (\u03c8l) \u2212 |X l 0 1|cos (\u03c8l) \u2212 |X l 1 2|cos (\u03c8l)cos (\u03b8l,1) )2 + ( y + |X l 5 E |sin (\u03c8l) \u2212 |X l 1 2|sin (\u03c8l)cos (\u03b8l,1) )2 + ( z \u2212 |X l 1 2| sin (\u03b8l,1) )2 \u2212 (|X l 3 4| )2 = 0 (110) where l = 1, 2, 3 and the symbol \u03c8l denotes the angles between the vector from the origin to the first joint of each l-th limb and the x-axis of the base reference frame (such that \u03c81, \u03c82, \u03c83 = 0, 2\u03c0 3 ,\u2212 2\u03c0 3 ). To simplify the analysis, we assume that the mass of each connecting rod in the forearm (m2) is divided evenly and concentrated at the two endpoints (X0 3 and X0 4)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001005_apex.2007.357700-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001005_apex.2007.357700-Figure5-1.png", + "caption": "Fig. 5. Experimental test setup.", + "texts": [ + " The high frequency voltage and current were measured and the reactances were found using (14) and (15). In the simulation, the machine is accelerated to speed of 30 rpm without load. The position estimation by computing high frequency reactances is show in Fig. 4. Note that the reactance calculation yields the rotor position with relatively good accuracy. A. System configuration The sensorless control system was evaluated in the laboratory using a commercial PMSM drive system. The PMSM data are listed in the Table I. Fig. 5 shows the configuration of the test platform. As shown in Fig. 5 the test setup is composed by a microcomputer equipped with a dedicated data acquisition and control board and two PMSM servodrives. The command board of one of the servodrives (CONVERTER 1) has been disabled and the control signals have been provided by microcomputer board. The command signals are generated by the microcomputer with a sampling time of 10 ,us. The data acquisition channel employs Hall effect sensors and 12 bit A/D converters. CONVERTER 2 is used to simulate different load conditions for testing the proposed solution" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000882_s1061920808040109-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000882_s1061920808040109-Figure2-1.png", + "caption": "Fig. 2. A quadrangle with configuration space two disjoint circles.", + "texts": [], + "surrounding_texts": [ + "In this section, we look more closely at two specific quadrangles, both of which are generic. 3.1. The data specifying the first example (i.e., the lengths of the four bars, including the immobile one) are l0 = 1, l1 = 0.3, l2 = 0.5, l3 = 0.3. Figure 1 below shows three (admissible) positions of the linkage: ABCD, AB\u2032C \u2032D, ABmaxC1D. Note that, in the position ABmaxC1D, the bars BmaxC1 and C1D are aligned. The \u201cfigure eight\u201d curve C in the middle of the picture is the geometric locus of the midpoint M of the mobile bar BC, while M1,M2,M3,M4 are its locations corresponding to the positions at which two bars are collinear. As the linkage moves, the point M sweeps out the \u201cfigure eight\u201d curve C with selfcrossing point at the origin O. To each point M on the curve C there corresponds exactly one position of the linkage, except for M = O, for which there are two positions (the reader is invited to draw them on the figure). Thus C is not quite the configuration space of the linkage under consideration, and topologically the configuration space is the circle. To see this, let us describe the canonical configuration space (that we denote by C0) of our quadrangle. First we introduce the coordinates of the mobile hinges B = (x1, y1), C = (x2, y2). These coordinates satisfy the system of equations \u23a7 \u23aa\u23a8 \u23aa\u23a9 x2 1 + (y1 + 1)2 = (0.3)2, x2 2 + (y2 \u2212 1)2 = (0.3)2, (x1 \u2212 x2)2 + (y1 \u2212 y2)2 = (0.5)2. (1) Elementary but tedious calculations show that the algebraic curve C0 \u2282 R 4 is in fact nonsingular and connected (i.e., is a smooth simple closed curve). RUSSIAN JOURNAL OF MATHEMATICAL PHYSICS Vol. 15 No. 4 2008 The projection from R 4 to R 2 given by the formula (x1, y1, x2, y2) \u2192 ( x1 + x2 2 , y1 + y2 2 ) (2) takes C0, the canonical configuration space of the quadrangle under consideration, to the \u201cfigure eight\u201d curve C shown in Fig. 1. 3.2. It should be noted that, in this case (as in other generic cases), although certain mobile points of the system describe semi-algebraic sets (e.g., the point B moves back and forth along the closed circular arc BminBmax), the configuration space of the system is not only algebraic, but smooth. Therefore there are no problems in defining the phase space of the system (in our case, it is the line bundle over the circle) and the velocity field of one of its possible motions (defined as a smooth section of this bundle). In the present paper, we will be mostly interested in motion without stopping, so that we will be dealing mostly with nonzero sections. However, even in our simple (generic) example, there exist regimes of motion requiring a more sophisticated description of the velocity field; this is discussed in the next subsection. 3.3. It is important to note that the smooth structure on the configuration curve is not sufficient to adequately describe the kinematics of our linkage: the curve actually possesses four distinguished points, which we denote by M1,M2,M3,M4 and call collineation points or switching points); they are indicated by stars on the configuration curve C in Fig. 1. At these points, a rebounding phenomenon in the kinematics of the linkage may occur. When the linkage moves so that M reaches one of these points, say M1, two of the bars constituting the linkage (in this case, DC1 and C1Bmax) become collinear and the regime of motion can switch (continue smoothly without stopping in two radically different ways): either so that M continues moving in the same direction along C, or so that it rebounds from the position M1 and starts moving along C in the opposite direction. The way the system will move after it reaches a collineation point depends on its dynamics: heavy bars with considerable inertia will tend to continue moving in the same direction, whereas certain technological constructions of the hinges have a certain \u201cbounce\u201d in them that makes the bars rebound as they straighten out, much like a billiard ball bounces off the boundary of the billiard table. Note that when the direction of motion of the point M along the curve C is instantly reversed, it is natural (but wrong!) to say that the velocity vector has a jump discontinuity. To formalize the RUSSIAN JOURNAL OF MATHEMATICAL PHYSICS Vol. 15 No. 4 2008 notion of smoothness in this situation, we need an appropriate definition of the phase space at the collineation points. We will come back to this question in Section 4. 3.4. The second generic quadrangle ABCD that we consider here is given by the following data \u30081; 5/8, 7/8, 5/4\u3009. The configuration space of the system consists of two closed disjoint algebraic curves. The two intersecting ovals in the middle of the picture constitute the image under the projection from R 4 to R 2 given by (2) of the canonical configuration space C0 of the quadrangle ABCD. The preimages of these two ovals do not intersect, and it is impossible for the linkage to pass continuously from a position on one of the ovals to one on the other. This quadrangle can be regarded as a simplified model of the ordinary automobile windshield wiper. An electric motor uniformly rotates the bar AB about the fixed hinge A. The bar BC then moves so that its extremity C runs back and forth between the points C1 and C2 along the arc of the circle of radius l1 = 5/8 centered at A, and so the brush FE on the extension DE of the third bar CD sweeps the windshield, moving back and forth from the limit positions F1E1 and F2E2. For other parameter values, the projection of the canonical configuration space to the (x, y)plane under the map (2) can consist of nonintersecting ovals (for an example, see Fig. 3(b)); our specific choice of parameters producing two intersecting curves in the figure was motivated by the intention of obtaining a motion of the brush FE looking like that of a real windshield wiper. 3.5. Just as in the case of the quadrangle in Subsection 3.1, the definitions of the phase space and of velocity fields of the second quadrangle are straightforward: they are, respectively, the line RUSSIAN JOURNAL OF MATHEMATICAL PHYSICS Vol. 15 No. 4 2008 bundle over the disjoint union of two circles and its smooth (nonzero) sections. This definition is adequate provided we stipulate that no collineation points are rebound points. What one must do if this is not the case is explained in the next subsection." + ] + }, + { + "image_filename": "designv11_83_0003385_ur52253.2021.9494662-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003385_ur52253.2021.9494662-Figure3-1.png", + "caption": "Fig. 3. Mecanum In-wheel motor", + "texts": [ + " Therefore, a cycloid gear-based mecanum in-wheel motor equipped with a reducer was developed for improved space usability within a mobile platform with low cost. The mecanum in-wheel motor was developed by redesigning the 3D printed cycloid reducer used in the manipulator. A 6-inch mecanum roller was mounted on the outer ring of the output terminal, and the gear ratio was reduced to 14:1 for highspeed running. The motor was developed in a module type to maximize the space usability within the mobile platform, creating a high-speed reaching up to 20 km/h. The internal structure of the mecanum in-wheel motor is shown in Fig. 3. A suspension module was designed for perfect contact between the mecanum wheel and the ground. The suspension module operates independently for each wheel. The equation below was designed so that the variable length is compressed by 1/3 and the maximum payload is 20 kg under the noload condition, considering the spring mass of the mecanum mobile platform. The spring was selected using (1). Fig. 3 shows the internal structure of the suspension module; the parameter values of the equation are summarized in Table III. Fs = Mg Ns \u2206x = Fs N (1) A mobile platform was developed using the mecanum in-wheel motor and the suspension module. It has the advantages of holonomic system owing to an omni-directional movement. The body consists of aluminum profiles in which the mobile platform can be run at a high speed of up to 20 km/h and has a maximum payload of 20 kg. Fig. 4 shows the mecanum mobile platform; the detailed specifications are summarized in Table IV" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001767_09544062jmes1076-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001767_09544062jmes1076-Figure4-1.png", + "caption": "Fig. 4 (a) Schematic of a 60\u25e6 punch model and (b) slip response map for a 60\u25e6 punch", + "texts": [ + " Further changes of the applied load, beyond the \u2018adhesion trough\u2019 (Fig. 3b), imply growth of the internal slip zone. This has not been investigated in detail, but it is expected that the growth in slip will progress towards both contact edges until sliding eventually occurs. Note that the presence of any bulk tension has no effect on the condition for incipient sliding. The contact geometry used to investigate behaviour when the edges make an interior angle of 60\u25e6 to the contact edge is shown in Fig. 4(a), and the \u2018substrate\u2019 is, again, a strip of thickness a. The total included angle at the edge of the contact is now below the figure at which the antisymmetric dominant eigensolution becomes bounded, and the symmetric solution has an eigenvalue of 0.6157, so that the order of singularity expected is rather less than in the previous example, but here the critical coefficient of friction (= g I r\u03b8 ) is fo = 0.962. The calibrations for the generalized stress intensity factors are as follows { KI a1\u2212\u03bbI KII a1\u2212\u03bbII } = [\u22120", + " IMechE Vol. 222 Part C: J. Mechanical Engineering Science JMES1076 \u00a9 IMechE 2008 at TU Muenchen on July 11, 2015pic.sagepub.comDownloaded from JMES1076 \u00a9 IMechE 2008 Proc. IMechE Vol. 222 Part C: J. Mechanical Engineering Science at TU Muenchen on July 11, 2015pic.sagepub.comDownloaded from It follows that, in the vast majority of practical problems, the coefficient of friction actually present is in the slipping regime, and the response of the contact is rather compartmentalized, as shown in Fig. 4(b). Thus, on pressing down the contact normally, fringes of edge slip will be present if the coefficient of friction does exceed fo, and edge separation, by hitting the \u2018walls\u2019 will ensue if the \u03c3o, Q history takes the loading trajectory away from the central region. In practice, there will always be at least some slip. The last geometry studied was the 90\u25e6 contact shown in Fig. 5(a), but here the layer is deeper, with a depth of 5a. As before, the initial attack on the problem assumed complete adhesion, and the contacting pair was modelled as a monolith within the finite element simulation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000348_intmag.2006.375779-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000348_intmag.2006.375779-Figure1-1.png", + "caption": "Fig 1 Machine cross section", + "texts": [ + " The machines used usually have axial rotor laminations; this is a modification from the synchronous reluctance machine. However, in the doubly-fed application the rotor experiences rotating flux waves which can produce lamination eddy currents which could cause excess losses. This paper seeks to investigate this and compare to a radial laminated rotor using a 2/6 pole arrangement, (with 2 pole power and 6-pole control windings). Modelling The machine is modelled using Cedrat FLUX 2D. The rotor laminations are modelled individually which produces a very dense mesh (Fig 1). The laminations are stacked and form a \u201cU\u201d shape along the axial length. There is spacing between the laminations to produce high q axis reluctance. In this machine the stacking factor is 0.5. The other modelling alternative is to use anisotropic material [5]. The machine was simulated using a series of static solutions then a series of transient solutions. Finally the machine was compared to an equivalent radially-laminated machine. Experiments A 2/6 pole test machine was set up; it was rewound from an axially-laminated synchronous reluctance motor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000901_icmech.2007.4280030-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000901_icmech.2007.4280030-Figure5-1.png", + "caption": "Figure 5 3 DOF Haptic Device", + "texts": [ + " On the other hand, an imitated endoscope is our original device which has one degree of freedom for translation and three degrees of freedom for rotation, a roll angle, a pitch angle and a yaw angle. A translational displacement and rotational angles are measured with potentiometers. We show function and specifications of imitated operational tools in Table 1 and an imitated endoscope in Table 2. tigure 4 Imitated Endoscope C Haptic device Each imitated operational tool is linked with the haptic device through holes pierced on the virtual human body. The haptic device adopts parallel link system with three degree of freedom to obtain the high accuracy and high rigidity\". As shown in Figure 5, a haptic device is consisted of a base plate, three DC servomotors with reduction gear which are connected to the upper plate with bending conformer link. A joint of two degree of freedom is on the upper plate and linked with an imitated operational tool. The upper plate always keeps level by restriction of a linkage and does coordinate detection of three translation degrees of freedom and generates of forces. We show functions and specifications of the haptic device in Table 3. It is possible to calculate the direction of the imitated operational tool from the position of upper plate and the hole pierced on the virtual human skin" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001107_00207160701477476-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001107_00207160701477476-Figure2-1.png", + "caption": "Figure 2. The surface D \u2217 with the un-parametrized region of \u03c4tr shown in black in an exaggerated way.", + "texts": [ + " The homogeneous form is used only to reduce the width of the parametric expressions. The R 3 parametrization may be derived, by division, from the homogeneous form in the usual way as: 1 a \u2212 c cos \u03b8 cos \u03c6 (\u03bc(c \u2212 a cos \u03b8 cos \u03c6) + b2 cos \u03b8, b(a \u2212 \u03bc cos \u03c6) sin \u03b8, b(c cos \u03b8 \u2212 \u03bc) sin \u03c6). Figure 1 shows D in the case of a = 6, b = 4 \u221a 2, (c = 2), and \u03bc = 3 (left); and the origins and directions of the angular parameters, \u03b8 and \u03c6, on the manifold (right). We denote the region parametrized by \u03c4u by D \u2217. Figure 2 shows the surface D \u2217 with the un-parametrized region, which is of area-measure zero, highly exaggerated. In this paper, we use a normalized form \u03c4tr : (0, 1)2 \u2192 D \u2217, which may be expressed as \u03c4tr (t, s) = \u23a1 \u23a2\u23a2\u23a3 \u03bc(c \u2212 a cos(2\u03c0t) cos(2\u03c0s)) + b2 cos(2\u03c0t) b(a \u2212 \u03bc cos(2\u03c0s)) sin(2\u03c0t) b(c cos(2\u03c0t) \u2212 \u03bc) sin(2\u03c0s) a \u2212 c cos(2\u03c0t) cos(2\u03c0s) \u23a4 \u23a5\u23a5\u23a6, where 0 < t, s < 1 are defined by t = (\u03b8/2\u03c0) and s = (\u03c6/2\u03c0). The parametrization \u03c4tr is maximal in the sense that the topology of D does not permit parametrizations, rational or transcendental, of greater range" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001655_pes.2007.385932-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001655_pes.2007.385932-Figure7-1.png", + "caption": "Fig. 7. Blackout material (Source: Tokyo Electric Power Company)", + "texts": [], + "surrounding_texts": [ + "Large blackouts are a common concern for the average person, including students. There was a large-scale blackout in New York in 2003, and also a blackout in Tokyo in 2006. Students are interested in knowing why these blackouts occurred, how restoration was done, and why it took such a long time to get the power back. It is a kind of mysterious phenomenon for students. We must obtain correct information about the blackouts from the power utility people. We need good cooperation with the industry for our lectures. For example, for the blackout which occurred in Tokyo in August, 2006, the author obtained power point materials which a TEPCO engineer used for presentations at IEEE, with permission granted for use in his classes. When the author explained the details of the blackout to students, the author could see that they were so excited and were interested to see the real practical application of what they learned in the lecture. To attract students to power areas, this kind of cooperation with power utilities is indispensable." + ] + }, + { + "image_filename": "designv11_83_0002732_j.engfailanal.2021.105451-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002732_j.engfailanal.2021.105451-Figure8-1.png", + "caption": "Fig. 8. Potential and kinetic energy difference, the first oscillating mode, undercarriage, slewing platform and pylons of BWE SchRs 630, spatial view and lateral plane.", + "texts": [], + "surrounding_texts": [ + "When it was determined that a model (and, in principle, laboratory) testing is possible and that the obtained results are applicable to the actual structure, it was decided to manufacture a sub-scaled (laboratory) model. Complete technical documentation and technology for model development have been developed. The material used for this structure was steel S355J2 + N. For practical reasons, the unification of thicknesses was performed. Plates whose thicknesses deviate by up to a maximum of 20% from the thicknesses obtained by a tenfold reduction of dimensions were used. The undercarriage and slewing platform with the lower part of the pylons in different phases of construction are shown in Fig. 11 and Fig. 12. The manufactured model has a weight of approximately 230 kg and a frame size of 1550x1650x950 mm. In order to facilitate the load input in the cut-off zone of the columns, a ribbed plate was introduced, which connects the structures of the cut-off pylons and enables further better load input. A. Petrovic\u0301 et al. Engineering Failure Analysis 126 (2021) 105451 A. Petrovic\u0301 et al. Engineering Failure Analysis 126 (2021) 105451 In the calculation of the laboratory model, a load correction in relation to sub-scaled structure was made. The load is reduced to forces in the axial directions (two horizontal in the plane of intersection in the X and Y axes and vertical in the Z direction). The results of the numerical analysis are given in Fig. 13. Comparing the corresponding places on the sub-scaled model and on the laboratory model, at first glance, the absolute coincidence of the stress field of these two models is noticeable, except that the problem of large stresses at the intersection line of pylons is solved by adding a plate that connects the pylons. As far as the displacement of the bearing points is concerned, the sub-scaled structure and A. Petrovic\u0301 et al. Engineering Failure Analysis 126 (2021) 105451 the laboratory model have almost the same displacements (see Figs. 14 and 15). A dynamic calculation of the laboratory model was performed. The results are similar to the results obtained on sub-scaled model. In our laboratory conditions, it was not possible to induce excitation of such a high frequency, so dynamic tests were abandoned. The dynamic calculation was used to once again identify the weak points of the structure. In limited laboratory conditions, it was not possible to cause multi-axis loading of the model, so it was decided to load the model with equal vertical forces (symmetrically). It was decided that the model is loaded only with vertical forces of 10 kN per pylon at the points of intersection of the vertical plates of pylons. Given the linearity of finite element calculations, this does not affect the research. The model is positioned in a rigid frame, which served to load the model. The vertical forces are caused by the hydraulic cylinders resting on the upper horizontal beam. Each of the performed measurements involved a gradual input of force using a hydraulic cylinder at two points of the model (symmetrically), with a step of 2 kN per individual input point, i.e. in a pair of 4 kN in total. The aim was to verify the numerical calculation of the laboratory model by experimenting on a physical (previously developed) laboratory model. By examining the physical model, it was finally confirmed that such structures can realistically be examined by model, while the numerical calculation only indicated that such structures, in theory, can also be examined by model. The experimental methods used are strain gauges method and the method for non-contact measurement of stresses and strains. The symmetrical load enabled (in addition to the already existing symmetry of the structure itself) parallel measurement with strain gauges and an optical system for digital image correlation. Testing of the construction by the method of strain gauges, implies positioning of strain gauges at previously selected critical places, with the aim of mapping the model. As a problem with the method of strain gauges, which is most often used for diagnostics, the positioning of gauge itself is imposed. The gauge should be placed in zones where there is a large A. Petrovic\u0301 et al. Engineering Failure Analysis 126 (2021) 105451 stress change gradient, so just locating the strain gauge has a big impact on the results. This problem is overcome by using the method of digital image correlation, which enables the recording of stress over the entire field. The system for non-contact optical measurement of displacement and deformation consists of sets of stereo cameras and lenses. The system also consists of a stand, a power supply and a PC system with Aramis software installed. The operation of the system is based on digital image correlation, which compares the position of points before and after deformation. The stochastic pattern, applied to the surface of the object to be deformed (irregular black dots on a white background), allows the system to identify the points (pixels) that it tracks during deformation. This system can measure displacements up to 0.001 mm. Thus, method of strain gauges and method for non-contact measurement of displacement and deformation would confirm each other, and the overall information obtained is more complete than if only one method was used. The readings of strain gauges during one measurement are shown in Fig. 16. From the readings of strain gauges, it can be seen that the T5, T6, T7 and T8 results, although they follow the load of the model, are unusable, because the values of the measured elongations are within the measurement error. This was expected based on the results of the calculation, because due to the limitations of the experiment, a reduced vertical load was adopted, which is not of sufficient intensity to revive all the gauges. Strain gauge T5 is positioned on upper horizontal plate of undercarriage, where vertical plate connects the horizontal plate. Strain gauge T8 is positioned on the upper horizontal plate of slewing platform where vertical plate connects the horizontal plate. Rosette (T1-T3) is positioned on the vertical plate of undercarriage, near the outer cylinder, in the zone of stress concentration. Fig. 17 shows the position of the rosette. The results obtained using strain gauges T1-T3 are compared with results obtained using system for optical non-contact measurement, and this group of the results is named MM 1\u20133. The position of the recording equipment during measurement conducted in spot MM 1\u20133 is shown in Fig. 18. A line in stress field obtained using DIC that imitates a strain gauge is shown in Fig. 19. Namely, the software provides the ability to determine the distance between any two points at any time. Thus, line is drawn along in the vertical direction in the rosette zone, which practically imitates the T2 strain gauge. According to the table, a high matching between experimental results and numerical calculation can be noticed. It can also be noticed that everything that can be measured with strain gauges, can also be measured with a system for non-contact measurement of A. Petrovic\u0301 et al. Engineering Failure Analysis 126 (2021) 105451 displacement and deformation, faster and more efficiently. Strain gauges T6, T7 and T9 are positioned as shown in Fig. 20. The measurement results at MM 6 are given below. In Fig. 21, a field of vertical displacements that is fully aligned with the field obtained by the numerical calculation can be seen. This is the point around which the pylons bend slewing platform. The measurement results at MM 9 are given below. At this point, a considerable difference between the numerical calculation and the experimental results was observed. The reason for this are the massive welded joints near this place, because the cylinder, the upper plate of slewing platform and the knot plate of sprit join there. The welded joint, on the one hand, increases the stiffness of this place, and on the other hand, it can cause the A. Petrovic\u0301 et al. Engineering Failure Analysis 126 (2021) 105451 A. Petrovic\u0301 et al. Engineering Failure Analysis 126 (2021) 105451 concentration of stress entered by welding. The results of the experiment confirm the numerical calculation of the laboratory model. By confirming that the laboratory computational model is good, all computational models that preceded it were indirectly confirmed." + ] + }, + { + "image_filename": "designv11_83_0003169_s00202-021-01352-z-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003169_s00202-021-01352-z-Figure1-1.png", + "caption": "Fig. 1 The proposed STPMIG: a explosive structure, b schematic of the generator", + "texts": [ + " 2 Structure and\u00a0working principle of\u00a0STPMIG In order to investigate the performance of the proposed double-rotor generator adopted for double-turbine wind applications, the structure and working principle of the generator must be investigated in this section. Hence, the configuration and structure of inner and outer rotors and the connection method of both rotors to auxiliary and main turbines are studied. The dynamic model of the novel STPMIG is also surveyed by considering the operation principle of the proposed direct-drive generator in double-turbine systems. Figure\u00a01 illustrates the explosive and 2D structures of the proposed generator. As shown, the STPMIG has two separate rotors, which can be directly connected to a double-turbine wind system. The outer rotor adopts permanent magnets and supplies the reactive power and the flux at both air gaps of the generator. The inner rotor utilizes the squirrel-cage 1 3 structure in order to minimize maintenance costs by eliminating slip rings and brushes. Also, in order to directly connect the proposed generator to a wind turbine system, it is necessary to select the number of generator poles as high as possible" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002240_012003-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002240_012003-Figure1-1.png", + "caption": "Figure 1. Geometry of the problem.", + "texts": [ + "1088/1742-6596/1730/1/012003 ( ) ( ) 2 2 2 1 1 1 1 1: 2 2 N N j i jj ij ijh hi j ij j ij ij j iji j ji i j i j ij mpp m W W r = = + \u2212 = \u2212 + \u2212 r u \u03c3 D u u (9) ( )( ) 2 1 1 N i j i jj h i ij j ij ji i j ij T Tm W = + \u2212 \u2212 = q r r (10) where ui \u2013 particle velocity, uij \u2013 velocity of particle i relative to particle j, \u03c3 \u2013 stress tensor, Fsn \u2013 surface tension force, Fs\u03c4 \u2013 Marangoni force, Fv \u2013 vapor pressure force, g \u2013 acceleration of gravity, Pi \u2013 vapor pressure per particle i, \u03bc \u2013 dynamic viscosity, rij \u2013 position of particle i relative to particle j, \u03b1 \u2013 model coefficient of surface tension force, \u03b2 \u2013 model coefficient of Marangoni force, ni \u2013 normal to the surface at the position of the particle i, Ci \u2013 color function, \u2207Ci \u2013 color function gradient, Vi \u2013 particle volume [1/\u043c3], Ti \u2013 particle temperature. A numerical experiment was carried out to simulate the process of wire surfacing using the developed model. The purpose of the simulation was to test the performance and adequacy of the constructed model. The problem of surfacing a wire made of steel on a substrate made of the same material in the jet transfer mode at a low power of the heat source is considered. The geometry of the problem is shown in Fig. 1. Table 1 shows the physical characteristics of steel grade 12X18H10T, required for numerical modeling by the smoothed particle method. IC-MSQUARE 2020 Journal of Physics: Conference Series 1730 (2021) 012003 IOP Publishing doi:10.1088/1742-6596/1730/1/012003 The numerical simulation parameters are as follows: Preliminary numerical implementations were carried out on an IBM 2x300 sas 15k multiprocessor computer (4xIntel Xeon E7520, 64 GB) using MPI multithreaded computing capabilities in the LAMMPS package" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002783_09544062211012724-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002783_09544062211012724-Figure6-1.png", + "caption": "Figure 6. Singular case where the length of one limb is 0.", + "texts": [ + " In Figure 5, modified sigmoid function, arctangent function and one of the algebraic functions are respectively f\u00f0x\u00de \u00bc 2 1\u00fe e x 1 (30) f\u00f0x\u00de \u00bc arctan\u00f0x\u00de (31) and f\u00f0x\u00de \u00bc xffiffiffiffiffiffiffiffiffiffiffiffiffi 1\u00fe x2 p (32) It can been observed in Figure 5 that the tanh\u00f0x\u00de function is a shifted and narrowed sigmoid function and tanh\u00f0x\u00de is scaled in [\u20131]. The tanh\u00f0x\u00de is closer to the Heaviside function compared to the others. Singularity analysis with loads A PM loses degree of freedoms at its singular configuration. For the prismatrically actuated 3-RPR mechanism, it has three singular types, as shown in Figures 6 to 9. The singularity depicted in Figure 6 occurs in the limit situation where at least one limb of the 3-RPR PM owns zero limb length. This theoretical situation is easy to avoid in practice. Then the effect of employed load on this situation is not further discussed here. The planar 3-RPR manipulator lose its rotational DoF when three axes of limbs intersect at the common point O1, as shown in Figure 7. Two necessary conditions need to be met to have this kind of singularity when employed loads are considered, that is, 1. Axes of the three limbs can intersect at the common point geometrically within the workspace of the manipulator" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003347_s40430-021-03120-3-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003347_s40430-021-03120-3-Figure1-1.png", + "caption": "Fig. 1 Structure of electric hydraulic pump", + "texts": [ + " The main research contents were as follows: (1) with the electric hydraulic pump in the new electrohydraulic power steering system as the research object, the source of the noise was traced Journal of the Brazilian Society of Mechanical Sciences and Engineering (2021) 43:407 1 3 Page 3 of 11 407 through noise testing and analysis; (2) the mechanism of noise generation was analyzed to provide theoretical solutions to reduce the noise; and (3) the feasibility and effectiveness of noise optimization of the electric hydraulic pump by blocking vibration transmission and suppressing surface vibration were also explored, which was verified by the vibration and noise test of the modified prototype of electric hydraulic pump. The research object of the present study was the new generation of electric hydraulic pump, which involves the high integration of oil pump, motor, and controller. Such integration can ensure the reliability of the whole electrohydraulic power steering system. The structure of the electric hydraulic pump is shown in Fig.\u00a01. To shorten the axial size of the assembly, the oil pump was directly connected with the motor rotor and embedded in the motor. The controller was integrated at the other end of the motor and was replaceable. Regarding the oil pump type, in order to meet the hydraulic requirements of steering gear of different new energy commercial vehicle types, the oil pump parameters were determined, as shown in Table\u00a01. For the motor type, the permanent magnet brushless single winding motor was adopted. A Hall element was used to detect the position of the motor rotor and it was fixed in the reserved Hall slot on the stator" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000790_s1560354708040072-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000790_s1560354708040072-Figure5-1.png", + "caption": "Fig. 5. A pencil with N faces and major radius a has faces of width b = 2a sin(\u03c0/N).", + "texts": [ + " (28), and indicates that rolling friction is important for a \u201ccircular\u201d pencil, while collisional energy losses are important for a hexagonal pencil. 4. ANALYSIS ASSUMING THE IMPULSE ACTS ONLY ALONG THE EDGE NEWLY IN CONTACT WITH THE PLANE This assumption was used in [1], but in an inconsistent manner. A more consistent analysis was given in [2]. 4.1. Asymptotic Velocities at the Beginning and End of a 1/N Turn If the pencil has N faces, then the angle between adjacent major radii is \u03b2 = 2\u03c0/N , as shown in Fig. 5. If a major radius has length a, then the width of a face is b = 2a sin(\u03b2/2) = 2a sin(\u03c0/N). At the end of a 1/N turn the center of mass of the pencil has velocity vector v perpendicular to the major radius from the center of mass to the edge that is the (old) instantaneous axis of rotation, as shown in Fig. 6. As the face of the pencil collides with the plane the instantaneous axis shifts to the adjacent major radius, and an impulsive force is exerted over the colliding face. REGULAR AND CHAOTIC DYNAMICS Vol" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002906_iemdc47953.2021.9449573-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002906_iemdc47953.2021.9449573-Figure2-1.png", + "caption": "Fig. 2. Upper half of the cross section of the analyzed PMSM", + "texts": [ + " Both models show good accordance and the computational time of the analytical model is about 200 times faster compared to the FEA model. This is important for coupling the two domains, because the thermal time constant is much larger than the electromagnetic one. Fig. 1 shows the phase arrangement of the analyzed machine and an exemplary inter-turn fault in phase Ph1. The stator winding is separated in a faulty and healthy part, which allows the simulation of various inter-turn, phase-to-phase or phase-to-ground faults. The corresponding upper half of the cross section is shown in Fig. 2. The colors of the slots match with the colors of the phase arrangement. In general, an interturn fault causes a circulating current IF in the shorted turns. This fault current depends on the operating point and the fault combination, which is given by the number of shorted turns wF and the fault resistance RF. The fault combination cannot be determined during operation, since there are multiple solutions [16]. The power loss generated by the fault is calculated by: Pfault = I2F \u00b7RF (1) An inter-turn fault causes second harmonics in the dq-voltages, due to the arising asymmetry in the stator winding" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000812_icit.2008.4608539-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000812_icit.2008.4608539-Figure8-1.png", + "caption": "Fig. 8. Trajectory of cutting edge", + "texts": [ + " Experiment 3 In order to evaluate the effects of the inclination angle conditions are adjusted as follows: Diameter of cutting tool R=0.1mm Clearance angle =30\u00b0 Number of cutting blade N=1 Shape of the cutting tool: Ball-end mill Depth of cut d=0.01mm Inclination angle = 30\u00b0 or 45\u00b0 Feed rate c=0.1 mm/tooth Two cases with the inclination angles =30\u00b0 and 45\u00b0 are tested. Obtained dimples are shown in Fig.6 and Fig.7. In Fig.6, deformations can be recognized on the dimples in the encircled regions. The reason can be understandable by considering the trajectory of cutting edge. Fig.8 shows how the cutting edge of the ball-end mill moves during the cutting process. The tool center is assumed to move along the dotted line. In collision area, the clearance face precedes the cutting edge. This means the target metal surface collides with clearance face. Therefore, irregular deformation is caused. In order to suppress these unfavorable phenomena, desirable range of the inclination angle is considered. Fig. 6. Dimples with deformation ( =30\u00b0) Suppose, the following cutting conditions are denoted as Diameter of cutting tool is R mm Clearance angle : Number of cutting blade is N Shape of the cutting tool: Ball-end mill Depth of cut is d mm Feed rate is c mm/tooth where is selected enough big value to suppress the collision phenomena shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000890_s12204-008-0171-z-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000890_s12204-008-0171-z-Figure1-1.png", + "caption": "Fig. 1 The model of pile group", + "texts": [ + " Once the axial forces in the fictitious piles have been solved, the desired real pile force distributions as well as the pile head settlement of a two-pile group under two identical vertical loads can be readily determined. The expression for the interaction factor is finally derived and the solution of the present approach is confirmed via comparisons with the available results. A mathematical formulation is presented for the analysis of the interaction factor between two piles under vertical load P0. As shown in Fig. 1, let {0, x, y, z} be a rectangular Cartesian coordinate frame spanning the homogeneous semi-infinite elastic soil medium B. The two embedded piles denoted by B\u2032 1 and B\u2032 2 are assumed to have the same length of L, diameter of d, and circular cross-sectional region of \u03a0z(0 < z < L), respectively. The center-to-center spacing of the two piles is denoted by S. As in the treatment by Muki and Sternberg of this class of problems, the embedding soil medium is extended throughout the half space and two fictitious piles, B\u22171 and B\u22172, are introduced throughout their original locations to account for the presence of the embedded piles (see Fig. 1). The Young\u2019s modulus E\u2217 of each fictitious pile is equal to the difference between that of the real pile and the extended soil, i.e. E\u2217 = Ep \u2212 Es, (1) where Ep is the Young\u2019s modulus of real piles, and Es is the Young\u2019s modulus of soil. In what follows we treat the extended soil B as a 3D elastic continuum characterized by material constants Es and \u03bcs. In contrast, the two fictitious piles B\u22171 and B\u22172 are regarded as 1D elastic continua as far as their constitutive laws and equilibrium conditions are concerned" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001851_978-3-540-74027-8_1-Figure14-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001851_978-3-540-74027-8_1-Figure14-1.png", + "caption": "Fig. 14. Plan view of the robot during rotation about an instantaneous centre of rotation, IC of R", + "texts": [ + " In fact, \u03b4\u03b8ICR, enlarges to, \u0394\u03b8ICR, and, \u03b4\u03b8planet, enlarges to, \u0394\u03b8planet. Hence Eqs. (1) and (2) lead to errors in the robot body following the required path. However, these errors are compensated by inserting a gain factor, k > 1, into the Eq. (1) like this: ICR planet P S 1k R1 R \u239b \u239e \u239c \u239f \u239c \u239f\u0394\u03b8 = \u00b1\u0394\u03b8 \u00d7 \u239c \u239f\u00b1\u239c \u239f \u239d \u23a0 (4) where, k > 1. A schematic view of the six-legged omni-directional robot is shown below in Fig. 13. Locomotion of the robot is produced by the leg tips moving in rectangular curved plane shapes. A plan view of the robot is shown in Fig. 14, which shows the scale of the computational problem to achieve Viennese waltz behaviour because for each leg the value of the leg tip walking angles, (P), and the amplitude of step, (amp), must be computed in real time. The computational problem is made more challenging because the leg tip loci are curved planes whose radii of curvature are the distances, L, from the IC of R. A plan has been worked out using a sun and planet wheel for obtaining Viennese waltz behaviour which means combined translation and rotation of the robot body" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002995_iciccs51141.2021.9432294-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002995_iciccs51141.2021.9432294-Figure2-1.png", + "caption": "Fig. 2. Free body representation of Quanser AERO.", + "texts": [ + " The rest of the paper is organized as: For the problem formulation given in section II, the Controller is designed in Section III. Section IV discusses simulation results followed by conclusions in section V and references. Quanser AERO consists of two rotors viz. front rotor and a back rotor coupled with two high-efficiency dc motor [4].The front rotor predominately affects the motion about the pitch axis while the back rotor mainly affects the motion about the yaw axis. The Quanser AERO test-bed model is shown in Fig.1 The free-body representation of the Quanser AERO is shown in Fig.2. When an input voltage Vp is applied to the front rotor, the speed of rotation causes a force Fp at a distance rp from the Y -axis. Simillarly, yaw motor voltage, Vy genrates a force Fy at a distance of ry from the Y -axis. Both the rotors of Quanser AERO having the same size and equidistant from each other [4], due to this front rotor affects 978-0-7381-1327-2/21/$31.00 \u00a92021 IEEE 1595 20 21 5 th In te rn at io na l C on fe re nc e on In te lli ge nt C om pu tin g an d Co nt ro l S ys te m s ( IC IC CS ) | 9 78 -1 -6 65 4- 12 72 -8 /2 1/ $3 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001107_00207160701477476-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001107_00207160701477476-Figure1-1.png", + "caption": "Figure 1. Left: the surface D. Right: the trigonometric parametrization \u03c4u of D.", + "texts": [ + " The \u2018usual\u2019 trigonometric parametrization of D is given, in homogeneous form, by \u03c4u(\u03b8, \u03c6) = \u23a1 \u23a2\u23a2\u23a3 \u03bc(c \u2212 a cos \u03b8 cos \u03c6) + b2 cos \u03b8 b(a \u2212 \u03bc cos \u03c6) sin \u03b8 b(c cos \u03b8 \u2212 \u03bc) sin \u03c6 a \u2212 c cos \u03b8 cos \u03c6 \u23a4 \u23a5\u23a5\u23a6 where 0 < \u03b8 < 2\u03c0 , 0 < \u03c6 < 2\u03c0 . The homogeneous form is used only to reduce the width of the parametric expressions. The R 3 parametrization may be derived, by division, from the homogeneous form in the usual way as: 1 a \u2212 c cos \u03b8 cos \u03c6 (\u03bc(c \u2212 a cos \u03b8 cos \u03c6) + b2 cos \u03b8, b(a \u2212 \u03bc cos \u03c6) sin \u03b8, b(c cos \u03b8 \u2212 \u03bc) sin \u03c6). Figure 1 shows D in the case of a = 6, b = 4 \u221a 2, (c = 2), and \u03bc = 3 (left); and the origins and directions of the angular parameters, \u03b8 and \u03c6, on the manifold (right). We denote the region parametrized by \u03c4u by D \u2217. Figure 2 shows the surface D \u2217 with the un-parametrized region, which is of area-measure zero, highly exaggerated. In this paper, we use a normalized form \u03c4tr : (0, 1)2 \u2192 D \u2217, which may be expressed as \u03c4tr (t, s) = \u23a1 \u23a2\u23a2\u23a3 \u03bc(c \u2212 a cos(2\u03c0t) cos(2\u03c0s)) + b2 cos(2\u03c0t) b(a \u2212 \u03bc cos(2\u03c0s)) sin(2\u03c0t) b(c cos(2\u03c0t) \u2212 \u03bc) sin(2\u03c0s) a \u2212 c cos(2\u03c0t) cos(2\u03c0s) \u23a4 \u23a5\u23a5\u23a6, where 0 < t, s < 1 are defined by t = (\u03b8/2\u03c0) and s = (\u03c6/2\u03c0)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002411_012034-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002411_012034-Figure2-1.png", + "caption": "Figure 2. Pipe Inspection Gauge (PIG) receiver [35].", + "texts": [ + " Pipeline inspection conducted by the engineer in the oil and gas industry utilized in-pipe robotic system or better known as pipe inspection gauge (PIG) that is driven by fluid. The system is launched from a launcher and then will end at the receiver [20], [32], [33]. A few criteria should be considered before doing the inspection and cleaning using the system such as, internal diameter of the pipe, radius of pipe bending layout, length, flow and pressure conditions, expected debris inside pipe, medium of transition to force the movement of in-pipe system either corrosive or not and pig trap [32], [34]. Figure 1 and Figure 2 show the PIG launcher and receiver produced by [35]. In the early study of the in-pipe inspection robotic system by [36], the researcher had developed the Smart Acquisition and Analysis Module (SAAM) for pipeline inspection. The system was ICATAS-MJJIC 2020 IOP Conf. Series: Materials Science and Engineering 1051 (2021) 012034 IOP Publishing doi:10.1088/1757-899X/1051/1/012034 fitted with differential pressure gauge, high and low sensitivity accelerometer, and temperature sensor. The system developed had a diameter of 100mm, an overall length of 381mm and was tested inside 8inch pipeline" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001278_s1068798x08110087-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001278_s1068798x08110087-Figure1-1.png", + "caption": "Fig. 1. Calculation schemes for a threaded joint with constant (a) and variable (b) pliability of the components.", + "texts": [ + "063 Two methods of calculating threaded joints under the action of an external rupture force F and a tipping moment M were outlined in [1, 2]: (1) under the assumption of constant pliability of the connected parts, when the threaded joint with n screws is replaced by n bushes connected by a rigid diaphragm parallel to the junction plane (Fig. 1a); (2) under the assumption of variable pliability of the connected parts, when the same bushes are connected by a rigid diaphragm inclined to the junction plane (Fig. 1b); on the compression side, the diaphragm touches the junction plane, while on the tensile side it is at a distance \u03a3 h from this plane. The total force F \u03a3 i 1 acting on screw i is calculated as follows: on the assumption of constant pliability (1) on the assumption of variable pliability (2) where F ti is the tightening force in the screw; \u03c7 = \u03bb P /( \u03bb P + \u03bb s ) is the basic load coefficient; \u03bb P and \u03bb s are the pliability values of the part and the screw, i.e., their change in length under the action of unit load; x and x 0 i are the distances from screw i to the y axis; x 0max is the maximum distance between the screws along the x axis, under the assumption that the junction axis y passes through the axes of the extreme screws, on the side of the junction compressed by moment M ; A s = \u03c0 d 3 /4 is the cross-sectional area of the screw; d 3 is the internal diameter of the screw thread; m = A b / A s is the ratio of the bush area to the cross-sectional area of the screw; I s = A s \u03a3 is the moment of inertia of the screws relative to the y axis; S s = A s \u03a3 x i is the static torque of the screws relative to the y axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001084_detc2007-34070-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001084_detc2007-34070-Figure2-1.png", + "caption": "Figure 2. Generalized forces and displacements in the plane of action", + "texts": [ + " The compliance characteristics are generally deduced Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/03/2016 T from preliminary modeling of teeth, by the finite element method for example. From this static calculation, it is possible to deduce the elastic characteristics of elastic coupling between meshing wheels. Distribution of the transmitted load p(s) can be reduced to the pitch point I through the normal force F, and the moment M around the axis perpendicular to the plane of action (see figure 2). Force F is described by equation (2) and moment M is expressed easily in the following form: \u222b \u0393 = dsspsLM )()( (3) Taking into account the assumption of small displacements, for each point Q(s), displacement D(s) can be described by normal displacement dn-s and rotation angle \u03bbs around the axis perpendicular to the plane of action and passing by the pitch point I, so that: Usual models of the meshing process are based on the assumption according to which the effects of the moment M and the angle \u03bbs are negligible" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001408_detc2008-50033-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001408_detc2008-50033-Figure2-1.png", + "caption": "Figure 2 Free body diagram of unsprung mass of two-axled vehicle in steady state cornering (Front View) [6]", + "texts": [], + "surrounding_texts": [ + "L \u2013 \u0394w (8)" + ] + }, + { + "image_filename": "designv11_83_0002873_s00170-021-07615-0-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002873_s00170-021-07615-0-Figure2-1.png", + "caption": "Fig. 2 Numerical model. (a) Meshing of FE model. (b) Boundary condition. (c) Contact surface", + "texts": [ + " First, the finite element model of the bolted connection is established, while the deterministic load is applied to the bolt connection after pretightening for simulation calculation. Croccolo [34, 35] assumed that the friction coefficient of the threaded joint friction interface would change during pre-tightening. The threaded connection is used in the finite element model because the friction coefficient of the contact surface needs to be considered, when the nut is pre-tightened. In this study, the friction coefficient of the contact surface is set to 0.15. The finite element mesh of each component and assembly is shown in Fig. 2. Taking into account the calculation accuracy and efficiency, the threaded connection and the overlapping area of the connecting plate where slippage occurs are finely meshed. The hexagonal shape of the bolt head and nut is simplified into a circle, the same as in previous work [36]. The material characteristics of the bolt connection are listed in Table 1. Figure 3 illustrates the interface pressure distribution diagram when a tightening torque of 50 nm is applied to the nut. In the bolt connection model after pre-tensioning, the stress concentration area appears at the first few turns of the thread, indicating that the screw and the nut are completely matched" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000332_50003-4-Figure3.33-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000332_50003-4-Figure3.33-1.png", + "caption": "Fig. 3.33. Enhanced model with deflected carcass and tread element that is followed from front to rear.", + "texts": [ + " With the updated belt deflection the next passage through the zones is performed and the calculation is repeated. Here, we will restrict the discussion to steady-state slip conditions and take a single zone with length equal to the contact length. In Section 2.5 an introductory discussion has been given and reference has been made to a number of sources in the literature. The complete listing of the simulation program TreadSim written in Matlab code is given in Appendix 2. For details we may refer to this program. Figure 3.33 depicts the model with deflected belt and the tread element that has moved from the leading edge to a certain position in the contact zone. In Fig.3.34 the tread element deflection vector e has been shown. The tread element is assumed to be isotropic thus with equal stiffnesses in x and y direction. Then, when the element is sliding, the sliding speed vector Vg, that has a sense opposite to the friction force vector q, is directed opposite to the deflection vector e. The figure depicts the deflected element at the ends of two successive time steps i- 1 and i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002468_j.ast.2021.106656-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002468_j.ast.2021.106656-Figure1-1.png", + "caption": "Fig. 1. Engagement geometry for the i-th and the j-th agent.", + "texts": [ + " Thus, at each time step, the current speed is taken as constant, thus facilitating the application of the algorithm to variable-speed agents. The nonlinear kinematic equations for the engagement under the aforementioned assumptions for the i-th and the j-th agent duos are r\u0307Di D j = V D j cos ( \u03b3D j \u2212 \u03bbDi D j ) \u2212 V Di cos ( \u03b3Di \u2212 \u03bbDi D j ) (1a) \u03bb\u0307Di D j = [ V D j sin ( \u03b3D j \u2212 \u03bbDi D j ) \u2212 V Di sin ( \u03b3Di \u2212 \u03bbDi D j )] /rDi D j (1b) where V and r are the speed of the vehicles and the range between the vehicles respectively. The angles \u03b3 and \u03bb are the flightpath and LOS angles respectively. Fig. 1 presents a schematic view of the engagement geometry for the i-th and the j-th agents, where XI \u2212 O I \u2212 Y I is a Cartesian inertial reference frame. The addition of the subscript \u20320\u2032 indicates the value of the parameter at the beginning of the engagement. The velocity vector of the vehicles is assumed perpendicular to the acceleration vector. Thus, their ratio defines the rate of the flight-path angle: \u03b3\u0307k = ak/Vk, k = { Di, D j } (2) where a is the normal acceleration. By assuming that the controller dynamics of all the participants can be represented by linear differential equations, the controller dynamics can be formalized as: x\u0307k = Akxk + Bkuk ak = Ckxk + dkuk k = { Di, D j } (3) where xk is the vector of an agent\u2019s internal state variables and uk is its controller", + " For a comprehensive presentation of the connection between the controller dynamics and the longitudinal state-space model comprised of the UAV\u2019s stability and control coefficients, please see [22, Chapter 5.5.3]. A numerical example of the aforementioned coefficients is provided in [22, Appendix E.2]. Note that while the controller dynamics are assumed to be linear, the kinematic equations are clearly non-linear, and are linearized in the following section. The relative displacement between the i-th and the j-th agent, normal to L O S Di D j 0 (see Fig. 1) is noted by yDi D j . By assuming that the velocities of the vehicles are constant, the kinematics around the initial LOSs can be linearized, thus the dynamics of all the participants can be written as x\u0307i = \u23a7\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a9 x\u0307Di D j ,1 = xDi D j ,2 x\u0307Di D j ,2 = aD j \u2212 aDi x\u0307Di = ADi xDi + BDi uDi x\u0307D j Di ,1 = xD j Di ,2 x\u0307D j Di ,2 = aDi \u2212 aD j x\u0307D j = AD j xD j + BD j uD j (4) where the state vector of the linearized problem is xi = [ yDi D j y\u0307Di D j xT Di yD j Di y\u0307D j Di xT D j ] (5) Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000750_oe.15.004671-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000750_oe.15.004671-Figure1-1.png", + "caption": "Fig. 1. Schematic diagram of a double-pass pumping Raman fiber laser", + "texts": [ + " [9] presented the exact analytical solution for the single-pass pumping RFL with the assistance of Lambert W function, but no explicit expression was provided. In this paper, we obtain an explicit analytical solution for double-pass pumping RFL under a linear-attenuation approximation for pump propagation. The approximation has been applied to the second-order RFL and proved to be valid [10]. The proposed explicit solution provides us a clear physical understanding to the optimal design of the laser. The schematic diagram of double-pass pumping Raman fiber laser is shown in Fig. 1. A pair of fiber Bragg grating reflectors, i.e. FBG1 and FBG2, form the Fabry-Perot resonant cavity at the first Stokes wavelength. An additional reflector (i.e. FBG0) with high reflectivity at the pump wavelength yields a double-pass pumping scheme. The reflectivity of FBG0 and FBG2 is larger than 99% at pump and Stokes wavelengths, respectively. FBG1 with relative low reflectivity at the Stokes wavelength can couple the Stokes lights out of the cavity. The forward- and backward-propagated pump and Stokes powers in Raman gain fiber meet the following well-known differential equations [1, 11] )( 1 11 0 1 0 0 0 \u2212+ \u00b1 \u00b1 +\u2212\u2212=\u00b1 PPg dz dP P \u03bb \u03bb\u03b1 (1a) )( 1 001 1 1 \u2212+ \u00b1 \u00b1 ++\u2212=\u00b1 PPg dz dP P \u03b1 (1b) where the subscripts i represent pump(i=0) and Stokes(i=1) waves" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000636_bfb0119382-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000636_bfb0119382-Figure7-1.png", + "caption": "Figure 7. (Left) Non-holonomic Model of a single Orientation Step O. The orientation 0 of the box as a function of the control parameter d, modeled in the non-holonomic configuration space ]~ x S 1. This figure shows one half (O) of an orientational flossing cycle (O, OR). (Right) Non-holonomic Model of a Translation Step T. x represents the position of the box. x O[", + "texts": [ + " Our system admits an isotropic non-holonomic model as follows: Let R1 and R2 be the positions of Bonnie and Clyde (respectively), and let r be one half the diameter of the manipulandum (for a sensible definition of the diameter of a polygon, see [12]). Although in general Ri is two dimensional, for the sake of illustration, assume that R1, R2 E ]~. With the rope taut, the distance between R1 and R2 may be closely approximated as fixed once the box is wrapped. Thus the combined, simultaneous configurations of R1 and R2 may be collapsed into a single parameter d E R, representing the rope position. Finally, let 0 be the orientation of the box, and x its translational position. Now, let us consider a single orientation step O (Figure 7-1eft). When the robots move to manipulate the box, it translates and rotates in a non-linear fashion that depends on the box geometry (Figure 7). For the orientation task, the progress this strategy makes is the average slope of the linearization of this curve in d-O space. This slope is approximately (27rr) -1. Non-holonomic analyses can also be done for translation (T), flossing (O, OR)*), translational ratcheting (T,G)*, and orientational ratcheting (O, G)*. (Figures 7, 8, and 9). By interleaving regrasp steps with translation steps, the box can be translated repeatedly; the motion in configuration space is like a 'ratchet ' (Figure 8)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002241_j.actaastro.2021.02.001-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002241_j.actaastro.2021.02.001-Figure7-1.png", + "caption": "Fig. 7. \u2018\u2018Signal Flag\u2019\u2019 steps after legs are bent. (a and b) Thighs are bent to 88\u25e6 from causes v\ud835\udc5a\ud835\udc4e\ud835\udc65 to tilt significantly away from being parallel to AP axis. (c) Thighs bent to 30\u25e6 from causes v\ud835\udc5a\ud835\udc4e\ud835\udc65 to tilt 5.5\u25e6 from AP axis. (d) Plot of v\ud835\udc5a\ud835\udc4e\ud835\udc65 in for each of the previous figures. v\ud835\udc5a\ud835\udc4e\ud835\udc65 vectors orient from the location of \ud835\udc29\ud835\udc50\ud835\udc5c\ud835\udc5a in world (inertial) coordinate system.", + "texts": [ + " Raising one arm above the head, results in an additional \ud835\udee5\ud835\udc29\ud835\udc50\ud835\udc5c\ud835\udc5a = (3.8 cm)?\u0302?. Rotation in this maneuver is the result of the motion of the arms. To maximize the amount of rotation in the rest of the body, it would be ideal to bring \ud835\udc29\ud835\udc50\ud835\udc5c\ud835\udc5a as close to the center of mass of the arm subsystem as possible. This would make the analysis in the previous paragraph seem ideal. However, let us consider the angle \ud835\udef7 to be the angle between v\ud835\udc5a\ud835\udc4e\ud835\udc65 and the AP axis. When the legs are brought toward the torso, v\ud835\udc5a\ud835\udc4e\ud835\udc65 tilts away from being parallel with the AP axis (see Fig. 7(a)), where \ud835\udef7 \u2248 20\u25e6. This would be enough to cause significant distortions away from stable motion if the arms rotate in . Additionally, when raising the arm overhead, \ud835\udef7 \u2248 22\u25e6, see Fig. 7(b). It also rotates out of the sagittal plane , giving it a significant ?\u0302? component. Fig. 7(d) shows the v orientation for these positions \ud835\udc5a\ud835\udc4e\ud835\udc65 as viewed from the front. If a person were to attempt this maneuver in weightlessness without an understanding of the orientation of v\ud835\udc5a\ud835\udc4e\ud835\udc65, they may conclude the movement is fundamentally not stable. The deviation can be reduced significantly by only bringing the knees to about 30\u25e6 forward of , see Fig. 7(c). This reduces \ud835\udef7 \u2248 5.5\u25e6, while keeping v\ud835\udc5a\ud835\udc4e\ud835\udc65 very close to . Rotation of the arms in would then cause much less wobble. We now find that \ud835\udee5\ud835\udc29\ud835\udc50\ud835\udc5c\ud835\udc5a = (1.9 cm)?\u0302? + (10.2 cm)?\u0302?. This is less than if the thighs are brought further from but allows for increased stability during rotation (see Fig. 8). If this movement is used as part of a choreography, it may be desired that it would be a continuous motion rather than a discrete set of rotations. The sequence proposed in this section provides a technique for continuous motion about the AP axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001413_j.jsv.2008.02.056-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001413_j.jsv.2008.02.056-Figure3-1.png", + "caption": "Fig. 3. Meridional stress diagram for H1.", + "texts": [ + " To determine the displacements d1H1 ; d2H1 ; d3H1 ; d4H1 , and dVH1 ; we first apply the virtual loading H 01 at and in the direction of ring coordinate 1 prior to application of the actual loading H1, where, typically, H1 is the radial force per unit length in the direction of coordinate 1, and where, typically, d1H1 represents the deflection in the direction of coordinate 1 due to H1, and where, typically, H 01 is a unit virtual radial force per unit length in the direction of coordinate 1. Subsequently, we apply the actual loading H1. The meridional stress diagram due to H1 as shown in Fig. 3 is identical to the meridional stress diagram (not shown) due to the virtual loading H 01. Under these loadings (and by using Eqs. (1), (8), and (10)), our actual stresses sx are seen to be sx \u00bc H1R1c Iz0 y0 \u00fe H1R1 A \u00bc H1R1c Iz0 \u00f0y cos a z sin a\u00de \u00fe H1R1 A . (18a) It may be seen from Fig. 3 and discussions hereinbefore relative to Eq. (13) that the stresses sy are represented very closely by sy \u00bc H1R1 sin a Rt y \u2018 \u00fe 1 2 , (18b) while the virtual stresses are sx\u00f0H 01\u00de \u00bc H 01R1c Iz0 \u00f0y cos a z sin a\u00de \u00fe H 01R1 A , (19a) sy\u00f0H 01\u00de \u00bc H 01R1 sin a Rt y \u2018 \u00fe 1 2 . (19b) ARTICLE IN PRESS T.A. Smith / Journal of Sound and Vibration 318 (2008) 428\u2013460438 The actual strains are 2x\u00f0H1\u00de \u00bc 1 E \u00f0sx nsy\u00de \u00bc H1R1 E c Iz0 \u00f0y cos a z sin a\u00de \u00fe 1 A \u00fe n sin a Rt y \u2018 \u00fe 1 2 , (20a) 2y\u00f0H1\u00de \u00bc 1 E \u00f0sy nsx\u00de \u00bc H1R1 E sin a Rt y \u2018 \u00fe 1 2 nc Iz0 \u00f0y cos a z sin a\u00de n A " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001001_aim.2008.4601625-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001001_aim.2008.4601625-Figure3-1.png", + "caption": "Fig. 3 Quadruped robot model in ADAMS/View.", + "texts": [ + " 1, We have: 2 2 1 3 3 2 2cos( ) cos( ) ( ) ( )o o o o i i i i i Ti Bi Ti Bia a a x x y y\u03b1 \u03b1+ + = \u2212 + \u2212 978-1-4244-2495-5/08/$25.00 \u00a9 2008 IEEE. 1 Proceedings of the 2008 IEEE/ASME International Conference on Advanced Intelligent Mechatronics July 2 - 5, 2008, Xi'an, China 4 3 3 2 2sin( ) sin( )i i i i i ia a a h\u03b1 \u03b1+ + = (2) 1tan( ) o o Ti Bi o o Ti Bi y y x x \u03b1 \u03ba \u2212= \u2212 where 1=\u03ba for 2=i or 3, 1\u2212=\u03ba for 1=i or 4 and where ih denotes the initial height of i oB . Maximum lateral stretch of a leg can be expressed as; We use ADAMS for our simulations and a scene from this program can be seen from the Fig. 3. III. GAIT GENERATION In this section three different straight going gait types are discussed. The basic principle of quadruped walking is that, the ground touching tip points should form a triangular support area which embraces the CoG so the stability of the body is already maintained and then joint actuators make the body shift. Therefore, walking can be considered as an inverse kinematics problem. The swinging leg has no contribution to the progress of the body. Legs swing in an appropriate sequence for the locomotion of the body and forming the new support plane in the next step" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001652_bf03177400-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001652_bf03177400-Figure1-1.png", + "caption": "Fig. 1. Design parameters and sensing area.", + "texts": [], + "surrounding_texts": [ + "The basic first step in designing the laser vision sensor was to determine the general design para meters based on function, environment and restric tions of the application. The detailed design model, such as size, weight, resolution , and measurement area of the sensor must be established from the general design conditions. In this study, the main role of laser vision sensor is to acquire the distance information of the measurement area with desired resolution. The design parameters related to this role are the distance and angle between the laser and CCO camera and the resolution of CCO. Since the latter is generally fixed , the former are the main design parameters. Fig. I shows the design parameters and sensing area. In Fig. I , the distance (C) and angle ( o) 2. Laser vision sensor system" + ] + }, + { + "image_filename": "designv11_83_0002022_mmvip.2007.4430742-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002022_mmvip.2007.4430742-Figure5-1.png", + "caption": "Fig. 5. Condition of catopter is a convex surface in RIM-FOS.", + "texts": [ + " When light is reflected through convex catopter, it is equivalent to the sending out light from the A\u2032 forming emanated light-cone \u2206 HGA' . In this case, the light which the angle is \u03b8\uff2e through the catopter is reflected, the reflecting light NH can not be received by RF, because of the result of the formula \u03b8\uff2e+ 2\u03b1 is bigger than the angle \u03b8N. The condition can be only considered, which the light in the area of available light-cone \u2206A\u2032JK (the light-cone angle is 2\u03b8N) can be received by RF. From the geometrical relationship showing by Fig. 5, we can get the following equations as. )2sin( )sin(' \u03b1\u03b8 \u03b1\u03b8 + += N NRCA C (5) CARdDAdEA C ''' \u2212+=+= (6) qEKEA N ==\u03b8tan' (7) )2tan(' \u03b1\u03b8 += NEAEH (8) where q is the radius of facula which the available reflex light-cone \u2206A\u2032JK forms in the fibre end. On the premise that reflector is completely smooth and the light intensity is uniform distribution in the reflex light- cone, the fibre intensity modulated coefficient is equivalent to the ratio of the facula overlapping area rS of the RF end face and effective light-cone which the radius is q in the fibre end face and the area of reflex light-cone (a radius of EH) in the fibre end face [2]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002772_978-3-030-67750-3_14-Figure11.3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002772_978-3-030-67750-3_14-Figure11.3-1.png", + "caption": "Fig. 11.3 Proposed layout", + "texts": [ + " In this proposed layout the number of workers is reduced to three persons to operate the production system. This proposed layout has reduced the workers and the material\u2019s trolley has been redesigned and changed to a location which is to near the conveyor A. The function of worker 1 is to transfer the part from conveyor A direct to the material\u2019s trolley. Then, worker 2 is stamping the product\u2019s logo to the part. While worker 3 requires to transfer the part to pallet into the next workstation. The time required and line balancing for this proposed workstation is as follows (Fig. 11.3): Tare weight (s) = 6.862 s Logo stamp (s) = 2.451 s Worker = 3 Line balancing = (6.862+ 2.451)\u00d7 100% (6.862)\u00d7 (3) Line balancing = (931.3) (20.6) Line balancing = 45.2% A comparison of the results before and after the improvement process is shown in Fig. 11.4. It compares from the existing layout of the company and the proposed layout in this research. Figure 11.4 indicates that improvements occur at bottleneck stations which is the tare weight station. By that, the objective of the research has been achieved by minimizing the cycle time" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001350_imece2008-67917-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001350_imece2008-67917-Figure1-1.png", + "caption": "Figure 1: Section View of the original projectile", + "texts": [ + " Find the effects of using polyurethane rubber to support the electronic board on transmitted accelerations to electronic components. 3. Create FEA models that can produce results that closely match corresponding experimental ones. Achieving these objectives can help engineers better understand the behavior of electronic components within projectiles as well as help improve FEA of these systems. U.S. Army ARDEC Tank-automotive & Armaments Command-Armaments Research, Development & Engineering Center (TACOM-ARDEC) provided the projectile under consideration. The projectile consists of several components as shown in Figure 1. All these components are threaded or bolted together. An experimental test fixture with the same mass and mass moment of inertia as the original projectile is used in this research, Figure 2 and Figure 3. The fixture has two components: impactor and housing. Impactor has a tapered end where an impact hammer is used to send load to an electronic board. The electronic board is situated between the impactor and the housing. A 4.5\u201d-15 thread is used to connect the impactor and the housing. The outer surfaces of the impactor and housing have flat surfaces to allow tightening them together" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000790_s1560354708040072-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000790_s1560354708040072-Figure4-1.png", + "caption": "Fig. 4. When the radius of length a of the pencil makes angle \u03b1 to the vertical, the perpendicular distance from the center of mass to the inclined plane is a cos(\u03b1 \u2212 \u03b8).", + "texts": [ + " (21) predicts an asymptotic rolling velocity of 5 cm/s. 2.3. Will the Pencil Lose Contact with the Plane? If the normal force of the inclined plane on the pencil goes to zero, the pencil will lose contact with the plane.2) The limiting case is that the component mg cos \u03b8 of the force of gravity perpendicular to the plane is just sufficient to provide the acceleration of the center of mass of the pencil in this direction. Since the perpendicular distance of the center of mass from the plane is a cos(\u03b1 \u2212 \u03b8), as shown in Fig. 4, the acceleration of the center of mass towards the plane is a\u03b1\u0308 cos(\u03b1 \u2212 \u03b8) = mga2 I sin \u03b1 cos(\u03b1 \u2212 \u03b8) = g k sin \u03b1 cos(\u03b1 \u2212 \u03b8), (22) recalling Eq. (2). The limiting condition is that, at the maximum angle \u03b8max of inclination for rolling without slipping, m times the acceleration (22) equals the normal component of mg, or k cos \u03b8max = sin \u03b1 cos(\u03b1 \u2212 \u03b8max). (23) Since angle \u03b1 is never far from \u03b8, the condition (23) is roughly that tan \u03b8max \u2248 k, (24) which is weaker than the condition (1), tan \u03b8 < \u03bc, for realistic values of the coefficient of static friction \u03bc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000971_msf.537-538.599-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000971_msf.537-538.599-Figure4-1.png", + "caption": "Figure 4. Schematic of modelled geometry along the symmetry plane", + "texts": [ + "114, Freie Universitaet Berlin, Berlin, Germany-12/05/15,10:20:02) shown in the flow chart in Fig. 1. Coupled thermal and metallurgical calculations were used for different calculations (Fig. 2). Geometry configuration. The process of laser transformation hardening and the geometry of the problem being solved are illustrated in Fig. 3. The dimensions of the workpiece used are 60 mm long, 10 mm thick and 56 mm wide. The single track laser hardening was applied exactly at the centre of the workpiece. For sake of simplicity, half of the workpiece was modelled as shown in Fig. 4. Thermal formulation. In thermal analysis, the transient temperature field T is a function of time t and the spatial coordinates (x, y, z), and is determined by the tree-dimensional non-linear heat transfer using heat conduction equation: ( ) { } t TCtyzxQTk \u2202 \u2202 =+\u2207\u2207 \u03c1,,, (1) where k is thermal conductivity, T is temperature, Q is the rate of heat generation per unit time, t is time, \u03c1 is the density and C is the specific heat capacity. The heat source, which moves along a defined path, is modeled as a heat flux density applied on the elements along the path" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002381_j.matpr.2021.01.637-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002381_j.matpr.2021.01.637-Figure4-1.png", + "caption": "Fig. 4. Tooth profile wear of 2%, 4%, 6%, et 8%.", + "texts": [ + " Considering the OBI triangle: sin I _ Ra \u00bc sinO _ IB \u00bc sin B _ R arc sin I _ Ra R 0 @ 1 A \u00bc B _ with Ra the head radius, R the pitch radius and the angle at I which is equal to I _ \u00bc p 2 \u00fe a, with a the pressure angle. aracteristic. is figure legend, the reader is referred to the web version of this article.) O _ \u00bc p I _ \u00fe B _ which allows us to obtain the segment [IB], IB \u00bc sin O _ sin I _Ra, [IB] is the retreat distance. Thus for radius going from the head to the clearance radius, for a point M of radius Ri, its abscissa S on the action line which corresponds to the segment [IM] is found, I being the origin. With Ri the radius included between the head and clearance radius (Fig. 4). In this study SolidWorks Finite element module was used for the modeling of straight cylindrical gears along with the guiding curves method. This method makes it possible to plot the curves from precise coordinates obtained in a text file provided by the VBA computer programming software and its pseudo code is illustrated in Fig. 6a. SOLIDWORKS software considers the macros from VBA, can read a file of points that it can represent with much precision in a form of curve. After reading all the curves of the tooth by this method, the representation of the latter consisting of the two main profiles which are the involute of circle and the trochoid" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002925_01423312211016188-Figure15-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002925_01423312211016188-Figure15-1.png", + "caption": "Figure 15. Test bearings with different size outer ring fault.", + "texts": [ + " In theory, force and compression of a single roller in the load bearing area can be calculated using relevant formulae, and the relationship between ud and Td can be further deduced. However, the internal force of working bearing changes dynamically and the fault state has differences from theoretical formulae. The proposed model-based measurement method can theoretically solve this problem, and is tested in this section. In order to verify method effectiveness, the vibration signals of outer ring fault bearing with two fault sizes were collected. The test bearings are shown in Figure 15, and the operating parameters of different samples are listed in Table 3. Other experimental conditions are the same as those in the section \u2018Model verification\u2019. The actual dual impulse time intervals are shown in Figure 16. By comparing pulse frequency with bearing fault characteristic frequency or using other proposed diagnosis methods, the bearing fault type can be determined, which is not discussed in this paper. We continue to use the proposed method to measure bearing fault, and the benchmark conditions are set to be the same as those in the section \u2018Rolling bearing local fault vibration model\u2019, the section \u2018Vibration mechanism analysis\u2019 and Figure 14(d), that is, 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001730_tmag.2007.914649-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001730_tmag.2007.914649-Figure1-1.png", + "caption": "Fig. 1. Ideal induction motor.", + "texts": [ + " From (17) and (19), we have (29) On combining (29) and (25), one obtains (30) By a similar procedure, the following expressions can be derived for and : (31) (32) (33) Finally, by applying (30), (31), (32), and (33) in (24), one arrives at an expression in the form of (7). When the torque is calculated with SFEM and there are two or more layers of elements in the air gap, the result depends on the factor assigned to the \u201cnodal derivatives.\u201d If (24) or (7) are used instead, this drawback is eliminated. The proposed method is verified by applying it in the calculation of torque for the IIM depicted in Fig. 1 and whose characteristics are listed as follows [9]: \u2022 The IIM presents four regions: each one is made of an isotropic, homogenous, and linear material. At each region, the permeability and the conductivity are constant scalars. \u2022 The surfaces are considered smooth; that is, teeth and grooves are disregarded. \u2022 The source is a traveling wave current sheet placed in the stator air-gap-boundary: , where should be an integer, since the number of pairs of poles is equal to . Now the torque is computed for the IIM by means of (7) and (24) using FEMA, as well as through the Maxwell stress and virtual work methods using SFEM, for three levels of refinement in the mesh: coarse, medium, and fine" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001537_09544062jmes814-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001537_09544062jmes814-Figure2-1.png", + "caption": "Fig. 2 Cross-section of a couple-stress fluid film journal bearing", + "texts": [ + " Mechanical Engineering Science JMES814 # IMechE 2008 at UNIV OF CONNECTICUT on May 19, 2015pic.sagepub.comDownloaded from spectra, Poincare\u0301 maps, bifurcation diagrams, and the Lyapunov exponent are applied to analyse the rotor\u2013bearing system. 2 MATHEMATICAL MODELLING Figure 1 shows a flexible rotor supported by two couple-stress fluid film journal bearings on a parallel with non-linear spring. Om is the centre of the rotor gravity, O1 the geometric centre of the bearing, O2 the geometric centre of the rotor, and O3 the geometric centre of the journal. Figure 2 shows the crosssection of the fluid film journal bearing, where (X, Y) is the fixed coordinate and (e, w) is the rotated coordinate, e being the offset of the journal centre and w being the attitude angle of the X-coordinate. Considering the journal centre O3, the resulting viscous damping forces in the radial and tangential directions are shown in Fig. 2. From the equilibrium of force, the forces applied to the journal centre O3 are Fx \u00bc fe cosw\u00fe fw sinw \u00bc kp \u00f0X2 X3\u00de 2 \u00f01\u00de Fy \u00bc fe sinw fw cosw \u00bc kp \u00f0Y2 Y3\u00de 2 \u00f02\u00de where fe and fw are the resulting viscous damping forces in the radial and tangential directions. To analyse this rotor\u2013bearing system, the following assumptions are made. 1. The rotor mass and the bearing mass are lumped at the midpoint. 2. The rotor, the bearing, and the support of bearing housing are radially symmetric and the damping at rotor midpoint due to aerodynamics is viscous" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002993_s00170-021-07405-8-Figure11-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002993_s00170-021-07405-8-Figure11-1.png", + "caption": "Fig. 11 Effective stress distributions of finished sun gear under different sizing amounts. a 0.1 mm. b 0.3 mm. c 0.5 mm", + "texts": [ + " The influences of key process parameters on forming accuracy of external gear and internal spline were investigated through FE simulations to determine the optimal process parameters of precision sizing operation. The generally observed influencing parameters included the sizing amount of external gear, interference value of internal spline, die bearing length, and friction factor. The sizing amount of external gear is generally considered a key process parameter influencing the tooth accuracy of sun gear. To obtain the effects of sizing amount on gear accuracy, the sizing amount was selected as 0.1, 0.2, 0.3, 0.4, and 0.5 mm for numerical simulation. Figure 11 shows the effective stress distribution of finished sun gear under different sizing amounts. When the sizing amount of external gear was 0.1 mm, less plastic deformation was observed at the tooth tip and root of external gear, while only elastic deformation occurred in the internal spline after the part was finished, as displayed in Fig. 11a.When setting the sizing amount to 0.3 mm, the tooth surfaces of internal-external gears had a proper amount of plastic deformation, and the shaping effect was better, as shown in Fig. 11b. This can be explained as the sizing amount increases, the deformation area gradually expands to inner layer. When defining the finishing amount as 0.5 mm, the tooth surfaces of both external gear and internal spline have undergone severe plastic deformation, which can no longer achieve the purpose of finishing, as illustrated in Fig. 11c. The deviation curves of internal-external gears under different sizing amounts are represented in Figs. 12, 13, and 14, respectively. It can be observed from Fig. 12 that the single profile deviation of external gear gradually reduced with the increase of tooth width. In addition, when the finishing amount of external gear was 0.1 ~ 0.3 mm, the total profile deviation was stable at about 0.016 mm; when the sizing amount was greater than 0.3 mm, the total profile deviation increased remarkably. Figure 13 shows helix deviation curves of external gear with finishing amount" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002204_s40435-021-00757-9-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002204_s40435-021-00757-9-Figure3-1.png", + "caption": "Fig. 3 A free body diagram of 2-DOF helicopter system", + "texts": [ + " 2i\u2013l, in the steady-state chattering phenomenon, can be observed with a NESO and an SMO. The effect of chattered estimated states and their derivatives can be seen in the control inputs in the comparative results.As the magnitude of chattering in the estimated states increases, it induces chattering in the control inputs and eventually in the system states. The 2-DOF helicopter model is a highly non-linear system with inputs coupling. A free body diagram of the 2-DOF helicopter system is shown in Fig. 3. It consists of two BrushlessDCmotors to controlmotions over the pitch axis and yaw axis. The non-linear dynamic model of 2-DOF Helicopter system [26] is given by \u03b8\u0308 = \u2212 Bp \u03b8\u0307 + ml2\u03c8\u03072 sin \u03b8 cos \u03b8 + mgl cos \u03b8 Jp + ml2 + kpp Jp + ml2 u p + kpy Jp + ml2 uy (23) \u03c8\u0308 =2ml2\u03c8\u0307 \u03b8\u0307 sin \u03b8 cos \u03b8 Jy + ml2 cos2 \u03b8 \u2212 By\u03c8\u0307 Jy + ml2 cos2 \u03b8 + kyp Jy + ml2 cos2 \u03b8 u p + kyy Jy + ml2 cos2 \u03b8 uy (24) where \u03b8 and \u03c8 are the pitch angle and yaw angle respectively. The motor voltages u p and uy are the control inputs to be designed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002432_tasc.2021.3064529-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002432_tasc.2021.3064529-Figure1-1.png", + "caption": "Fig. 1. Different types of HTS-LFSMs (a) HTS-LFSM with segmented secondary. (b) HTS-LFSM with slotted secondary.", + "texts": [ + " Recent literatures show that Linear Flux-Switching Permanent Magnet (LFSPM) motors have the merits of higher power factor, higher efficiency than LIM does while also has simple structure of secondary [5]. Despites these merits, due to the adoption of Permanent Magnets (PMs), the flux-weakening ability of LFSPM motors is influenced, leading to a narrow speed range, which limits their further application in long-stroke transportation, such as urban railway transit system. To solve the aforementioned problem, it is feasible to adopt High Temperature Superconducting (HTS) coils to replace the PMs [6], [7]. As Fig. 1 shows, HTS Linear Flux-Switching Motor (HTS-LFSM) combines both high electromagnetic performance and good flux-weakening ability at the same time, thus becoming a promising candidate for urban railway transit system. So far, researches on HTS-LFSM has been focused on its topology and working principles. In [8], a linear modular HTS flux-switching motor is proposed, but its power is relatively low. In [9], experiments are conducted to a segmented HTS fluxswitching motor, suggesting the feasibility of such topology" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001405_1352793.1352901-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001405_1352793.1352901-Figure2-1.png", + "caption": "Figure 2. Distance between the robot and the obstacle.", + "texts": [ + " If the CV value is bigger than the threshold value, an obstacle exists in that position. We denote the distance between the obstacle and robot as \u201cD\u201d The distance of x and y between the robot and the obstacle can be obtained by equations (1), (2). ,\u03b8SinDx \u00d7= (1) .\u03b8CosDy \u00d7= (2) After x and y values are calculated, the robot finds the position of the obstacle using the current coordinate of the robot. As the grid dimension is acm, x and y are divided by a to give the distance between the obstacle and robot within the array. Figure 2 shows the method to calculate the distance between the robot and the obstacle. In this example (Figure 3) the distance between the robot and obstacle is 50cm in the direction of the x axis and 0cm of the y axis. Each grid unit is 30cm. So, the value from x/30 and y/30 is the distance between the robot and obstacle. The obstacle is located 2 units to the left and 0 above the current position of the robot. The value of CV increases by 1 to indicate the possible existence of an obstacle. In the experiment, because the robot checks obstacles every 100ms, the value of CV increases at 10 per second" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003385_ur52253.2021.9494662-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003385_ur52253.2021.9494662-Figure4-1.png", + "caption": "Fig. 4. Mecanum mobile platform", + "texts": [ + " The spring was selected using (1). Fig. 3 shows the internal structure of the suspension module; the parameter values of the equation are summarized in Table III. Fs = Mg Ns \u2206x = Fs N (1) A mobile platform was developed using the mecanum in-wheel motor and the suspension module. It has the advantages of holonomic system owing to an omni-directional movement. The body consists of aluminum profiles in which the mobile platform can be run at a high speed of up to 20 km/h and has a maximum payload of 20 kg. Fig. 4 shows the mecanum mobile platform; the detailed specifications are summarized in Table IV. TABLE IV SPECIFICATION OF MECANUM MOBILE PLATFORM Size 0.604\u00d7 0.598\u00d7 0.205 m Weight 15.3 kg Payload 20 kg Wheel 6 Inch Max velocity 20 km/h Battery 25.6 V, 16.5Ah, LeFeSO4 126 Authorized licensed use limited to: University of Glasgow. Downloaded on August 12,2021 at 13:49:11 UTC from IEEE Xplore. Restrictions apply. A mobile manipulator that can move in any direction at a high speed was developed by combining the 6-DOF manipulator and the mecanum mobile platform presented above", + " The inverse kinematics of the manipulator can be computed through a numerical analysis using the Jacobian. The Jacobian defined based on the end-effector is shown in (3). Jb(\u03b8) = [Jb1 Jb2 Jb3 Jb4 Jb5 Jb6] (3) where, Jbi vector can be calculated following equation and Ad corresponds to adjoint. Jbi = Ad e\u2212[Bn]\u03b8n \u00b7\u00b7\u00b7e\u2212[Bi+1]\u03b8i+1 (Bi) (i = n\u2212 1, \u00b7 \u00b7 \u00b7 , 1, Jbn = Bn) The mecanum mobile platform is a holonomic system in which omni-direction movement is possible. The kinematics of the robot in the case of the mecanum mobile platform is shown in (4). Fig. 4 shows kinematic model. u = HVbase = 1 r \u2212l \u2212 \u03c9 1 \u22121 l + \u03c9 1 1 l + \u03c9 1 \u22121 \u2212l \u2212 \u03c9 1 1 \u03c9bz vbx vby (4) where, Vbase = [ \u03c9bz vbx vby ]T and H corresponds to the pseudo inverse Jacobian of the mecanum mobile platform. The wheel speed of the mobile platform can be calculated using (4). The odometry of the mecanum mobile platform refers to obtaining the robot\u2019s position information based on the rotation information of the wheel. Equation (4) can be changed to (5) to obtain the velocity component Vbase of the mobile platform for odometry" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002838_012021-Figure10-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002838_012021-Figure10-1.png", + "caption": "Figure 10. Stress Distribution in Structural Steel", + "texts": [ + "1. Total Deformation The Max. and Min. Total Deformation in Structural Steel is 0.18644 mm and 0 mm respectively shown in Figure 9. The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.2.2. Stress Distribution The Max. and Min. Stress Distribution in Structural Steel is 183.73 MPa and 1.8481 MPa respectively shown in Figure 10. 3.2.3. Strain Distribution The Max. and Min. Strain Distribution in Structural Steel is 0.0010678 and 0.000012423 respectively shown in Figure 11. 3.3. Analysing Testing Result of Kevlar 29 3.3.1. Total Deformation The Max. and Min. Total Deformation in Kevlar 29 is 0.19844 mm and 0 mm respectively shown in Figure 12. The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002255_tmag.2021.3057124-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002255_tmag.2021.3057124-Figure4-1.png", + "caption": "Fig. 4. Finite element model of turbine generator based on proposed modeling method.", + "texts": [ + " Furthermore, we actually employ the two-potential method using edge element [16], [17] to effectively model the stator coils. Though the basic formulation of the proposed method is not changed due to the two-potential method, the right-hand terms in (2) and (3) are modified (see appendix for more detail). Finally, the proposed method is also valid for various radial-gap-type electric machines, not only for turbine generators. III. VALIDITY CHECK OF PROPOSED METHOD A 900MVA-class turbine generator is modeled using the proposed modeling method. Fig. 4 shows the separated FE model of the generator. The rotor surface is precisely discretized into fine elements so that the distributions of eddy current density can accurately be computed. In the periodic-region model, the stator cores and air ducts are integrally modeled using the homogenization method [18]. In addition, the model length of periodic region, i.e., L P , is determined based on the typical cross-slot pitch and wedge length. The total number of elements in this model is about 2000 000" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001413_j.jsv.2008.02.056-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001413_j.jsv.2008.02.056-Figure5-1.png", + "caption": "Fig. 5. Meridional moment diagram for T 03.", + "texts": [ + " (23b) By applying the principle of virtual work, we find our expression for determining d2H1 as 2pR2H 0 2d2H1 \u00bc Z v 2 \u00f0H1\u00des\u00f0H 02\u00dedv: (24) By substituting Eqs. (20) and (23) into Eq. (24) and performing the integrations, one finds d2H1 \u00bc H1RR1 E cd Iz0 1 A \u00fe \u2018 sin2 a 6R2t . (25) To find the rotational displacement d3H1 due to the H1 loading, we apply the virtual loading T 03 at and in the direction of coordinate 3 before application of the H1 loading. The meridional moment diagram for the T 03 loading is shown in Fig. 5. The virtual meridional moment M 0 x due to the T 03 loading is seen from Fig. 5 and from the previous development and discussions relative to Eq. (15) to be M 0 x \u00bc T 03 y \u2018 \u00fe 1 2 R1 R . (26) The virtual stresses sx caused by the T 03 loading are given by Eq. (8) as sx\u00f0T 0 3\u00de \u00bc Mz0y 0 Iz0 \u00bc T 03R1 Iz0 \u00f0y cos a z sin a\u00de. (27a) ARTICLE IN PRESS T.A. Smith / Journal of Sound and Vibration 318 (2008) 428\u2013460440 The virtual stresses sy due to M 0 x (Fig. 5) may be found from the previous development and discussions relative to Eq. (15) to be sy\u00f0T 0 3\u00de \u00bc M 0 xz Ix \u00bc 12T 03z t3 y \u2018 \u00fe 1 2 R1 R . (27b) By invoking and applying the principle of virtual work, we find our equation for determining d3H1 to be 2pR1T 0 3d3H1 \u00bc Z v 2 \u00f0H1\u00des\u00f0T 03\u00dedv. (28) Upon substituting Eqs. (20) and (27) into Eq. (28) and performing the integrations, one finds d3H1 \u00bc H1RR1 EIz0 c\u00fe n\u20182sin a cos a 12R \u00fe nc\u2018 sin a 2R . (29) To determine the rotational displacement d4H1 caused by the H1 loading, we apply the virtual loading T 04 at and in the direction of coordinate 4 before applying the H1 loading" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002687_j.seta.2021.101240-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002687_j.seta.2021.101240-Figure8-1.png", + "caption": "Fig. 8. Flux lines Trends of Modified Pole- PMDC Motor using Infolytica MotorSolve Software.", + "texts": [ + " Moreover, this modified PMDC motor net magneto force will be reduced. The modified pole PMDC motor is designed with Infolytica MotorSolve Software is illustrated in Fig. 7. Flux lines Trends of Modified Pole PMDC Motor has been shown with the help of \u2018Infolytica MotorSolve Software. After cutting the pole magnets into pieces, it has been observed that, the concentration of flux lines are more towards the gap in magnets of main poles i.e. South Poles and North Poles of the Magnets assemblies. Flux lines of modified PMDC motor is shown in Fig. 8. An air gap of the flux density distribution is the first-rate concurrence, concurrently armature speed and RMS current is full of very high. Moreover, the position of the motor is increased also flux density value varies rapidly. An air\u2013gap flux line of modified Pole-PMDC motor is shown in Fig. 9. The position of the motor is increasing then the cogging torque is decreasing exponentially. Moreover, a comparison of cogging torque in basic PMDC motor and modified Pole-PMDC motor is exposed in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003330_s00707-021-03026-0-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003330_s00707-021-03026-0-Figure1-1.png", + "caption": "Fig. 1 A schematic illustration of the piezoelectric ribbon-substrate structure", + "texts": [ + " To make a brittle piezoelectric ribbon form a wavy configuration on a soft substrate, there are two strategies. One strategy is stretching the soft substrate, bonding the piezoelectric ribbon on the compliant substrate, and releasing the substrate. Then, the ribbon forms a wavy configuration on the soft substrate [30]. The other strategy is tuning the voltage applied to the piezoelectric ribbon, which will cause the piezoelectric ribbon to buckle on the soft substrate [31, 32]. In this paper, the second strategy is considered, as shown in Fig. 1. To gain a comprehensive understanding of the dynamic behaviour of the wavy ribbon-substrate structure subjected to Gaussian white noise excitation, an analytical model is required, which will be established in the following. For the wrinkled (also named as buckled) piezoelectric ribbon, the density of its kinetic energy is calculated by K 1 2 \u03c1 ( \u2202w3 \u2202\u03c4 )2 , (1) where \u03c1 indicates the density, \u03c4 is the time, and wi denotes the displacement of the piezoelectric ribbon in the corresponding xi coordinate. x1, x2, x3 represent the three perpendicular directions of the length, width, and thickness of the ribbon-substrate structure, respectively, as shown in Fig. 1. The total electric enthalpy density takes the form of 1 2 \u03c3i j\u03b5i j \u2212 Di Ei , (i, j 1, 2, 3), (2) where \u03c3i j and Di represent the stresses and the electric displacements; \u03b5i j and Ei are the strain tensors and the electric fields. For the wavy piezoelectric ribbon, the linear constitutive equations for stresses \u03c3i j and electric displacements Di of piezoelectric materials can be expressed as [32\u201334]\u23a7\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a9 \u03c3xx \u03c3yy \u03c3zz \u03c3yz \u03c3xz \u03c3xy \u23ab\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23ac \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23ad \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 c\u030411 c\u030412 c\u030413 0 0 0 c\u030412 c\u030411 c\u030413 0 0 0 c\u030413 c\u030413 c\u030433 0 0 0 0 0 0 c\u030444 0 0 0 0 0 0 c\u030444 0 0 0 0 0 0 (c\u030411 \u2212 c\u030412) / 2 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 \u23a7\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a9 \u03b5xx \u03b5yy \u03b5zz 2\u03b5yz 2\u03b5xz 2\u03b5xy \u23ab\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23ac \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23ad \u2212 \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 0 0 e\u030431 0 0 e\u030431 0 0 e\u030433 0 e\u030451 0 e\u030451 0 0 0 0 0 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 \u23a7\u23a8 \u23a9 Ex Ey Ez \u23ab\u23ac \u23ad, (3) \u23a7\u23a8 \u23a9 Dx Dy Dz \u23ab\u23ac \u23ad \u23a1 \u23a3 0 0 0 0 e\u030451 0 0 0 0 e\u030451 0 0 e\u030431 e\u030431 e\u030433 0 0 0 \u23a4 \u23a6 \u23a7\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a9 \u03b5xx \u03b5yy \u03b5zz 2\u03b5yz 2\u03b5xz 2\u03b5xy \u23ab\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23ac \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23ad + \u23a1 \u23a3 k\u030411 0 0 0 k\u030411 0 0 0 k\u030433 \u23a4 \u23a6 \u23a7\u23a8 \u23a9 Ex Ey Ez \u23ab\u23ac \u23ad, (4) where c\u0304i j , e\u0304i j and k\u0304i j are the reduced elastic, piezoelectric and dielectric coefficients [31], respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002530_cerma.2007.4367740-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002530_cerma.2007.4367740-Figure6-1.png", + "caption": "Figure 6. Radio calculation of the curve trajectory followed to circumvent an obstacle", + "texts": [ + " In order to avoid a dead-locking between the tractor and the obstacles we will use an influence distance calculated in function of the number of trailers. This consideration can be explained from the fact that the robot needs a larger place to turn around the obstacles depending on the number of modules that it tows. The influence\u2019s distance used by the algorithm must be equal to the radius ri, where i stands for the number of modules. The robot must begin to circumvent the obstacle from this distance. Figure 6 shows how this influence distance is obtained. The anti-collision constraints on the trailers must be mapped to the tractor\u2019s velocity space to prevent the collisions. There exist some parameters which must be identified: 1. The tow where the constraint occurs must be identified. 2. The constraints\u2019 passage matrix is generated. 3. The constraint produced on the velocity of the trailer must be relocated within the tractor\u2019s velocity space. Once the constraint generated on the tow is mapped to the tractor\u2019s velocity space, actions must be undertaken to avoid the collision" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000790_s1560354708040072-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000790_s1560354708040072-Figure2-1.png", + "caption": "Fig. 2. The center of mass of a hexagonal pencil falls through height a sin \u03b8 during each 1/6 turn, where a is the width of a face of the pencil. For an N-sided pencil whose faces have width b, the center of mass falls by b sin \u03b8 during each 1/N turn.", + "texts": [ + " During each 1/6 turn, \u03b8 \u2212 \u03c0/6 \u03b1 \u03b8 + \u03c0/6. (3) The equation of motion (2) is a so-called Mathieu equation, which does not lend itself to analytic solution. Instead, we use an energy analysis to estimate the asymptotic linear and angular velocity of the rolling pencil. 1)Typical hexagonal pencils have rounded edges such that spontaneous rolling commences for \u03b8 \u2248 \u03c0/9. REGULAR AND CHAOTIC DYNAMICS Vol. 13 No. 4 2008 During each 1/6 turn, the center of mass of the pencil falls by height a sin \u03b8, as shown in Fig. 2. Thus, gravity adds energy \u0394E = mga sin \u03b8 (4) to the pencil each 1/6 turn. At the end of each 1/6 turn, the pencil collides with the surface of the inclined plane. Let \u03b5 (0 \u03b5 1) be the coefficient of (in)elasticity for this collision, meaning that if the pencil has kinetic energy E just before a collision, it has kinetic energy \u03b5E just after the collision. The parameter \u03b5 is essentially the square of the coefficient of restitution defined for one-dimensional collisions. Then, at the beginning of the second 1/6 turn, the pencil has kinetic energy E2 = \u03b5\u0394E, assuming that the pencil starts from rest", + " However, this assumption is inconsistent with conservation of angular momentum about the edge newly in contact with the plane (Eqs. (32)-(33)), and also inconsistent with the torque analysis (36) about the axis of the pencil. The kinetic energy lost in the collision is given by \u0394E = 1 2 I(\u03c92 \u2212 \u03c9\u20322) = kmv2 2 ( 1 \u2212 \u03c9\u20322 \u03c92 ) , (37) where v is the velocity of the center of mass of the pencil just before the collision. In the large-N limit where \u03c9\u2032 = \u03c9, no energy is lost during the collision. During each 1/N turn the center of mass of the pencil falls by height h = b sin \u03b8, recalling Fig. 2, and gravitational potential energy mgh is converted into kinetic energy. The asymptotic condition is that the potential energy gained during each 1/N turn equals the kinetic energy lost during the collision at the end of that turn. REGULAR AND CHAOTIC DYNAMICS Vol. 13 No. 4 2008 The asymptotic velocity ve of the center of mass at the end of a 1/N turn is given by mgb sin \u03b8 = \u0394E = kmv2 e 2 ( 1 \u2212 \u03c9\u20322 \u03c92 ) . (38) or v2 e = 4ag k sin \u03c0 N 1 \u2212 \u03c9\u20322/\u03c92 sin \u03b8. (39) Since \u0394E \u2192 0 in the large-N limit, there is no finite asymptotic velocity for any nonzero value of the inclination \u03b8 in this limit" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001109_imece2007-42921-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001109_imece2007-42921-Figure6-1.png", + "caption": "Figure 6. Stress field (Von Mises) for the second layer deposition at t = 9 s", + "texts": [ + " Since the heat losses at these regions are less than middle segments and first layer, this result in higher temperature at these points. Therefore, the clad heights as shown in Figure 3 and the stresses increase at the end-point of the upper layers. Te relationship between the temperature distribution and geometry of the deposited material in multilayer LSFF are studied in detail by Alimardani et al. [8]. Because of the higher temperature, and consequently higher thermal stress at the end segments, these locations are more likely subject to the delamination and crack formation as shown in Figure 6. To verify this result obtained from numerical model, the microscopic views of the second layer at three different cross-sections normal to the x axis (as schematically shown in Figure 3) are presented in Figure 7. As shown in Figure 5, the maximum stresses vary along the deposition tracks and within the same layer. The maximum stresses are larger at the beginning and end points of layers 2 to 4, while it was lower at the middle section. The magnitude of the maximum stress is almost half at the middle section, as compared to beginning and end segments", + " In addition, the number of micro-cracks per unit width 6 Copyright \u00a9 2007 by ASME erms of Use: http://www.asme.org/about-asme/terms-of-use is almost half at the middle section, as compared to the beginning and the end sections. This fact can be predicted based on the numerical results illustrated in Figure 5. Maximum stresses for the first layer are two times larger for this layer compared to layers 2 to 4, and also the high stresses at the two end segments of the wall result in more cracks at these regions. This can also be seen in Figure 6. Microstructure plays a very crucial role in determining the property of a fabricated component using LSFF. In the present work, the results of 3D modeling are further extended to study the microstructure of the component. A detailed observation of the optical and scanning electron micrographs at various crosssections of the fabricated thin wall were undertaken and presented in Figure 8. As seen in this figure, the microstructure of the thin wall consists of fine dendrites, cellular and also mixture of both" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000828_3.43615-Figure10-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000828_3.43615-Figure10-1.png", + "caption": "Fig. 10. Components of the loop.", + "texts": [ + "4 36 15 eluding a centrifugal force WVq acting on the vehicle at the bottom of the loop. The terms q and R are the angular velocity in pitch and turning radius, respectively. The lift being developed is nW, where n is the load factor. Consider a and d to be made up of two parts, the amount required for trim flight and the amount required for the maneuver. Then L = 5man) + Lqq (3a) Li = -L/tri and M = ^ Ms(8trim + <5man) or M = + (3b) Mqq (3c) (3d) The condition shown in Fig. 9 is thus a sum of the two conditions shown in Fig. 10, the level flight trim at weight W, plus the \"trim\" to overcome the centrifugal force. The aerodynamic force and moment in the second condition shown in Fig. 10 also depend on the rotational velocity q. The maneuverability can then be studied on the basis of the system shown in the right-hand sketch of Fig. 10. The left-hand side was investigated in the previous section of this paper. The force and moment system in accordance with the procedure outlined in this paper is transferred to the neutral point. The nondimensional force system, equivalent to Fig. 4, is shown in Fig. 11. The angular velocity with the cap over it q is a nondimensional angular velocity. The freestream velocity VQ and a characteristic length I are used for the nondimensionalizing. The term /x is the nondimensional relative mass parameter", + " The system will therefore be stable if the square root is imaginary or less than 1 .0. Therefore, for stability, 0 (8) Thus, the fact whether the vehicle is stable or not (disregarding the type of transient) is dependent on the force and moment system due to the angle of attack of the body a and the pitching velocity q. The relationships of Eq. (8) obtained for stability suggest that the behavior of the vehicle at the bottom of a loop of constant radius will indicate whether or not it is stable (see Fig. 9). The force system of the right-hand side of Fig. 10 for the additional forces arising from the steady turn is repeated in Fig. 12. The value of the moment about the neutral point, noted as CM' in Fig. 11, is now written in terms of the moment and force about the e.g. In addition, the moment due to the control surface is written in terms of the control force acting at some distance IT from the e.g. Consider the system of Fig. 12 in equilibrium prior to instantaneously zeroing the maneuver value of the control surface deflection. If the moment about the neutral point is negative, then making 5man zero will cause the moment to bring the vehicle out of the loop" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002832_j.matpr.2021.05.047-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002832_j.matpr.2021.05.047-Figure6-1.png", + "caption": "Fig. 6. Deformation analysis at crank position.", + "texts": [], + "surrounding_texts": [ + "The initial Strandbeests developed by 71 year old Dutch artist was made using plastic pipes, nylon threads and adhesive tapes [5]. In this case, the Dutch artist Theo Jansen made use of solution which is economical and low technology that is interpretable to common people around for its use. Jansen has covered the entire Strandbeests with plastic bottles provided with switch that supplies adequate amount of air in and out of the system to have movement of this Strandbeests. But in the specified proposal Fig. 7. Elastic Strain Analysis at crankshaft position. Fig. 8. Principal Stress analysis on crankshaft due mountings\u2019 load effect. system, the material used for legs of motion is made of such a material that has stress and strain bearing capacity proportional to load mounted onto it and its self-weight. The below Fig. 10 shows a system that is desired to get implemented for agricultural applications. As mentioned earlier system comprises of certain number of mountings, and more importantly the system is balanced upon number of legs, linkages provided as per the cost and design aspects. Each linkage is provided with more than one degree of freedom for efficient working conditions [2]. In the above figure, the numbers represent following: 1. Plough with spike mechanism 2. Camera systems 3. Antenna for control of entire system 4. Fertilizers 5. Seed sowing arrangement 6. Solar panels 7. Windrower/Mower 8. Sensors 9. Secondary storage The system is provided with primary units and secondary units that act as power house for the system performance. The primary units form the solar panels that are provided at the edges of board or plate and secondary unit is the secondary storage provided in the above figure. This secondary storage can either be made use for placing screw turbine that can generate power for system in still water or any other power generating means that are under estimated economy and easy to replace for any other storage purposes. The vertical half portion of the system (as seen by top view) consists of transmission head and seed separator machines. The data transmission used can be facilitated with antennas so as to have no intermittent operations that may eventually leads to time loss. The seed separator dust generated is put down on agricultural fields such that it gets uniformly blended with the soil in short period of time interval and they are also provisions provided for seed sowing. The spiked wheels provided have an actuation that provides operation of soil loosing and its related actions for desired time which is controlled by use of controllers and wireless technology helps the farmers to control and operate from anywhere, In order to have visual inspections of the operations being performed Fig. 9. Different ways of obtaining locomotion on agricultural fields using Theo Jansen mechanism [4]. upto the mark of satisfactions camera sub system has been implemented at both the ends of the system. All the legs are connected to each other through two boards made up of suitable material such as steel, composites, etc. which have high load bearing capacity so that it has enough strength to hold the mountings. The mountings include the power systems and its mechanisms, antenna systems, spiked wheels, secondary storage, solar panels, and controllers. Additionally the system can include camera or video sub systems so as to provide the farmers knowledge and information about the detailed portions of agricul- tural fields by being anywhere through use of communication technologies such as Wi-Fi technologies or IoT (Internet of Technology) applications to make the system user friendly for farmers. Apart from these mountings the other basic mountings includes the operational devices such as fertilizer spraying, weed removing, seed sowing and other various activities performed by the farmer manually or through utilization of tractors and domestic animals. To transmit power from motor to linkages to provide motion utilization of belt drives or chains can be done [1]. The power being received from motor is distributed through gears and shafts to required sub systems such as those of all leg linkages and spiked wheels which is mounted to either bottom or side surface of one of the board or through the slots provided on the board and this depends on the space requirements, operational requirements and other design requirements for desired effectiveness and efficiency. The speed of response, the speed of motion of the system and steering or the rotation of system towards left or right of its line of motion is handled by the use of suitable electronic devices such as Pulse Width Modulation (PWM), use of time interrupt controls, microcontrollers, Bluetooth applications and transistors like MOSFETS [7]. The same proposed system is compared and hypothesized to adopt the similar simulation environment as mentioned in [6]. The simulation which would be taken has to be ensured with the algorithm that involves design specification as per desired needs and has to be scaled proportionally because it has been practiced that the digital simulation of Theo Jansen mechanism often fail due to difference in lengths of links of displayed and proposed linkage configurations. The algorithm implemented must have suitable user flexible factors such that it provides the standard mechanical parameters essential such as centre of gravity, centre of pressure as well as flexibility in making changes to the lengths of links to the required dimensions. The below Fig. 11 displays the demonstration of algorithm flexibility with respect to link lengths such that Fig. 11. Use of algorithm in approaching perfect path trace without multiple corrections by user [6]. desired link length is approached automatically once input initially by the user. The important thing to be observed is that these lengths of different links have synchronization with the orbits of locomotion in this geometry phase of construct. These digital simulation models can be of various size and shape that are in accordance with the Theo Jansen mechanism configurations. The claw type and rotation type [6] architectural configurations which give possible design benefits for the applications in agriculture fields can be built using GH systems. The popular digital models used to construct such a system include the Rhino and Grasshopper simulation softwares. Also the simulation models presently built have the option of dynamic analysis of linkages of multiple leg coordination in accomplishing the required task. Such a digital dynamic models provide scope of variety of parametric configuration\u2019s construction as shown in below Fig. 12 which indirectly describes the broad spectrum of different functions the configuration able to handle efficiently [6]." + ] + }, + { + "image_filename": "designv11_83_0000220_cbo9780511546877.006-Figure4.23-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000220_cbo9780511546877.006-Figure4.23-1.png", + "caption": "Figure 4.23: A single constraint was added to the point p on A2, as shown in (a). The curves in (b), (c), and (d) depict the variety for the cases of f 1 = 0, f 1 = 1/8, and f 1 = 1, respectively.", + "texts": [ + " Set the x coordinate to 0, which yields a1a2 \u2212 b1b2 + a1 = 0, (4.65) Cambridge Books Online \u00a9 Cambridge University Press, 2009use, available at https://www.cambridge.org/core/terms. https://do .org/10.1017/CBO9780511546877.006 Downloaded from https://www.cambridge.org/core. The Librarian-Seeley Historical Library, on 01 Dec 2019 at 10:09:11, subject to the Cambridge Core terms of P1: JZP book CUNY490-Lavalle 0 521 86205 1 April 14, 2006 15:46 144 MOTION PLANNING and allow any possible value for y. As shown in Figure 4.23a, the point p must follow the y-axis. (This is equivalent to a three-bar linkage that can be constructed by making a third joint that is prismatic and forced to stay along the y-axis.) Figure 4.23b shows the resulting variety V (a1a2 \u2212 b1b2 + a1) but plotted in \u03b81 \u2212 \u03b82 coordinates to reduce the dimension from 4 to 2 for visualization purposes. To correctly interpret the figures in Figure 4.23, recall that the topology is S 1 \u00d7 S 1, which means that the top and bottom are identified, and also the sides are identified. The center of Figure 4.23b, which corresponds to (\u03b81, \u03b82) = (\u03c0, \u03c0), prevents the variety from being a manifold. The resulting space is actually homeomorphic to two circles that touch at a point. Thus, even with such a simple example, the nice manifold structure may disappear. Observe that at (\u03c0, \u03c0) the links are completely overlapped, and the point p ofA2 is placed at (0, 0) inW . The horizontal line in Figure 4.23b corresponds to keeping the two links overlapping and swinging them around together by varying \u03b81. The diagonal lines correspond to moving along configurations such as the one shown in Figure 4.23a. Note that the links and the y-axis always form an isosceles triangle, which can be used to show that the solution set is any pair of angles, \u03b81, \u03b82 for which \u03b82 = \u03c0 \u2212 \u03b81. This is the reason why the diagonal curves in Figure 4.23b are linear. Figures 4.23c and 4.23d show the varieties for the constraints a1a2 \u2212 b1b2 + a1 = 1 8 , (4.66) Cambridge Books Online \u00a9 Cambridge University Press, 2009use, available at https://www.cambridge.org/core/terms. https://do .org/10.1017/CBO9780511546877.006 Downloaded from https://www.cambridge.org/core. The Librarian-Seeley Historical Library, on 01 Dec 2019 at 10:09:11, subject to the Cambridge Core terms of P1: JZP book CUNY490-Lavalle 0 521 86205 1 April 14, 2006 15:46 THE CONFIGURATION SPACE 145 and a1a2 \u2212 b1b2 + a1 = 1, (4.67) respectively. In these cases, the point (0, 1) in A2 must follow the x = 1/8 and x = 1 axes, respectively. The varieties are manifolds, which are homeomorphic to S 1. The sequence from Figure 4.23b to 4.23d can be imagined as part of an animation in which the variety shrinks into a small circle. Eventually, it shrinks to a point for the case a1a2 \u2212 b1b2 + a1 = 2, because the only solution is when \u03b81 = \u03b82 = 0. Beyond this, the variety is the empty set because there are no solutions. Thus, by allowing one constraint to vary, four different topologies are obtained: 1) two circles joined at a point, 2) a circle, 3) a point, and 4) the empty set. Since visualization is still possible with one more dimension, suppose there are three links, A1, A2, and A3", + " The Librarian-Seeley Historical Library, on 01 Dec 2019 at 10:09:11, subject to the Cambridge Core terms of P1: JZP book CUNY490-Lavalle 0 521 86205 1 April 14, 2006 15:46 THE CONFIGURATION SPACE 147 We now describe a general methodology for defining the variety. Keeping the previous examples in mind will help in understanding the formulation. In the general case, each constraint can be thought of as a statement of the form: The ith coordinate of a point p \u2208 Aj needs to be held at the value x in the body frame of Ak . For the variety in Figure 4.23b, the first coordinate of a point p \u2208 A2 was held at the value 0 in W in the body frame of A1. The general form must also allow a point to be fixed with respect to the body frames of links other than A1; this did not occur for the example in Section 4.4.2 Suppose that n links, A1, . . . ,An, move in W = R 2 or W = R 3. One link, A1 for convenience, is designated as the root as defined in Section 3.4. Some links are attached in pairs to form joints. A linkage graph, G(V, E), is constructed from the links and joints" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001921_6.2007-1205-Figure11-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001921_6.2007-1205-Figure11-1.png", + "caption": "Figure 11. Proportional navigation guidance geometry.", + "texts": [ + " The above considerations will be given attention in future works. In this paper, to facilitate the development of the simulation environment, a simple proportional navigation law22 is implemented. Let \u03b3\u0307c be the 12 of 18 American Institute of Aeronautics and Astronautics commanded heading rate. Let \u03bb\u0307 be the line of sight (LOS) rate. The proportional navigation law is given by \u03b3\u0307c = Nc\u03bb\u0307 (26) where Nc is a designer-chosen constant navigation parameter. The guidance geometry used here is shown in Fig. 11. vp is the pursuer\u2019s velocity. \u03bb is calculated from the geometry in Fig. 11 using the known xp(t) and the xte(t) provided by the estimation module. \u03bb is passed through a first order filter and differentiated to obtain \u03bb\u0307, which is then multiplied by the guidance gain Nc. The guidance filter serves only to slightly smooth some of the sharpness of the LOS signal due to rapid changes in the target\u2019s estimated position. A saturation limit \u03b3\u0307cmax is placed on the commanded heading rate in order to impose a maximum turn rate on the pursuer. The block diagram of the guidance module is shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000312_978-3-540-44410-7_8-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000312_978-3-540-44410-7_8-Figure4-1.png", + "caption": "Fig. 4. Dual arm system supported by a vehicle.", + "texts": [ + " Moreover note that, in order to also avoid any possible interference between \u03b6\u0307 \u2014now smoothly generated via (29), (24) and z\u0307 which is maintained active via the smoothed form (14))\u2014 we should furtherly shape \u03b1(\u00b5), with respect to \u03b1(\u00b5), in such a way as to be certainly unitary within the whole finite support where \u03b1(\u00b5) > 0 (see Fig. 3). As it can be easily realized, with such a choice for \u03b1 and \u03b1, signal z\u0307 is always null, even during the smooth transition phase of \u03b8\u0307 (and then \u03b6\u0307). 4 Control of a Dual-Arm Nonholonomic Mobile Manipulator The results of the previous section are hereafter extended to the case of a dual-arm nonholonomic mobile manipulator of the type of Fig. 4, when performing grasping operations. As a matter of fact, a grasping operation to be performed by the overall system simply corresponds to the global task of having the two end-effector frames < e1 >, < e2 > asymptotically converging to the goal frames < g1 >, < g2 > respectively (Fig. 4), while obviously maintaining the desired minimum level of manipulability for each arm. By still assuming each arm to be separately controlled, as done in the previous section, let us start again by considering the following global candidate Lyapunov function, with an obvious meaning of the introduced terms, V := 1 2 (eT 1 e1 + eT 2 e2) (30) whose time derivative is V\u0307 = \u2212eT 1 (x\u03071 + X\u03071)\u2212 eT 2 (x\u03072 + X\u03072) := \u2212eT (x\u0307 + X\u0307). (31) Then, by performing the same analysis leading to (23) in the previous section, we get V\u0307 = \u2212eT (x\u0307\u2217 + S\u03b8\u0307) = \u2212eT ( \u02d9\u0302x + z\u0307 + S\u03b8\u0307) = \u2212eT [ (x\u0307 + \u03b6\u0307) + z\u0307 + S\u03b8\u0307 ] (32) where now x\u0307\u2217 := [ x\u0307\u2217T1 x\u0307\u2217T2 ]T S := [ST 1 ST 2 ]T \u02d9\u0302x := [ \u02d9\u0302x T 1 \u02d9\u0302x T 2 ]T z\u0307 := [ z\u0307T 1 z\u0307T 2 ]T (33) x\u0307 := [ x\u0307 T 1 x\u0307 T 2 ]T \u03b6\u0307 := [ \u03b6\u0307T 1 \u03b6\u0307T 2 ]T The (12\u00d7 6) matrix S is still of full column rank type, and the overall coordination signal \u03b6\u0307 is to be suitably chosen" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002135_s00202-020-01163-8-Figure13-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002135_s00202-020-01163-8-Figure13-1.png", + "caption": "Fig. 13 The axial cut of the proposed TSCMG", + "texts": [ + " Figure\u00a011 shows the output of the middle rotor belt. The important point to notice is the parameters of the middle rotor, which play a key role in the torque of the inner and outer rotors. That is why we can say that the middle rotor is the most critical and effective component of a multispeed magnetic gear, the magnetic field of which affects other rotors at the same time. Furthermore, Fig.\u00a012 illustrates the modulators and rotors of the proposed magnetic gear built in this work. An axial cut of the proposed TSCMG is graphed in Fig.\u00a013, which exhibits the operation mechanism of the gear thoroughly along with the arrangement of modulators, rotors and other appurtenances. Figure\u00a014 indicates the layout of the magnetic gear built in the laboratory, which includes three separate shafts on the inner, middle and outer rotors. The shaft of the middle rotor is rotated by the belt. This shaft, as per Fig.\u00a015, is connected to the test bench via the belt. The test bench includes one motor, three torque-meters, three speed-meters, one drive 1 3 and other elements that are measured point to point under different torque speeds" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001190_09544062jmes539-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001190_09544062jmes539-Figure3-1.png", + "caption": "Fig. 3 Three-dimensional distributions of pressure, film thickness, and temperatures in the middle layer of the film and on the plane of symmetry predicted by the steady-state thermal and non-Newtonian EHL solution for Ue = 2 \u00d7 10\u221211, G = 4949, W0 = 6.143 \u00d7 10\u22126, ke = 2.5, and \u03be = 1.0", + "texts": [ + "593 \u00b5m, respectively, suggesting that the surface deflection at the centre of the contact is 6.749 \u00b5m and this is consistent with the Hertzian solution, indicating a heavily load case. The predicted minimum film thickness is 0.375 \u00b5m. The three-dimensional distributions of the pressure and film thickness, the temperature in the middle layer of the film (i.e. the layer of z/h = 0.5), denoted by Tm, and the temperature on the plane of symmetry (i.e. the plane of y = 0), denoted by Tsymmetry, are given in Fig. 3, meanwhile, the profiles of pressure and film thickness and the contour map of the temperature on the plane of symmetry are plotted in Fig. 4. For a vibration amplitude A = 6 \u00b5m, similar to the undeformed reference centre distance, and a vibration frequency f = 1000 Hz, time-dependent cyclic solution was obtained taking both the thermal and non-Newtonian effects into account. At the four Proc. IMechE Vol. 221 Part C: J. Mechanical Engineering Science JMES539 \u00a9 IMechE 2007 at UNIV OF ILLINOIS URBANA on March 16, 2015pic" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002317_tac.2021.3058960-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002317_tac.2021.3058960-Figure3-1.png", + "caption": "Fig. 3: Planar arrangement of orbits on the grid where neighboring feature opposite rotation.", + "texts": [ + " See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. an orbit; Fig. 2) around each center ( s \u03c0 `i, s \u03c0 mi) through the equation cos ( \u03c0 s \u03c1i cos \u03b8i + `i ) cos ( \u03c0 s \u03c1i sin \u03b8i + mi ) = C (1) where C \u2208 R is a circulation parameter. Now fix |C| in (1) to some positive constant, and for i = 1, . . . , N set the right hand side of (1) to either |C| or \u2212|C| so that neighboring (in the north/south/east/west direction) frames are assigned opposite values (Fig. 3). Motivated by the geophysical gyre flow model of [15], [22], [47] and setting a positive amplitude orbit parameter A > 0, the phase dynamics of an agent moving along an orbit are defined as follows. Definition 4. Given an initial position on an orbit with circulation parameter C, the phase dynamics of an agent is in the following form \u03b8\u0307i = \u2212sgn(C)\u03c0A \u03c1i [ cos \u03b8i sin(\u03c0\u03c1i s cos \u03b8i) cos(\u03c0\u03c1i s sin \u03b8i) + sin \u03b8i sin(\u03c0\u03c1i s sin \u03b8i) cos(\u03c0\u03c1i s cos \u03b8i) ] , fC(\u03b8i) , (2) with T = 8 \u03c0A \u222b \u03c0/2 s\u03c0\u22121 arcsin \u221a |C| (sin \u03c0x s \u2212 C2)\u22121/2 dx as a period of oscillation", + " Now define v , max ` \u2211 (i,j)\u2208E` W\u03c0(\u2206\u03d1ij) computed along one of the paths from the root to a leaf. That particular leaf node given by argmax` \u2211(i,j)\u2208E` W\u03c0(\u2206\u03d1ij) is identified as the network\u2019s trailing leader, with phase image denoted \u03d1min. Given the shape of the gyre orbits (1) for |C| \u2192 0 and the fact that \u03b4 s, the proximity condition for rendezvous is satisfied for two neighboring agents drifting along orbits i and j in two distinct scenarios: (i) away from the saddle points of the ambient flow (Fig. 3), where for some small \u03b5 > 0, | \u03d1i+\u03d1j 2 \u2212 k\u03c0 2 | < \u03b5, k = 0,\u00b11,\u00b12, or (ii) near the saddle points of the ambient flow, where | \u03d1i+\u03d1j 2 \u2212 k\u03c0 4 | < \u03b5, k = \u00b11,\u00b13. In case (i) exactly two agents are in rendezvous within each ball of radius \u03b4, whereas in case (ii) up to four agents can be in rendezvous with each other (Fig. 7). The control objective is stated as follows. Problem 1. limt\u2192\u221e(\u03d1i \u2212 \u03d1j) = 0, \u2200i, j \u2208 {1, . . . , N}. Figure 5 illustrates the transition that the desired control intervention aims to achieve" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003527_s00707-021-03025-1-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003527_s00707-021-03025-1-Figure5-1.png", + "caption": "Fig. 5 For any nonhorizontal and nonstationary slopes l, they can be converted from horizontal one l2 by translation and rotation. The center of rotation is located at pc. ps1 is the interim point translated from landing point ps and rotated by [\u03b8x , \u03b8y , \u03b8z] to get ps2. Similarly, pf , pf 1, and pf 2 are the contact point locations of the robot before and after translation or rotation, respectively", + "texts": [ + " Different scenes are defined for verifying the feasibility and effectiveness of the ground contact model and balance controller. There is a total of fifty contact points under each toe and heel. And all the simulations are begun at the double support. Moreover, the environment such as the directions, magnitudes, and the locations of perturbations and types of the ground are unknown to the controller. Besides, only the joint angles, joint velocities, joint torques, and GRF at each sole can be measured by sensors (Fig. 5). 5.1 Maintaining balance on the stationary horizontal beam Firstly, we test the robot while standing on a stable horizontal beam. Figure 6 shows a series of results when the robot encounters an instantaneous perturbation in the sagittal plane. The external force (150 N) is applied at 0.1 s, locates at the CoM and lasts 0.1 s. The desired position of CoM is where the robot would be standing upright, and the desired linear and angular momenta are set to zero. The gravity of the robot is equivalent to a step response for the ground contact model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002679_14644207211012726-Figure9-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002679_14644207211012726-Figure9-1.png", + "caption": "Figure 9. Deviation from the nominal geometry as a function of strut angle for Opt setting; two samples are analyzed. BF: bed fusion; MM: maximum material; LM: least material.", + "texts": [ + " On the right of Figure 8, the comparison between LEVD curves of Ref and Opt specimens is reported. The correlation factors R2 for the distribution fitting are listed as well. A good fit was achieved, with R2 greater than 0.94 in all cases. In the Opt case, an anomalous behavior of the fitted distribution is observed, as the 15 curve is above the 0 one. Comparing the two specimens, the Opt porosity is lower than the Ref, especially in the right portion of the plot, i.e., where pores are bigger and more detrimental to mechanical performance. The results of the dimensional check are shown in Figure 9, for two Opt samples. The contour map of the variation with respect to the nominal geometry reveals that the overall geometry is not distorted or bent globally, but imperfections are clearly visible (maximum deviations are approximately 0.2 mm) and seem to be mainly due to over-melting and parasitic particles in the down-skin surfaces, especially in 0 and 15 struts. BF, MM, and LM diameters relative to Opt samples were plotted as a function of the strut angle. Two spider-web specimens were inspected in order to get eight measurements per strut angle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002776_03093247211018820-Figure9-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002776_03093247211018820-Figure9-1.png", + "caption": "Figure 9. Dynamics model of lumped mass method.", + "texts": [ + " It should be pointed out that VFIFE method has poor computing speed in low-speed simulation problems, which takes about 20 h to finish the simulation. Since it is difficult for FEM to deal with high-speed rotational analysis (supercomputing center maybe needed), this paper compared simulation results between VFIFE and LMM to verify accuracy and stability of the proposed method under 15,000 r/min rotational speed. Simulation model and boundary conditions To ensure comparability between different methods and models, LMM model is defined as purely torsional concentrated parameter model, as shown in Figure 9. Input excitations of LMM are obtained by static FEM analysis, including transmission error, meshing stiffness, and contact force etc. Because of a large number of literature studied on spiral bevel gear dynamics based on LMM,34\u201336 dynamic equations are not detailed in this article. Due to ignorance of structural feature, LMM could not derive stress data or contact pattern. Therefore, contact force, displacement and acceleration are summarized to illustrate dynamic simulation results. Contact force" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001413_j.jsv.2008.02.056-Figure12-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001413_j.jsv.2008.02.056-Figure12-1.png", + "caption": "Fig. 12. Typical axisymmetric shell structure and loading. HB1 and HB2 are the radial boundary forces per unit circumferential length. TB3 and TB4 are the meridional moments per unit circumferential length at the shell boundaries.", + "texts": [ + " (53a), and proceeding in the same manner as used to find the free ring displacements with axial restraint at R2, we find the displacements d10, d20, d30, d40, and dV0 under pressure and axial loadings with axial restraint at R1 to be d10 \u00bc T01R2c EIz0 HpRR3 EA nV p1R1c\u2018 2 cos2 a 12EIz0 \u00fe n cos a 2Et \u00f02LR1 Vp1R1\u00de \u00fe nT01R\u20182 sin a cos a 12EIz0 nHpR3 sin a 2Et \u00fe \u2018 sin a cos a 6EtR \u00f03LR1 2V p1R1\u00de , (80) d20 \u00bc T01R2d EIz0 HpRR3 EA \u00fe nVp1R1d\u2018 2 cos2 a 12EIz0 \u00fe n cos a 2Et \u00f02LR1 V p1R1\u00de nT01R\u20182 sin a cos a 12EIz0 \u00fe nHpR3 sin a 2Et \u2018 sin a cos a 6EtR \u00f03LR1 V p1R1\u00de , (81) d30 \u00bc T01 R2 EIz0 \u00fe nR\u2018 sin a 2EIz0 \u00fe nVp1R1\u2018 2 cos2 a 12EIz0 , (82) d40 \u00bc T01 R2 EIz0 nR\u2018 sin a 2EIz0 nV p1R1\u2018 2 cos2 a 12EIz0 , (83) dV0 \u00bc T01R2\u00f0R2 R1\u00de EIz0 nVp1\u00f0R2 R1\u00deR1\u2018 2 cos2 a 12EIz0 \u2018 cos2 a 2ERT \u00f02LR1 V p1R1\u00de \u00fe nHpR3 cos a Et . (84) By using Eqs. (80)\u2013(84) in lieu of Eqs. (69)\u2013(73) for free ring displacements under combined pressure and axial loadings with axial restraint at R1, one will find the total free ring displacements as given already in Eqs. (74)\u2013(76). We show a typical shell structure and loading in Fig. 12. The boundary forces HB1, TB3, HB2, and TB4 represent prescribed boundary forces per unit circumferential length at each boundary. The thickness of the shell and the applied pressure loading p may vary along the shell meridian. The analysis will be performed by the standard flexibility method of indeterminate structural analysis. Our matrix form of the continuity equations for this solution is aF \u00feD0 \u00bc 0, (85) where a is the flexibility matrix for the base structure, F is the column matrix of redundant forces to be determined, and D0 is the column matrix of displacements under external loadings and temperature changes for F \u00bc 0", + " Our revised expressions for individual ring element axial deformations due to unit total circumferential values of the forces in the coordinate directions 1, 2, 3, and 4 are dV1 \u00bc d1V \u00bc R 2pE c\u00f0R2 R1\u00de Iz0 \u00fe nl2\u00f0R2 R1\u00de sin a cos a 12RIz0 l sin a cos a 2R2t n cos a Rt , (87a) dV2 \u00bc d2V \u00bc R 2pE d\u00f0R2 R1\u00de Iz0 \u00fe n\u20182\u00f0R2 R1\u00de sin a cos a 12RIz0 \u00fe \u2018 sin a cos a 2R2t n cos a Rt , (87b) dV3 \u00bc dVv \u00bc R\u00f0R2 R1\u00de 2pEIz0 1\u00fe n\u2018 sin a 2R , (87c) dV4 \u00bc d4V \u00bc R\u00f0R2 R1\u00de 2pEIz0 1 n\u2018 sin a 2R . (87d) For the analysis of the structure, the base structure will be obtained by dividing the structure into a series of conical ring elements as indicated in Fig. 12. The separate ring elements will be numbered as indicated in Fig. 13, and the coordinate system for the redundant radial forces and moments will be selected as shown in Fig. 13. The column matrix of unknown forces and moments can now be represented as \u00bdF \u00bc \u00bdP1;P2; . . . ;PN 1;MN ;MN\u00fe1; . . . ;M2N 3;M2N 2 T. (88) The column matrix of displacements for the base structure, with the redundants removed, can be represented as \u00bdD0 \u00bc \u00bdD10;D20; . . . ;DN 1;0;DN ;0;DN\u00fe1;0; . . . ;D2N 3;0;D2N 2;0 T. (89) The flexibility matrix for the structure is represented as \u00bda \u00bc \u00bdaij ; i \u00bc 1; " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001851_978-3-540-74027-8_1-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001851_978-3-540-74027-8_1-Figure4-1.png", + "caption": "Fig. 4. Sun and planet wheel modeling robot Viennese Waltz behaviour", + "texts": [ + " So also does the study of motion behaviour patterns and motion strategies of players in a premier league football match. It will be shown in this paper that the Viennese waltz behaviour pattern is a highly useful motion, possibly a fundamental motion, for omnidirectional robots. For example the behaviour pattern can be used to program a robot to back out of a tight corner whilst rotating, or, for a robot football player to circulate around an opponent whilst tracking the translating opponent. The classic \u201csun and planet wheel\u201d, Fig. 4, is used to model Viennese Waltz behaviour. The planet wheel is a disc attached to the robot. There is no slip between the planet and the sun wheels. The planet wheel could be on the outside of the sun wheel, Fig. 5, in which case the rotation of the planet wheel is reversed. The robot is shown, Fig. 6, backing out of a corridor representing retreating motion, i. e. translating, and simultaneously rotating such that the head of the robot is ready to face an opponent in the direction of its exit path" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001537_09544062jmes814-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001537_09544062jmes814-Figure1-1.png", + "caption": "Fig. 1 Model of a flexible rotor supported by two couple-stress fluid film journal bearings", + "texts": [ + " Owing to the non-linearity of the couple stress fluid film forces and to simplify the computations, this study assumes the journal bearing to be a so-called \u2018short bearing\u2019. The dynamic equations are solved using the Runge\u2013Kutta method. The dynamic trajectories of the rotor and bearing centre, power Proc. IMechE Vol. 222 Part C: J. Mechanical Engineering Science JMES814 # IMechE 2008 at UNIV OF CONNECTICUT on May 19, 2015pic.sagepub.comDownloaded from spectra, Poincare\u0301 maps, bifurcation diagrams, and the Lyapunov exponent are applied to analyse the rotor\u2013bearing system. 2 MATHEMATICAL MODELLING Figure 1 shows a flexible rotor supported by two couple-stress fluid film journal bearings on a parallel with non-linear spring. Om is the centre of the rotor gravity, O1 the geometric centre of the bearing, O2 the geometric centre of the rotor, and O3 the geometric centre of the journal. Figure 2 shows the crosssection of the fluid film journal bearing, where (X, Y) is the fixed coordinate and (e, w) is the rotated coordinate, e being the offset of the journal centre and w being the attitude angle of the X-coordinate" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002969_j.matpr.2021.05.300-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002969_j.matpr.2021.05.300-Figure7-1.png", + "caption": "Fig. 7. Al nitride clutch plate temperature results.", + "texts": [ + " Calculation of stresses, and greatest weights utilizing EulerLagrange conditions. Developing 3D model of MPCs utilizing Solid edge programming. Exporting the model of MPCs from strong edge to ANSYS 19.2 work bench for FEA. Static and dynamic examination of MPCs utilizing ANSYS work seat at various working condition. Determining the wear, most extreme disfigurement and equal worries for every material utilizing FEA Showing the connection of disfigurement and obstruction property of every material utilizing figures and tables (Fig. 1. Fig. 2. Fig. 3. Fig. 4. Fig. 5. Fig. 6. Fig. 7. Fig. 8. Fig. 9. Fig. 10. Fig. 11. Fig. 12. Table 1). Comparing and conversation of results for choice of better covering materials. Fig. 5. Cast iron clutch plate heat flux results. Fig. 6. Cast iron clutch plate heat flux results. Grip disappointment and harm because of extreme frictional warmth and warmth changes to the grasp counter mate circle frequently happens to a car grips. This circumstance adds to warm weakness to the segment which causes the grasp counter mate plate to break and twist" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001663_oceanskobe.2008.4530919-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001663_oceanskobe.2008.4530919-Figure2-1.png", + "caption": "Fig. 2. n-link underwater robot model", + "texts": [ + " In this paper, to obtain higher control performance we propose a disturbance compensation control method based on the discrete-time RAC method [11]. The influence of hydrodynamic modeling errors is treated as a disturbance, and the disturbance is compensated by using a disturbance observer. To verify the effectiveness of the proposed method, experiments using the 2-link underwater robot are performed. The experimental results show that the proposed control method has a good control performance. II. MODELING [11] An UVMS model used in this paper is shown in Fig. 2. It has a vehicle (robot base) and a n-DOF manipulator. Symbols are defined as follows: n number of joint ZI inertial coordinate frame Zi link i coordinate frame (i = 0, 1, 2, , n; link 0 means vehicle) 'Rj : coordinate transformation matrix from Zj to i 978-1-4244-2126-8/08/$25.00 \u00a92008 IEEE Ej: j x j unit matrix The tilde operator stands for a cross product such that ra = r x a. All position and velocity vectors are defined with respect to XI. A. Kinematics In this subsection, based on the reference [14] kinematic and momentum equations are derived. First, from Fig. 2 a time derivative of the end-effector position vector Pe 1S n VPeVO(Pe -ro) + {ki(pe Pi)}Iji. (1) On the other hand, relationship between end-effector angular velocity and joint velocity is expressed with we =wo+E ki/i i=1 (2) Pe position vector of manipulator end-effector pi position vector of origin of Zi ri position vector of gravity center of link i vo linear velocity vector of origin of Eo Ve linear velocity vector of end-effector +bo roll-pitch-yaw attitude vector of Eo b roll-pitch-yaw attitude vector of end-effector wi angular velocity vector of Zi We angular velocity vector of end-effector Xi relative angle of joint i 0 relative joint angle vector (= [01q, On,I/T) ki unit vector indicating a rotational axis of joint i m mass of link i Ma, added mass matrix of link i with respect to Zi Ii inertia tensor of link i with respect to Zi Ia, added inertia tensor of link i with respect to Zi xo : position and attitude vector of Eo(= [rTO fT]T) xe position and attitude vector of end-effector(= [PT: feT]T) Xo linear and angular vector of E0(= [vT, woTIT) Xe linear and angular vector of end-effector(= IVeT w TIT) ag: position vector from joint i to gravity center of link i ab- : position vector from joint i to buoyancy center of link i 1i length of link i Di width of link i Vi volume of link i p fluid density Cd: drag coefficient of link i g gravitational acceleration vector From Eqs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000551_2007-01-1309-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000551_2007-01-1309-Figure8-1.png", + "caption": "Figure 8: Mould flow analysis of housing", + "texts": [], + "surrounding_texts": [ + "The control electronics for the actuator comprises a circuit board on which nearly all required components for the electrical functions are mounted. The actuator has five electrical external interfaces: Kl30, Kl31, CAN-High, CAN-Low, Aux Supply for powering the auxiliary motor. A 16-bit microcontroller is used. In addition to the internal watchdog of the microcontroller, an external watchdog is also employed for monitoring; no second microcontroller is used for monitoring purposes. The safety function is guaranteed through the interaction with other control devices at system level. An H-bridge is used to control the main motor. The Hbridge is supplemented with a measuring circuit for recording the H-bridge current. In a failure incident case, the auxiliary motor is controlled externally. After a failure incident, or for diagnostic purposes, the actuator also has to be able to control the auxiliary motor. In this case it will suffice if the auxiliary motor turns in one direction. The heat load resulting from electrical losses in a Shiftby-Wire application is rather small, since driving mode changes only rarely occur. Most losses occur in the voltage regulator unit and in the microcontroller. This aspect, in addition to the size of the actuator, allow operation at ambient temperatures of 120 \u00b0C , and even at 140 \u00b0C for a short time period; specific heat dissipation measures for the electronics are not necessary. But due to the temperature requirements, circuit board materials with augmented glass transition point are used. Moreover, through-hole solder connections are not used. Electrical contacting to the in-molded lead frame elements in the housing is done via press-fit connections. The entire circuit board is conformal coated as protection against the expected build up of grit and residue in the actuator as a result of the mechanic wear." + ] + }, + { + "image_filename": "designv11_83_0003164_s43236-021-00268-y-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003164_s43236-021-00268-y-Figure2-1.png", + "caption": "Fig. 2 Load cycle in a food mixer", + "texts": [ + " Food mixers are started with a load. Due to their intermittent operation, the load varies for consecutive cycles. The load is reduced when the food particles get crushed. The initial load torque is decided by the size and texture of the food in the mixer. The texture of the food and the speed of the mixer influence the slope at which the load torque falls with respect to time. Assuming that the losses alone are supplied at the end of the mixer operation, the load torque curves are drawn, as shown in Fig.\u00a02. This figure also 1 3 shows the intermittent operation of the mixer for two ON\u2013OFF cycles. An equation for describing Fig.\u00a02 is given in (1): With Tlo specified, which is the initial load at t0. The constant \u2018a\u2019 is chosen based on the initial size and texture of the food contents. When the mixer operates at a high speed, the food particles are crushed quickly and the load requirement is reduced within a short span of time. When the machine is switched OFF in the middle of operation, the load torque stays constant and starts falling when it is switched ON again. Curves 1 and 2 in Fig.\u00a02 show load curves for two different initial textures and sizes of food, assuming the machine speed is kept constant. Meanwhile, in an actual scenario, the machine speed increases when the load falls, which leads to a further drop in the load curves as shown by 1\u2032 and 2\u2032. The food gets crushed as a function of speed and time. It reaches a final level after which the food particles cannot be crushed any further. Therefore, the load torque is reduced to a minimum value, which remains constant until the mixer is turned off" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002191_j.matpr.2020.12.112-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002191_j.matpr.2020.12.112-Figure4-1.png", + "caption": "Fig. 4. Force and Displacement.", + "texts": [ + " This point bargains about the compelled part appearing of the posts and imitating them by contributing specific physical representations and limit situations to reenact the fascinating situation plans ended likely. The segment stayed demonstrated utilizing ANSYS 14.0 software Plan Modeler programming as a strong perfect and ANSYS 14.0 worktable was utilized aimed at the assess- ment in the static examination. The key work method of a Limited Component Examination system is tended to like in Fig. 1(A) Fig. 2 Fig. 3 Fig. 4 Fig. 5 Fig. 6 Fig. 7 Fig. 8 Fig. 9 Fig. 10. The stainless harden fabric property Bulk (q): 7950 kg/m3 Young\u2019s modulus (E):206000Mpa Poisson\u2019s ratio: 0.3 The CAD typical of the bar is fit hooked on a limited quantity of components utilizing ANSYS 14.0 software inherent lattice calculation. The pillar contact locale is fit and interlinked to empower estimation of the power collaboration between them limited component investigation or FEA representing to a genuine task as a \u2018\u2018work\u201d a progression of little, consistently formed tetrahedron associated components, as appeared in the above fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002020_acemp.2007.4510512-Figure24-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002020_acemp.2007.4510512-Figure24-1.png", + "caption": "Fig. 24 Distribution of magnetic power loss.", + "texts": [ + " The gap field of this motor has the 1.5 times of he magnetization of permanent magnet. Figure 21 shows the magnetic flux distribution. From this result it was found that the magnetic path becomes short, therefore the inductance increases as this motor. Figs. 22 and 23 show the distribution of the magnetic field strength and the magnetic flux density, respectively. As same as procedure of the vector magnetic characteristic analysis, the distribution of magnetic power loss is obtained as shown in Fig. 24. The magnetic power loss increases in the place that both vectors B and H are large and the axis ratio in rotational magnetic flux is large. It is found that the major magnetic power loss is being generated at the tooth division of the stator core. 5 Conclusions In this paper we proposed the dynamic integrated typed E&S model for analyzing the eddy current effect under distorted magnetic flux conditions. And furthermore it was clarified how the effect of eddy current influence the vector magnetic characteristic" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001415_kem.353-358.2479-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001415_kem.353-358.2479-Figure2-1.png", + "caption": "Fig. 2 The finite element model of the rotor system", + "texts": [ + " As shown in Fig.1(b). Pn is normal impact-rub force, Pt is tangential impact-rub force. \u03c9 is the rotor rotating speed. Assuming that kr is impact-rub stiffness, e is the rotor eddy radial displacement and e 2 =x 2 +y 2 , \u03b4r is impact-rub clearance. f is tangential friction coefficient. When the rotor-stator impact-rub occurs, the impact-rub force is showed as follows: Pn=kr(e-\u03b4r) (2) Pt=f Pn Rotor system mathematical model. The finite element models of rotor-bearing system shown in Fig. 1 is shown in Fig. 2. Here, the small points denote the nodes, the numbers denote the node mark, the large black points express the two disk concentration quality. The entire rotor system movement equation can be expressed as: QKqqDqM =++ &&& (3) Here, M is the mass matrix including shaft and disk, D is the damped matrix including bearings damping, gyro movement and the internal inertia of axis, K is the whole stiffness matrix including looseness equivalent stiffness, bearings stiffness, impact-rub stiffness, etc. Q is a vector of the combine external forces, q&\u3001 q&& are generalized speed and acceleration vector" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000886_pime_conf_1965_180_339_02-Figure10.1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000886_pime_conf_1965_180_339_02-Figure10.1-1.png", + "caption": "Fig. 10.1. Use of a load-cell", + "texts": [ + " Strain-gauges of a resistance type enable compact load-cells to be provided which can be inserted between the bearing and the main frame of the equipment so that the load can be continuously monitored. Thus, in the case of a three-high mill, continuous wear of the lower roll bearing was being experienced and, after the material composition had been checked and found to be satisfactory, it was decided to investigate whether or not there were unexpected features regarding the load using a load-cell, Fig. 10.1. In the absence of a proprietary Vol180 Pt 3K at UNIV CALIFORNIA SANTA BARBARA on July 22, 2015pcp.sagepub.comDownloaded from 278 F. T. BARWELL AND D. SCOTT cell of the required capacity, one was constructed in the laboratory workshop especially for the trials so as to fit into the restricted space available between the lower bearing housing and the screw. This resulted in an undesirably low height-diameter ratio of the load-cell which had to be accepted. Four resistance-wire strain-gauges were bonded to the outer surface of the load-cell in diametrically opposed pairs and connected to a bridge circuit" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002794_5.0049922-Figure13-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002794_5.0049922-Figure13-1.png", + "caption": "FIGURE 13. Deformed shape of the glass-epoxy composite driveshaft at the critical speed.", + "texts": [ + " The black line is indicating the ratio of whirl frequency and rotational velocity. The red triangular marker are the intersection between black line and each whirl mode at which the rotating shaft resonates with the peak whirl frequency [14]. The rotational velocity corresponding to this intersection is the critical speed. The first forward and second backward critical speeds are 16319 rpm and 18113 rpm. The first critical speed is taken as a base for design so the deformation under this critical speed is plotted in black in Figure 13. The elliptic orbit shown in Figure 13, have proved the shaft is stable at critical speed and vibrating with the amplitude 20.992mm. The finite element analysis of glass-epoxy composite driveshaft has been carried out using ANSYS Composite PrepPost (ACP). The shaft is analyzed for torsional strength, torsional buckling strength, modal, and critical speed. The ply wise failure analysis of glass- epoxy driveshaft have been performed to predict the first ply failure and safety margin. The torsional strength and torsional buckling strength of glass-epoxy composite driveshaft are 3589 Nm and 25757 Nm, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000162_bf01972478-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000162_bf01972478-Figure2-1.png", + "caption": "Fig. 2. Arthograph for knee joint", + "texts": [ + " The original data were obtained by using an apparatus in which the sinusoidal motion was derived from that of a pendulum, and later by a crank driven by a variable speed motor which permitted investigation of a wider *) Paper presented to the British Society of Rheology Conference on Rheology in Medicine and Pharmacy, London, April 14 15, 1970. range of frequencies (from 0.03 3 cycles/second), velocities (15 radians/see minimum) and accelerations (280 radians/sec maximum). A portable version of this apparatus is shown in fig. 1. To contrast these results with those of a weight bearing joint, an arthrograph operating on similar principles was devised for the knee (fig. 2). The lower leg was attached to an adjustable holder, which could aecommodate any size of leg. It was strapped on so that no movement was permitted at the ankle joint during oscillation. Oscillations could be started from any position of flexion or extension. No tests were done in the final 20 30 ~ of full extension, due to the rotation of the lower leg which occurred at that amplitude. The knee was aligned carefully with the axis of rotation of the drive shaft, so that no gross lateral and vertical movement occurred" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002969_j.matpr.2021.05.300-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002969_j.matpr.2021.05.300-Figure3-1.png", + "caption": "Fig. 3. Al oxide clutch plate temperature results.", + "texts": [ + " Calculation of stresses, and greatest weights utilizing EulerLagrange conditions. Developing 3D model of MPCs utilizing Solid edge programming. Exporting the model of MPCs from strong edge to ANSYS 19.2 work bench for FEA. Static and dynamic examination of MPCs utilizing ANSYS work seat at various working condition. Determining the wear, most extreme disfigurement and equal worries for every material utilizing FEA Showing the connection of disfigurement and obstruction property of every material utilizing figures and tables (Fig. 1. Fig. 2. Fig. 3. Fig. 4. Fig. 5. Fig. 6. Fig. 7. Fig. 8. Fig. 9. Fig. 10. Fig. 11. Fig. 12. Table 1). Comparing and conversation of results for choice of better covering materials. Fig. 5. Cast iron clutch plate heat flux results. Fig. 6. Cast iron clutch plate heat flux results. Grip disappointment and harm because of extreme frictional warmth and warmth changes to the grasp counter mate circle frequently happens to a car grips. This circumstance adds to warm weakness to the segment which causes the grasp counter mate plate to break and twist" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001015_aas-007-0628-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001015_aas-007-0628-Figure3-1.png", + "caption": "Fig. 3 Graphical depiction of formation shape regulation", + "texts": [ + " The geometry-based formation control approach is comprehensible, but the main difficulty is to deal with the speed uncertainty of the leader. The offset of the follower is one of the key factors to show the formation structure in formation coordinates. In general, the initial poses of individual robots are random, and regulating the initial headings firstly is necessary to simplify the offset control rules. In other words, all the initial headings should be turned to the desired formation orientations. Note that the error of offset should be eliminated in a progressive process. Fig. 3 interprets the quantities of formation shape regulation graphically. O\u0303p denotes the origin of the formation coordinate system, and Of is the origin of the follower polar coordinates. P denotes the position of the leader, D is the desired following position, and F is that of the follower at the present time. d is the original offset, and dq is the desired offset. G is the real regulated target of the follower. Offset d can be deduced by d = (\u03c1f \u2212 \u03c1p)sign(\u03c9p) (14) To adjust the offset, the formation control system should satisfy the differential equations \u00bb \u03c9\u0307 h\u0307 \u2013 = \u02c6 k1 k2 \u02dc \u00bb 1 \u2212sign(\u03c9p) \u2013 (d\u2212 dq) (15) where k1, k2 are the user-selected controller gains and 0 \u2264 k1, k2 \u2264 1", + " Thus, the following results can be derived \u03c9f = k1(d\u2212 dq) + \u03c9p vf = (\u2212k2sign(\u03c9p)(d\u2212 dq) + Rn)(k1(d\u2212 dq) + \u03c9p) Note that the followers should not go ahead of their leader in our scheme, so the following inequality constraint must be satisfied: (\u03b8p \u2212 \u03b8\u0302f )sign(\u03c9p) \u2265 0 (17) where \u03b8\u0302f is the polar angle of the real target G in coordi- nates O\u0303p . Finally, the control inputs of the followers can be obtained from the following offset regulation OTR: \u00bb \u03c9f vf \u2013 = 2 4 k1(d\u2212 dq) + \u03c9p \u2212k1k2sign(\u03c9p)(d\u2212 dq) 2 + k1(d\u2212 dq)Rn+ \u03c9pRn \u2212 k2sign(\u03c9p)(d\u2212 dq)\u03c9p 3 5 (18) where k1, k2 \u2208 [0, 1] are control gains, which indicate the rate of convergence in offset regulation. As can be seen in Fig. 3, spacing distance l between two robot locations P and F is regulated to its desired distance lq. However, the adjustable process of spacing distance should be progressive in our approach, which can be realized by the following recursion. ld = (r1ld + r2lq)/ 2X i=1 ri (19) where r1 and r2 are the adjustable spacing distance coefficients. Thus, desired following angle \u03b8d can be updated by \u03b8d = \u03b8p + \u00b5 arccos((\u03c12 p + \u03c12 d \u2212 l2d)/2\u03c1p\u03c1d) (20) Under the dynamic formation framework, the controlled quantities of the leader are piecewise constants" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000171_detc2004-57029-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000171_detc2004-57029-Figure8-1.png", + "caption": "Fig. 8 Workspace in three dimensions( ) o0=\u03c9", + "texts": [ + " So is the rcond2. When both rcond1 and rcond2 are 1, the dexterity of the PMT is best. 3. Examples The main structural parameters of the prototype of PMT developed in our laboratory are as follows rA=720mm \uff0c W \uff1d 780 mm \uff0c 2/\u03c0\u03b8 = \uff0c mm\uff0c200=Br 5/2\u03c0\u03c6 = \uff0ca4=180mm 3.1 The workspace of the prototype Based on the given structural parameters, through numerical calculation, some graphical representations of the workspace are obtained in three dimensions. The workspace graphs in line with and are shown in Fig.8 and Fig.9. It is obvious that the bigger the agile cutter sloping angle is the smaller the workspace. The workspace corresponding to the bigger cutter sloping angle is contained by the workspace corresponding to the smaller one. o0=\u03c9 o25=\u03c9 3.2 Dexterity analysis of the prototype The reciprocals of condition numbers on the section and pose 1000=x 0== \u03b2\u03b1 are shown in Fig.10. The reciprocals of condition numbers on nloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/09/2016 Te the section 1000=x and pose are shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002838_012021-Figure19-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002838_012021-Figure19-1.png", + "caption": "Figure 19. Stress Distribution in E-Glass", + "texts": [ + " Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.5. Analysing Testing Result of E-Glass 3.5.1. Total Deformation The Max. and Min. Total Deformation in E-Glass is 0.23872 mm and 0 mm respectively shown in Figure 18. The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. 3.5.2. Stress Distribution The Max. and Min. Stress Distribution in E-Glass is 62.053 MPa and 0.37676 MPa respectively shown in Figure 19. 3.5.3. Strain Distribution The Max. And Min. Strain Distribution in E-Glass is 0.0008937 and 0.0000069443 respectively shown in Figure 20. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.6. Analysing Testing Result of Basalt Fiber 3.6.1. Total Deformation The Max. And Min. Total Deformation in Basalt Fiber is 0.19542 mm and 0 mm respectively shown in Figure 21. The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000624_bfb0110380-Figure29-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000624_bfb0110380-Figure29-1.png", + "caption": "Fig. 29. Nyquist Plots", + "texts": [ + " T h e y all sa t i s f ied the specif ications. Note t h a t the p l an t is uns t ab l e over p a r t of i ts p a r a m e t e r range. This has no effect on pe r fo rmance . Fig. 28-a shows uni t s t ep responses for 81 uns t ab le cases, Fig. 28-b for 81 s t ab le cases, Fig. 28-c for 27 cases of a pa i r of p l an t poles on i m a g i n a r y axis. T h e y are ind i s t ingu i shab le . I t is i n t e re s t ing to see the mechanism whereby the system is stable for both the unstable and stable plant cases. Fig. 29 shows the Nyquist encirclements which prevail for the 3 plant cases of: (a)unstable P(s) with two right half plane poles, (b)stable P(s ) , and (c)pair of poles on imaginary axis (giving semi-infinite plant template) . Figs. 30-a,b,c shows simulations of the same 3 classes of L(j(w) on the Nichols Chart. Fig. 31-a presents the system output due to unit disturbances entering directly at the system output , so output y(0) = d(0), because no practical feedback can be instantaneous, as it would require infinite loop bandwidth" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001413_j.jsv.2008.02.056-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001413_j.jsv.2008.02.056-Figure1-1.png", + "caption": "Fig. 1. Typical ring element cross section and coordinate systems. The x coordinate is measured circumferentially. AA, BB, CC, and DD are the corners of the ring element.", + "texts": [ + " The solution for the free vibration characteristics of rotationally symmetric shells with meridional variations in the shell parameters by means of his multisegment direct numerical integration approach was also obtained ee front matter r 2008 Elsevier Ltd. All rights reserved. v.2008.02.056 ARTICLE IN PRESS Nomenclature a flexibility matrix for the base structure, Fig. 13 aij deflection of base structure in coordinate i due to a unit value of force in coordinate j, elements of flexibility matrix a A cross-sectional area of ring element c value of coordinate y0 at coordinate 1, Fig. 1 d value of coordinate y0 at coordinate 2, Fig. 1 Di0 deflection of base structure in coordinate i due to a combined external loading and temperature change, elements of column matrix of displacements with F \u00bc 0 E Young\u2019s modulus F column matrix of redundant forces F 01 unit virtual longitudinal force per unit length applied at ring coordinate 1 h height of ring element measured parallel to ring element axis of symmetry H1 radial force per unit length applied at ring coordinate 1 H2 radial force per unit length applied at ring coordinate 2 H 01 unit virtual radial force per unit length applied at ring coordinate 1 H 02 unit virtual radial force per unit length applied at ring coordinate 2 Hp radial load per unit length at R3 due to external pressure HT P total circumferential radial loading from HP Iz 0 moment of inertia of ring cross section about z0 axis \u2018 meridional length of ring element, Fig. 1 L longitudinal load per unit length applied at ring coordinate 1 Mz 0, My 0 induced internal moments about z0 and y0 axes, respectively, due to applied loadings Mj total internally redundant meridional moment in coordinate j of the base structure my0 virtual moment applied about y0 axis at points A and B, Fig. 2 p uniform pressure on ring element surface Pi total internally redundant radial force in coordinate i of the base structure Q temperature rise of ring R radius of centroid of ring element cross section R1 radius of ring element cross section at ring coordinate 1 R2 radius of ring element cross section at ring coordinate 2 R3 radius of center of pressure for uniform pressure on outer face of ring element s variable distance measured along ring element meridian from coordinate 1 S direct force due to radially applied loads t thickness of ring element of rectangular cross section t1 thickness of ring element measured nor- mal to meridional centerline at coordinate 1 t2 thickness of ring element measured normal to meridional centerline at coordinate 2 T torque per unit length about centroid of ring cross section Tp torque per unit length about centroid of ring cross section due to external pressure TL torque per unit length about centroid of ring cross section due to L T01 torque per unit length about centroid of ring cross section due to combined pressure and longitudinal loading with pressure loading resisted at R1 T02 torque per unit length about centroid of ring cross section due to combined pressure and longitudinal loading with pressure loading resisted at R2 T3 meridional moment per unit length applied at ring coordinate 3 T4 meridional moment per unit length applied at ring coordinate 4 T 03 unit virtual meridional moment per unit length applied at ring coordinate 3 T 04 unit virtual meridional moment per unit length applied at ring coordinate 4 Vp1 longitudinal load per unit length resisted at ring coordinate 1 due to pressure loading Vp2 longitudinal load per unit length resisted at ring coordinate 2 due to pressure loading Vp longitudinal load as force per unit length at R3 due to external pressure T", + " Bending and direct stresses in the plane of the cross section due to edge moments, radially applied forces, longitudinal loads, pressure loadings, and thermal effects will be assumed to vary linearly between the edges of the cross section. Deformations due to shearing stresses and direct stresses through the thickness of the ring cross section will be neglected. It will be assumed that the material is homogeneous, isotropic, and linearly elastic and that the shell structure is subjected to loadings that are statically applied. The geometry and the coordinate system for a typical ring element cross section are shown in Fig. 1. The orientation of the middle surface of the rectangular cross section of the ring element with respect to the axis of symmetry of the circular ring is defined by the angle a. The coordinate x is measured in the circumferential direction of the ring at the centroidal axis of the cross section. With respect to the cross section, the coordinate y is measured along the middle surface of the element from the centroidal axis and the coordinate z is measured therefrom in a direction normal to the middle surface. The coordinate y0 is measured from the centroidal axis in a line parallel to the axis of symmetry of the ring element and z0 is measured therefrom in a direction normal to the axis of symmetry as shown in Fig. 1. Before proceeding with the development of the influence coefficients for a typical ring element, it is necessary to determine the circumferential stresses over the ring element cross section due to both uniformly distributed applied torques T3 and T4 and radially applied forces H1 and H2. The circumferential stresses sx will be expressed in terms of the y0\u2013z0 system of coordinates. However, for purposes of integration over the ring element cross section to determine the flexibility influence coefficients, the y\u2013z system of coordinate will be used. It is thus seen that the circumferential stresses will be as determined for a coordinate system of y0\u2013z0 which has no axis of symmetry. To aid in the development of the expressions for the forces and stresses on any ring element cross section shown in Fig. 1, plans and sections showing these forces are given in Fig. 2. Shown in Fig. 2 are a plan view of the ring element, a cross section showing the y0\u2013z0 coordinate system, and a plan view of half of a ring element showing real and virtual forces applied thereto. From the half-plan view, it is seen that, for a uniform torque T around the element, equilibrium requires that Mz0 \u00bc Z p=2 0 TR cosfdf \u00bc TR: (1) To determine My0, cut the structure on the diametral line A2\u2013B2 and apply a unit virtual moment my0 \u00bc 1 as shown in Fig", + " (8) are only those stresses due to bending of the ring element due to axisymmetric torques T1, T2, H1, H2, and other applied loadings about the centroidal axis of the element. In the case where radial forces (typically H1) are also applied, there are developed in the ring element not only bending stresses but also direct stresses which must be added to the bending stresses. Typically, for a radial force H1 applied at ARTICLE IN PRESS T.A. Smith / Journal of Sound and Vibration 318 (2008) 428\u2013460436 coordinate 1 in Fig. 1, the direct circumferential force S in the ring element becomes S \u00bc Z p=2 0 H1R cosfdf \u00bc H1R1. (9) The stress sx in the ring element due to direct forces (typically for the H1 application) is therefore sx\u00f0D\u00de \u00bc H1R1 A . (10) In the development of the expressions for the ring element influence coefficients, both bending stresses and any developed direct stresses will be considered. In addition to the stresses sx discussed immediately hereinbefore, the meridional stresses sy are to be considered. These are the stresses caused by the meridional bending moments Mx about the circumferential coordinate line x described in Fig. 1, by the radial forces normal to the axis of symmetry of the ring element, and by any applied or induced forces acting on the ring element. In general, for the development of the ring element influence coefficients, the forces H1 and H2 are assumed to be uniformly distributed across the thickness t of the element at the respective points of application, and they are zero at the opposite edge of the element. The meridional direct stresses sy have been assumed to vary linearly between these two edges of the element. In a similar vein, the bending stresses sy have been assumed to vary linearly between the element edges. The stresses sy due to each of the applied forces H1, H2, T3, and T4 based upon the above assumptions may be developed directly from Fig. 1, where the coordinate s shown there is given by s \u00bc y\u00fe 0:5\u2018 (11) and where the circumference of the ring element at any value of s is given by b \u00bc 2p\u00f0R1 \u00fe s sin a\u00de. (12) The stresses sy(H1) are therefore sy\u00f0H1\u00de \u00bc 2pR1H1 sin a\u00f01 s=\u2018\u00de bt \u00bc R1H1 sin a\u00f0y=\u2018 \u00fe 0:5\u00de t\u00bdR1 \u00fe \u00f0 y\u00fe \u2018=2\u00desin a (13) The stresses due to H2 are sy\u00f0H2\u00de \u00bc \u00f0R2H2 sin a\u00des t\u00f0R1 \u00fe s sin a\u00de\u2018 \u00bc R2H2 sin a\u00f0y=\u2018 0:5\u00de t\u00bdR1 \u00fe \u00f0 y\u00fe 0:5\u2018\u00desin a . (14) The total applied circumferential bending moment due to T3 is MT x \u00bc 2pR1T3\u00f01 s=\u2018\u00de, while the circumferential moment of inertia is IT x \u00bc 2p\u00f0R1 \u00fe s sin a\u00det3=12, where MT x and IT x are the values of Mx and Ix around the total circumference", + " Smith / Journal of Sound and Vibration 318 (2008) 428\u2013460 449 d40 \u00bc T02 R2 EIz0 nR\u2018 sin a 2EIz0 nV p2R2\u2018 2 cos2 a 12EIz0 , (72) dV0 \u00bc T02R2\u00f0R2 R1\u00de EIz0 nVp2\u00f0R2 R1\u00deR2\u2018 2 cos2 a 12EIz0 \u2018 cos2 a 2ERt \u00f02LR1 \u00fe Vp2R2\u00de \u00fe nHpR3 cos a Et . (73) Our complete shell structure will be composed of a number of ring elements extending in the meridional direction of the shell from one boundary to the opposite boundary. We assume that one boundary edge of the shell is that edge located at coordinates 1 and 3 as shown in Fig. 1 and that a radial force HB1 at coordinate 1 and a shell bending moment TB3 at coordinate 3 may be applied as force boundary conditions at that boundary. For the opposite boundary edge of the shell, we assume the boundary edge located at coordinates 2 and 4 as shown in Fig. 1 to constitute that boundary and that a radial force HB2 at coordinate 2 and a shell bending moment TB4 at coordinate 4 may be applied as force boundary conditions at this opposite boundary. We designate the first above-described boundary element as element 1. The second described boundary element will be designated as element N, where N is the total number of ring elements in the complete shell structure. For the boundary element 1, our free ring displacements will be given by D10 \u00bc d10 \u00fe d1T \u00feHB1d1H 001 \u00fe TB3d1T 003 , (74a) D20 \u00bc d20 \u00fe d2T \u00feHB1d2H 001 \u00fe TB3d2T 003 , (74b) D30 \u00bc d30 \u00fe d3T \u00feHB1d3H 001 \u00fe TB3d3T 003 , (74c) D40 \u00bc d40 \u00fe d4T \u00feHB1d4H 001 \u00fe TB3d4T 003 , (74d) DV0 \u00bc dV0 \u00fe dVT \u00feHB1dV H 001 \u00fe TB3dV T 003 , (74e) where the double primes on H1 and T3 denote unit values per unit of circumferential length", + " The circumferential strain, Ax, may be determined for the points of interest on any ring from 2 \u00f0n\u00de x\u00f0AA\u00de \u00bc D\u00f0n\u00de1 D\u00f0n\u00de3 \u00bdt \u00f0n\u00de 1 =2 sin a\u00f0n\u00de R \u00f0n\u00de 1 \u00fe \u00f0t \u00f0n\u00de 1 =2\u00de cos a \u00f0n\u00de , (98a) ARTICLE IN PRESS T.A. Smith / Journal of Sound and Vibration 318 (2008) 428\u2013460456 2 \u00f0n\u00de x\u00f0BB\u00de \u00bc D\u00f0n\u00de1 \u00fe D\u00f0n\u00de3 \u00bdt \u00f0n\u00de 1 =2 sin a\u00f0n\u00de R \u00f0n\u00de 1 \u00f0t \u00f0n\u00de 1 =2\u00de cos a \u00f0n\u00de , (98b) 2 \u00f0n\u00de x\u00f0CC\u00de \u00bc D\u00f0n\u00de2 \u00fe D\u00f0n\u00de4 \u00bdt \u00f0n\u00de 2 =2 sin a\u00f0n\u00de R \u00f0n\u00de 2 \u00fe \u00f0t \u00f0n\u00de 2 =2\u00de cos a \u00f0n\u00de , (98c) 2 \u00f0n\u00de x\u00f0DD\u00de \u00bc D\u00f0n\u00de2 D\u00f0n\u00de4 \u00bdt \u00f0n\u00de 2 =2 sin a\u00f0n\u00de R \u00f0n\u00de 2 \u00f0t \u00f0n\u00de 2 =2\u00de cos a \u00f0n\u00de ; n \u00bc 1; 2; 3 . . . ;N, (98d) where the subscripts 1, 2, 3, and 4 for D signify displacements in the directions 1, 2, 3, and 4 shown in Fig. 1. The circumferential stresses, sx, may be determined for the points of interest on any ring element from s\u00f0n\u00dex\u00f0j\u00de \u00bc E\u00bd2 \u00f0n\u00de x\u00f0j\u00de bQ\u00f0n\u00de \u00fe ns\u00f0n\u00dey\u00f0j\u00de, n \u00bc 1; 2; 3 . . . ;N, j \u00bc AA;BB;CC;DD. (99) The procedures that have been developed herein were programmed for the solution of actual cases by use of an electronic computer. In order to determine the accuracy of typical solutions by the flexibility method, the solution for a flat circular plate by the flexibility method will be compared with the solution obtained by use of the governing differential equations as a first case" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002864_0954406221995852-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002864_0954406221995852-Figure1-1.png", + "caption": "Figure 1. Snake-like robot with no slipping wheel.", + "texts": [ + " ~M 1=2 5VD 1=2V 1 (41) While ~M is Hermitian symmetric, the diagonalizing matrix V can be made an orthogonal matrix by appropriately choosing the eigenvectors. Then, the inverse of V is simply the transpose, so that b5VD 1=2VT (42) In order to solve dynamic equations (36), it is also necessary to find _b. Differentiating Eq. (40) with respect to time yields _bb1b _b5 d dt M (43) Algebraic equations (43) give _b. Eventually, knowing distinct b and _b, which satisfy the conditions of equation (39), equations of motion (36) can be written as _u25bT F \u00f0bTM\u0302b1bT M _b\u00deu2 (44) The first example is a planar 3-link snake-like robot (3LSR) illustrated in Figure 1. The robot consists of three links connected through two revolute joints. Each link is rigid, with a uniformly distributed mass. The 3LSR is placed on a horizontal surface. Each link is connected to a wheel with no slipping constraint, so this robot has three nonholonomic constraints. In this case, the system has five generalized coordinates and two degrees of freedom. Each joint is equipped with torque actuators s1 and s2. The GCs selected for this problem are as follows q \u00bc fx y h1 h2 h3g (45) To satisfy no slipping condition, the velocity of links at wheel contact points to the ground,~Vwi, in direction perpendicular to link axis should be zero" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001162_fedsm2007-37066-Figure23-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001162_fedsm2007-37066-Figure23-1.png", + "caption": "Figure 23 Instantaneous relative flow patterns and pressure distributions around a ski jumper for various styles.", + "texts": [ + "25 s that considerably higher pressure caused by frontal stagnation of the relative flow widely distributes not only along the front surface of the ski jumper\u2019s body but also on the back surfaces of both skis, whereas there are not so remarkable differences between top views of pressure distributions at t=0.04 s and 0.25 s. Therefore, it is understood that the flight style of V-shape is certainly effective to obtain considerable increase of lift (vertical) force to the ski jumper during the flight. Figure 23 shows instantaneous relative flow patterns around the ski jumper for the parallel style and the delta style. It is confirmed that the pressure near the back waist surface are high at t =0.25 s for both case. It is thought that the high pressure region formed near the waist of the jumper is caused by the wake structure formed to the back of the jumper immediately after shooting out. After the time, for delta style, the vertical wake is wildly formed behind the back of the jumper and the pressure of the back at t =0", + "04 s when movement of skis stars from the initial parallel-phase to the each style phase (V-shape, parallel, delta) and Tfinish is the time 0.25 s when the ski jumper finishes the ski-movement. It is very interesting that both values of lift and drag force areas remarkably increase during the transition period of ski-movement. It is thought that drag and lift forces acting of the jumper with the parallel and the delta style during the period of 0.2s < t <0.4s become lower then the case of the V-shape style because the pressure near the waist of the jumper becomes large as shown in fig.23. It is interesting that lift force acting on the jumper with the delta style during 0.8s < t <1.0s becomes almost the same value of the lift force for the case of the V-shape style, but the drag force acting on the jumper with the delta style is smaller than that of the case of the V-shape style. In this study, complex and unsteady flows around an isolated 100 m runner and a ski-jumper were calculated by applying an advanced vortex method based on the Biot-Savart law. Since the method provides completely grid-free Lagrangian calculation of unsteady and vortical flows, dynamic ownloaded From: http://proceedings" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002945_978-3-030-69178-3_15-Figure4.6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002945_978-3-030-69178-3_15-Figure4.6-1.png", + "caption": "Fig. 4.6 Collision avoidance vectors (Adapted from [25])", + "texts": [ + " The first is to stop the robot movement when the close proximity of an operator has been detected, and continue the robotmovement after the operator walks away. This scenario is suitable for operations with a limited degree of freedom, for example, inserting a pin or when moving in close proximity to the workpiece. The second scenario is applicable to such operations as transportation where the path can be dynamicallymodified to avoid collisionwith human operators and other obstacles. As shown in Fig. 4.6, when a short distance between one of the points of interests (POI) on the robot end-effector and operator (a) is detected, the direction of movements of the end-effector is modified as p1 to maintain a defined minimal distance. However, if operator (b) appears in proximity to the robot arm and the distance di between the operator and the robot becomes less than a safety threshold, the robot will stop moving. Once the situation is relaxed and dPOI is greater than the threshold, the end-effector will move back in direction p2 to its destination position. The representation of vectors taken into consideration for pathmodification is also shown in Fig. 4.6. The collision vector c is calculated as a vector between the POI on the end-effector and the corresponding closest point of detected obstacles. The vector representing the direction of the robot movement vc is decomposed into two orthogonal components vc\u2016c and vc\u22a5c, respectively, parallel and perpendicular to the collision vector c. The parallel component is calculated in Eq. (4.6) as a dot product of the movement vector and collision vector and has the direction of the collision vector. The collision vector is a unit vector with the direction of the collision vector" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002274_j.precisioneng.2021.01.004-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002274_j.precisioneng.2021.01.004-Figure2-1.png", + "caption": "Fig. 2. Geometric relation: (a) Leg length and leg angle, (b) Parallel spring (c) Approximation to a spring-mass model.", + "texts": [ + " The leg length is the line length between the body and the tiptoe. The leg angle is the angle between the vertical line which crosses the center of the body and the line of the leg length. The leg length and the leg angle are the variables of the polar base Y. Abe and S. Katsura Precision Engineering 69 (2021) 36\u201347 coordinate. The ideal characteristics of the leg spring are that the elastic force is proportional to the leg length, and that the elastic force is not determined by the leg angle. Fig. 2 shows the geometric relation of the robot body, the spring, and the leg angle. Assuming that the common mode and the differential mode of [q1, q2] are qc and qd, the variables [qc, qd] are defined as [ qc qd ] = 1 2 [ 1 1 1 \u2212 1 ][ q1 q2 ] . (5) The leg length L and the leg angle \u0398 are expressed using qc and qd as L= 2lcosqc (6) \u0398= qd. (7) From (6) and (7), the leg length and the leg angle are independent variables in terms of qc and qd. The leg spring at the knee joint is set parallel to the line of the body and the tiptoe as Fig. 2 (b). The elastic force of the leg spring can be induced from the partial derivative of the potential energy. The potential energy of the leg spring is calculated as U = 1 2 Ksp ( L 3 \u2212 l0 )2 . (8) Here, l0 is the natural length of the spring. The elastic force Fk of the leg spring in terms of [L,\u0398] is induced from the partial derivative of the potential energy by [L,\u0398] as Fk = [ \u2202U \u2202L , \u2202U \u2202\u0398 ] = [ Ksp 3 (L \u2212 l0), 0 ] . (9) As (9), the elastic force affect the leg length and does not affect the leg angle. This is same as the ideal characteristics of the spring in a spring-mass hopping. Because the center of gravity of the legged robot is set almost in the body by the parallel-link mechanism, the total dynamics become near a spring-mass model as Fig. 2 (c). The motion equation of the robot in the world coordinate is described. The motion equation in the world coordinate is J(qw)q\u0308w + g ( qw, q\u0307w ) + \u03c4k(qw)= \u03c4ref \u2212 \u03c4fric + JT acoW Fe. (10) Here, J(qw) is the inertia matrix, g(qw, q\u0307w) is the nonlinear terms of the rigid-body dynamics, \u03c4k(qw) is the elastic torque of the leg spring, \u03c4ref is the torque reference of the actuators, \u03c4fric is the friction torque, JacoW is Jacobian matrix of the tiptoe in the world coordinate, and Fe is the external force from the environment to the tiptoe" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001514_vss.2008.4570735-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001514_vss.2008.4570735-Figure3-1.png", + "caption": "Fig. 3. Limit cycle arising when w = 1.2", + "texts": [ + " The eigenvalues of the linearized system of equations at this point stipulate that in the increasing direction of c2, the system undergoes Hopf bifurcation at this operating point and turns into an unstable one displaying spontaneous oscillations due to the limit cycle. In this regime, cell mass varies in between 0.1219 and 0.1466 while the nutrient amount fluctuates in between 0.8243 and 0.8996. At the points of crossing the imaginary axis, the eigenvalues of the linearized model are approximately equal to 0 \u00b1 j1.7543, from which we infer that the self sustained oscillations are quite fast. In Fig. 3, the limit cycle and the convergence of the neighboring trajectories are illustrated for w = 1.2. In fact, limit cycles can occur for all values of admissible inflow rates, i.e. 0 \u2264 w \u2264 2. According to Bendixson theorem (See [14], [15]), since the quantity H := \u2202 \u2202c1 ( \u2212c1w + c1(1 \u2212 c2)e c2 \u03b3 ) + \u2202 \u2202c2 ( \u2212c2w + c1(1 \u2212 c2)e c2 \u03b3 1 + \u03b2 1 + \u03b2 \u2212 c2 ) = \u22122w + h(c1, c2) (3) does not vanish and does not change sign in \u039b \u2286 \u2126, no limit cycles can exist entirely in \u039b. For a given constant w, the curve of sign change for H is moved to the curve described by h(c1, c2) = 2w" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002838_012021-Figure12-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002838_012021-Figure12-1.png", + "caption": "Figure 12. Total Deformation in Kevlar 29", + "texts": [ + " Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.2.2. Stress Distribution The Max. and Min. Stress Distribution in Structural Steel is 183.73 MPa and 1.8481 MPa respectively shown in Figure 10. 3.2.3. Strain Distribution The Max. and Min. Strain Distribution in Structural Steel is 0.0010678 and 0.000012423 respectively shown in Figure 11. 3.3. Analysing Testing Result of Kevlar 29 3.3.1. Total Deformation The Max. and Min. Total Deformation in Kevlar 29 is 0.19844 mm and 0 mm respectively shown in Figure 12. The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.3.2. Stress Distribution The Max. and Min. Stress Distribution in Kevlar 29 is 36.704 MPa and 0.26231 MPa respectively shown in Figure 13. 3.3.3. Strain Distribution The Max. and Min. Strain distribution in Kevlar 29 is 0.00054321 and 0.000005542 respectively shown in Figure 14" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000335_3.6255-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000335_3.6255-Figure2-1.png", + "caption": "Fig. 2 Compound asymmetry models comprising radial e.g. offset and A) opposing coplanar trim, B) orthogonal", + "texts": [ + " When this occurs, the spin rate builds up in unison with the pitch frequency and a condition of persistent resonance then exists accompanied by large and possibly catastrophic excursions in angle of attack. It is promoted by one or more forms of inherent or acquired compound asymmetry in excess of some critical threshold level. Various criteria for its occurrence and avoidance have been developed in the literature.1\"6 This Note deals with the transient response of a vehicle in rolling trim at first intersection of the spin and pitch frequencies, and the susceptibility to steady resonance immediately thereafter for the two types of compound asymmetry shown in Fig. 2. Transient Resonance Response The characteristic excursion in angle of attack which occurs at the altitude of transient resonance is usually not of great concern. However, it does become important when considering resonance susceptibility, since the roll torques associated with this transient response will determine whether p exceeds coc, which is a necessary condition for lock-in. It has been demonstrated analytically by Kanno7 that the angle-of-attack response of a slender entry vehicle to transient resonant excitation can be conveniently expressed as a function of the nondimensional parameter X, where X = p<>(l - Ix/2Iy)/pVE smyE (1) The analysis presupposes a linear flight path, constant velocity, and constant roll rate p0", + "/p)(W/*y) (9) The invariance of the stream function requires that *(y = A) = t(y = 5) (10) From Eqs. (9) and (10), f**UdY = fipudy (11) Evaluation of the integral on the left-hand side of Eq. (11) and further simplification yields [1 - (y/S)1'\" d(|) + AT)] = Jo [1-, ' N) (12) Evaluation of the integral in Eq. (12) for various Mach numbers and compressible power-law exponents, l/n, yields the results shown in Fig. 1. The control volume selected along with other necessary geometric quantities for the two-dimensional wedge-shock/ boundary-layer interaction model is shown in Figure 2. The stream wise boundaries of the control volume are the compression surface downstream of the corner and a streamline passing through the corner shock in the inviscid region of the flow. The flow in the boundary layer passing through the control surfaces is assumed to be parallel to the compression surfaces. This precludes the existence of large flow disturbances in the boundary layer upstream of the compression corner; this is substantiated by experimental data in Ref. 2 where it was found that, in the absence of separation, the upstream influence was considerably less in extent than a length equivalent to one boundary-layer thickness" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002004_50015-9-Figure12.6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002004_50015-9-Figure12.6-1.png", + "caption": "Figure 12.6 A total knee replacement joint. (a) Schematic diagram (after Walker and Sathasiwan, 1999). (b) Photograph of a stainless steel prosthesis (courtesy R. Grimer, Royal Orthopaedic Hospital, Birmingham).", + "texts": [ + "3 Knee joints Total knee replacements (TKR) are increasingly being made to give pain-free improved leg function to many thousands of patients suffering from osteoarthritis or rheumatoid arthritis. As with hip replacements, TKR involves removing the articulating surfaces of the affected knee joint and replacing them with artificial components made from biomaterials. Such operations have a successful history and a failure rate of less than 2% per year (Figure 12.4). A typical replacement joint consists of a tibial base plate or tray, usually made of stainless steel, Co\u2013Cr or titanium alloy, with a tibial insert (UHMWPE) that acts as the bearing surface (see Figure 12.6). The femoral component largely takes the shape of a natural femoral condyle made from the above alloys, and articulates with the bearing surface, together with the kneecap or patella. The patella may be all polyethylene or metal backed. The components are either fixed by cement or uncemented. Cemented components use acrylic cement (polymethyl methacrylate). Uncemented components rely on bone ingrowth into the implant. Titanium used on its own shows evidence of bone ingrowth, but more recently this has been improved by use of hydroxyapatite (HA) coating" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002732_j.engfailanal.2021.105451-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002732_j.engfailanal.2021.105451-Figure1-1.png", + "caption": "Fig. 1. Bucket wheel excavator SchRs 630.", + "texts": [ + " For modelling of the excavator SchRs630 pylons, slewing platform and undercarriage, taking into account the appearance of the excavator SchRs630 structure, the structure was modelled by plate elements. Classical plate theory is applied. Axial bearing that connects the slewing platform and undercarriage is modelled by beam elements. This steel structure is made of steel S355J2 + N. All load case calculations were performed in accordance with relevant DIN standard, [33]. It should be noted that the units stress units in this figure, and all similar figures hereinafter, are expressed in kN/cm2, which corresponds to 10 MPa, e.g. 1.52e01 kN/cm2 would be 152 MPa (see Fig. 1). Loading of the structure includes all the weights (dead load) and estimated workload in the form of vertical, lateral and frontal force (overall digging force), according to valid standards. Reactions in bearings on pylons are calculated based on kinematic models of bucket wheel excavator. Loading of pylons is shown in Fig. 2. The results of the numerical analysis, in the form of the deformation and stress field, are shown in Figs. 3 and 4. Displacement of bearing balls is shown in Fig. 5. It can be seen that load distribution by bearing circumference is uneven" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001522_cce.2008.4578933-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001522_cce.2008.4578933-Figure5-1.png", + "caption": "Figure 5. Exponential decay technique", + "texts": [ + " In this paper, two AMC Research and Higher Degree by Research Committee 1-4244-2426-9/08/$20.00 \u00a92008 IEEE 54 typical IMO search patterns, expanding square pattern and sector search pattern, are considered as shown in Figures 2-3. The reference trajectory generator in the guidance system is a vessel simulator using the Nomoto\u2019s first-order manoeuvring model. Details can be found in [2][9]. When the ship goes along the desired trajectory, the reference heading angle can be adjusted by the exponential decay technique in [15][8][9]. Figure 5 illustrates the exponential decay technique. Coordinates of way-points (xk,yk) for a desired trajectory in the earth-fixed reference frame can be selected from the system database and are based on the manoeuvring patterns, such as IMO search patterns for maritime search and rescue mission at sea. The database consists of \u2022 Desired way-point positions: wpt.pos: {(x0,y0), (x1,y1), ..., (xk,yk)} (1) \u2022 Desired speeds between way-points: wpt.speed: Ud = {u0, u1, u2, ..., uk} (2) \u2022 Desired heading angles: wpt", + "5 C AU\u2212 + \u03c1 = \u03c4 (8) where \u03c1w is the density of sea water, Cd is the drag coefficient, A is the projected cross-sectional area of the submerged hull of ship in the x-direction, and (m \u2013 mx) is the mass included hydrodynamic added mass. The course dynamics is chosen as d dr\u03c8 = (9) d d rTr r K+ = \u03b4 (10) where T and K are ship manoeuvrability indices, rd is the desired yaw rate and \u03b4r is rudder angle. The guidance system has two inputs, thrust \u03c4x and rudder angle \u03b4r. The guidance controllers can be chosen as PI and/or PID types, see [2][9]. When the ship goes along the desired trajectory, the reference heading angle can be adjusted by the exponential decay technique as shown in Figure 5. See detailed information in [9][15][9]. Heading and position errors when the ship is moving along the desired trajectory are calculated as follows 1 d d 2 d d 3 d e (x x)cos (y y) sin e (x x) sin (y y)cos e \u2212 \u03c8 + \u2212 \u03c8 = = \u2212 \u2212 \u03c8 + \u2212 \u03c8 \u03c8 \u2212 \u03c8 e (11) where e1 = path tangential tracking error e2 = cross-track error (normal to path) e3 = heading error If the rudder-roll damping controller is switched on the vector of errors including roll error (e4) becomes 1 d d 2 d d 3 d 4 e (x x)cos (y y) sin e (x x) sin (y y) cos e e 0 \u2212 \u03c8 + \u2212 \u03c8 \u2212 \u2212 \u03c8 + \u2212 \u03c8 = = \u03c8 \u2212 \u03c8 \u2212 \u03c6 e (12) If the speed controller is on the speed error will be calculated 5 de U U= \u2212 (13) In order to get data (velocity, position) the navigation system consists of GNSS receivers with D-GNSS receivers when the vessel is running along a coast where D-GNSS and correction signals are available" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002886_iemdc47953.2021.9449529-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002886_iemdc47953.2021.9449529-Figure5-1.png", + "caption": "Fig. 5. Existing topology A from [4] with winding factors kmws = 0.97, khws = 0.7, khwf = 0.7: a) Geometry b) circuit diagram with diode rectifier (dashed) and differential step down converter (dotted); c) main machine and harmonic machine stator MMF at same current density over half a mechanical rotation", + "texts": [ + " For each topology, a HESM is designed for a rated torque of 30 Nm and a speed of fmech = 50 Hz. The machine specifications are obtained using analytical scaling laws such that similar torque, voltage rating, magnetic loading, electric loading and excitation power are achieved. The specifications are summarized in Table I. For each of the four HESMs denoted A-D the analytical model (1) is parametrized by FEA. The comparison between the topologies is performed by calculating the rated torque of each machine when excited by the excitation system. In Fig. 5a and b the existing topology is shown which requires no geometry or winding modifications to the pm machine and has no auxiliary windings on either stator or rotor. Instead, the two parallel paths of the stator and rotor winding are fed by power electronics. For the stator a B12 inverter enables feeding each phase (=parallel path) U1, U2, V1, V2, W1 and W2 with separate currents Ikc(t) = I\u0302h cos(\u03c9ht+ (k\u2212 1) 2\u03c0 3 \u2212 (c\u2212 1)\u03c0)) (6) + I\u0302m cos(\u03c9mt+ (k\u2212 1) 2\u03c0 3 ) Authorized licensed use limited to: California State University Fresno", + " The harmonic currents (index h) of two parallel paths k1 and k2 flows in opposing directions as if they were seriesconnected. Thus, for ph pole-pair operation, each inverter leg only provides half of the total phase voltage of phase k. The maximum modulation index yields kh mod = 2\u221a 3 . The rotor winding consists of four coils wound around every second salient-pole and is a single-phase variant of the stator winding. The winding layout and the corresponding MMFs of pm operation and ph operation are shown in Fig. 4. The rotor winding is fed by the currents IFr = Ifm + Ifh cos r\u03c0 with r = 1, 2, 3, 4. (7) In Fig. 5b, it can be seen that a rectifier and a buck converter couple the rotor winding paths of pole pair ph and pm. The converter sets a constant duty cycle of Ta = 8 % which enables very low harmonic copper losses. For a more detailed explanation of the function of the rotor circuitry refer to [4]. The results for the rated operation of topology A are shown in Fig. 5 In two-speed grid-connected induction motors, poleamplitude-modulation (PAM) windings are used to allow switching between different synchronous speeds which at fixed grid frequency fel corresponds to two pole-pair numbers pm and ph. The change between synchronous speeds is achieved by changing the winding configuration with the help of mechanical switches. While this switch configuration does not enable the simultaneous creation of both pm and ph fields, the concept of topology A can be transferred to PAM windings by feeding the coil groups which are normally switched by mechanical switches from different inverter legs (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003296_j.mechmachtheory.2021.104440-Figure10-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003296_j.mechmachtheory.2021.104440-Figure10-1.png", + "caption": "Fig. 10. Mechanical model of the system with rigid shafts", + "texts": [ + " A new vector {r} is introduced, and the 1st-order ODEs are then written as: {r} = {q\u0307} (12) { {q\u0307} {r\u0307} } = ( [0] [I] \u2212 [M] \u2212 1 [K] \u2212 [M] \u2212 1( \u03a91[G] + [C]) ){ {q} {r} } + ( {0} [M] \u2212 1 {Q} ) (13) The equations of motion of Eq. (13) are solved using a stiff numerical backward differencing scheme with quasi-constant step size. The maximum step size is set to 5 \u03bcs to ensure accurate capturing of the transient effects. Information on the system parameters is supplied in Table 1. In order to investigate the effects of shaft flexibility on the response of the system, a baseline case is constructed with rigid shafts in place of flexible elements (Fig. 10). The engaged gear pair and rolling element bearings are the same as those used in the flexible problem, and the system is driven by a kinematic input to the pinion. Each shaft has only 1 lateral degree of freedom in each of the xand y-directions, therefore the bearings supporting the shaft act as springs in parallel and take equal deflection. The reaction forces B. Friskney et al. Mechanism and Machine Theory 165 (2021) 104440 developed in each rolling element bearing in the same axis will depend upon their individual stiffness values" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002838_012021-Figure16-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002838_012021-Figure16-1.png", + "caption": "Figure 16. Stress distribution in S-Glass", + "texts": [ + " Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.4. Analysing Testing Result of S-Glass 3.4.1. Total Deformation The Max. and Min. Total Deformation in S-Glass is 0.18974 mm and 0 mm respectively shown in Figure 15. The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. 3.4.2. Stress distribution The Max. And Min. Stress distribution in S-Glass is 59.887 MPa and 0.3646 MPa respectively shown in Figure 16. 3.4.3. Strain Distribution The Max. and Min. Strain distribution in S-Glass is 0.00069941 and 0.000005478 respectively shown in Figure 17. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.5. Analysing Testing Result of E-Glass 3.5.1. Total Deformation The Max. and Min. Total Deformation in E-Glass is 0.23872 mm and 0 mm respectively shown in Figure 18. The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002993_s00170-021-07405-8-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002993_s00170-021-07405-8-Figure8-1.png", + "caption": "Fig. 8 Finite element model of precision sizing", + "texts": [ + " The splined punch was first passed through the internal tooth by a press, and then the external gear was reshaped during the cold sizing process. The finite element simulation was conducted using the commercial software DEFORM. The error billet was established by measured tooth thickness andM value. In order to focus on the deformation of a complete tooth and save computer processing time, the FE model of precision sizing operation for cold extruded sun gear was performed using a 1/26 section of billet and tools (punch and die) due to the symmetrical structure, as shown in Fig. 8. The billet was considered an elastoplastic body, while the tools were defined as rigid bodies because less deformation was observed in the precision sizing process. The billet was divided as tetrahedron mesh and materials around the internal and external teeth were locally refined with the ratio of 0.1. The friction factor was assumed to be constant and the friction type was shear. The conditions of finite element simulation are summarized in Table 3. The billet material is AISI 4120 alloy steel, which is widely used in manufacturing gears because of its good performance in the cold forming process [20, 22]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002838_012021-Figure13-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002838_012021-Figure13-1.png", + "caption": "Figure 13. Stress Distribution in Kevlar 29", + "texts": [ + " Analysing Testing Result of Kevlar 29 3.3.1. Total Deformation The Max. and Min. Total Deformation in Kevlar 29 is 0.19844 mm and 0 mm respectively shown in Figure 12. The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.3.2. Stress Distribution The Max. and Min. Stress Distribution in Kevlar 29 is 36.704 MPa and 0.26231 MPa respectively shown in Figure 13. 3.3.3. Strain Distribution The Max. and Min. Strain distribution in Kevlar 29 is 0.00054321 and 0.000005542 respectively shown in Figure 14. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.4. Analysing Testing Result of S-Glass 3.4.1. Total Deformation The Max. and Min. Total Deformation in S-Glass is 0.18974 mm and 0 mm respectively shown in Figure 15. The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001650_pes.2007.386081-Figure9-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001650_pes.2007.386081-Figure9-1.png", + "caption": "Fig. 9. Equivalent circuit for superimposed quantities under a reverse fault", + "texts": [ + "0 << opk , the diameter of characteristic circle S2 is SYopSYSY ZZkZZZZ +>\u2212+=\u2032\u2212\u2032 2)/12/()(2 Where SY ZZ + is the radius of the impedance characteristic circle of (3), accordingly the operating area of S2 is so larger that the corresponding relay has greater capability against fault resistance. (c) Circle S3 corresponds to C2 in Fig.4 and (5). Because of 1=Ck , inR kk = and RC kk > , then YY ZZ >\u2032 . Since equation (5) is regarded as the blocking criterion of the inphase area, this characteristic is of advantage to ensure the resistive tolerance of (4). The equivalent circuit for superimposed quantities under a reverse fault in a two-source system is shown in Fig.9. LZ is the whole line impedance. For convenience, let RLL ZZZ +=\u2032 . The fault components for reverse fault can be described according to Fig.9 as follows. )( YL ZZIU \u2212\u2032\u2206=\u2032\u2206 && and )(|0| LF ZZIU \u2032\u2212\u2206= && . Then (11) can be rewritten as: oo 270 ])[()1( ])([)1( 90 < \u2212\u2032+\u2212\u2212+ \u2032\u2212\u2212\u2212+\u2212< YLRCRC LRCYRC ZZkkZkk ZkkZZkk Arg Where CRZZ gF /\u2212\u2212= . Let: \u23a9 \u23a8 \u23a7 \u2212+\u2212\u2032+\u2032=\u2032 +\u2212\u2212\u2032\u2212+=\u2032\u2032 )1/()( )1/())(( RCYLLR CRYLCRYY kkZZZZ kkZZkkZZ (15) It follows that oo 27090 < \u2032\u2032\u2212 \u2032\u2212< Y R ZZ ZZ Arg (16) According to (16) and the detailed parameters of (4)-(6), the corresponding operation characteristics of (1) and (4)-(6) for reverse fault can be drawn as shown in Fig.10. And the corresponding relations between each curve and equation are shown in Table II", + " (c) The operating characteristic lies to the first quadrant during reverse faults whereas measured impedance lies to the third one, so the criterion itself has clear directionality. According to compensated voltage for forward faults, (7) of Criteria 2 can be rewritten as follows. \u03b1\u03b1 \u2212> + \u2212>\u2212 00 180360 S Y ZZ ZZ Arg (17) Its characteristic curve is an offset circle to the first quadrant as shown in Fig.11. Considering the angle difference between the pre-fault compensated voltage and the source voltage under load condition, the characteristic circles in Fig.9 and Fig.11 will shift a certain angle. For the relay at the sending end, the characteristic circles will shift to negative real axis, and this is contrary at the receiving end. The above analyses show that the operating characteristics of each group criterion are all circular characteristics on impedance plane, so the relay is called the fault component distance relay with double circular characteristics in this paper. A. Integral algorithm To ensure operating speed for forward faults near the start of the protected zone, the first group criterion is carried out with integral algorithm based on R-L model [6][7], while the zero-sequence current compensation is canceled to avoid transient overreaching [8]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000629_9780470612286.ch1-Figure1.7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000629_9780470612286.ch1-Figure1.7-1.png", + "caption": "Figure 1.7. Description of frame R0 in frame R-1", + "texts": [ + " Relation j-1Tj becomes: j-1Tj = Rot(y,\u03b2j) Rot(x,\u03b1j) Trans(x,dj) Rot(z,\u03b8j) Trans(z,rj) [1.43] The nominal value of \u03b2j is zero. When zj-1 and zj are not parallel, \u03b2j is not identifiable. We also note that when zj-1 and zj are parallel, we identify either rj-1 or rj. The maximum number of parameters per joint frame is therefore four. 1.4.2.2. Parameters of the robot\u2019s location We associate index \u20131 to the world frame Rf. Since this reference frame can be chosen arbitrarily, six parameters (\u03b3z, bz, \u03b1z, dz, \u03b8z, rz) (see Figure 1.7) are needed, in order to define the robot base frame R0 in Rf. These parameters are defined by [KHA 91]: Z = -1T0 = Rot(z,\u03b3z) Trans(z,bz) Rot(x,\u03b1z) Trans(x,dz) Rot (z,\u03b8z) Trans(z,rz) [1.44] and since \u03b11 and d1 are zero, we can write that: -1T1 = Rot(x,\u03b10) Trans(x,d0) Rot(z,\u03b80) Trans(z,r0) Rot(x,\u03b1'1) Trans(x,d'1) Rot(z,\u03b8'1) Trans(z, r'1) [1.45] with: \u03b10 = 0, d0 = 0, \u03b80 = \u03b3z, r0 = bz, \u03b1'1 = \u03b1z, d'1 = dz, \u03b8'1 = \u03b81 + \u03b8z, r'1 = r1 + rz. Hence, equation [1.45] represents two transformations of the classic type [1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001020_icsma.2008.4505592-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001020_icsma.2008.4505592-Figure2-1.png", + "caption": "Fig. 2 Laser scans with points connected by lines", + "texts": [ + " If mobile robot turns around sharp, laser range finder makes over a range of calculation of angle. In result, angle histogram matching method calculates wrong rotated value by match between other angle values. To solve this problem, we use vector when an angle calculate. It calculates a vector element of points of neighborhood. Based on the vector element calculated, the method makes a histogram. By correlation method, it calculates a rotated value. The scan data is calculated a magnitude and direction of a point adjacent as look in Fig. 2. Fig. 3 shows an angle histogram that composed an ingredient of direction using a data of Fig. 2. Fig. 3 shows direction angle histogram of scan data. 2.2 Translation A translation is used a method of Thomas et al. [1]. The translation is calculated by generating two histograms represented the distribution of the distances of the scan points from the x-axis histogram and y-axis histogram. Firstly, the current data is compensated with a rotated value. And then two scan data are made parallel to x-axis based on the main direction. Main direction is peak value of a histogram. Two scan data make x-axis histograms" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002838_012021-Figure11-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002838_012021-Figure11-1.png", + "caption": "Figure 11. Strain Distribution in Structural Steel", + "texts": [ + " The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.2.2. Stress Distribution The Max. and Min. Stress Distribution in Structural Steel is 183.73 MPa and 1.8481 MPa respectively shown in Figure 10. 3.2.3. Strain Distribution The Max. and Min. Strain Distribution in Structural Steel is 0.0010678 and 0.000012423 respectively shown in Figure 11. 3.3. Analysing Testing Result of Kevlar 29 3.3.1. Total Deformation The Max. and Min. Total Deformation in Kevlar 29 is 0.19844 mm and 0 mm respectively shown in Figure 12. The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.3.2. Stress Distribution The Max. and Min. Stress Distribution in Kevlar 29 is 36.704 MPa and 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003436_j.apacoust.2021.108345-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003436_j.apacoust.2021.108345-Figure2-1.png", + "caption": "Fig. 2. Gear pair meshed in FEM software.", + "texts": [ + " The driven wheel has friction on its axle and it is modelled as an elastic solid. The driving wheel is a rigid solid since the model is used to study the rattle movement and not the stresses produced during the gear process. As well, these simulation conditions are suitable for the free load situation of the gear system proposed in this document. Considering one of the solids as rigid and the other as elastic makes calculation time more efficient. The profiles of the teeth of both wheels are modelled as a contact with friction. As shown in Fig. 2, meshing is more detailed on the tooth profiles than on the inside of the wheel to improve contact calculation. Fig. 3 shows an example of the output of the dynamic model, where the angular output speed of the driven wheel _u2 is seen oscillating in a rattle movement around the corresponding value given by its transmission ratio. The points marked on the curve show the instant where an impact occurs between the gear teeth. The data of these points is used as an input of the analytical sound model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000024_2005-01-3371-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000024_2005-01-3371-Figure6-1.png", + "caption": "Fig. 6 Normalized error in shear stress distribution over the gear;", + "texts": [ + " 4(a) clearly shows a discrepancy in the radial stress distribution in a given region, and this local region turns out to be correspondent to be the damaged tooth. Same observation is made on inspection of Fig. 4(b) with respect to the tangential stress distribution. Figs. 5(a-b) show the line plots of the error in radial and tangential stress distributions, respectively. The abscissa denotes the grid point index; the first 420 grid points represent the first tooth; the last 420 grid points represent the nineteenth tooth. The damage is clearly predicted around the first and the 19th teeth. Finally, Fig. 6 shows the shear stress error distribution identifying the damaged tooth. stress This study thus establishes that for structural health monitoring, physics-based modeling and simulation is required to provide nominal solutions, both normal and damage, as reference solutions, which can then be used to train machine learning algorithms to subsequently test the incoming sensor data over a distributed network of sensors to detect damage. The first-principles modeling and simulation can also be used to guide an optimal placement of sensors over a given structural system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000007_robot.2003.1241646-Figure9-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000007_robot.2003.1241646-Figure9-1.png", + "caption": "Fig. 9. Calculating the mining data of neural network", + "texts": [ + " The BPNN has three inputs and two outputs, and the activation function in the output layer is also the linear summation function as shown in Fig. I, hut here we use the hyperbolic tangent function as the activation function in the hidden layer, because this function is very close to a linear function fx(x) = x and the nonlinear part of this function is used to overcome some disturbance which is described in the next section. In order to use a bigger training rate, the weight update rule includes a momentum term. B. Training data of neural network Fig. 9 shows how to calculate the training data of the NN. At time t - 1, the local coordinates E,, COi and the angle between them, e,. me expressed by COilt-l) and ai(t-i). Similarly, at time t, we have '&), Coi(t) and qt). The position of point mi, the middle point between two passive joints of the end-effector on the arm in xi, 'xrni = rxmi iy,,,;]T, can be calculated through (5) and (7)- (ll), then at times t - 1 and t , we denote them as i(t-l)xz: and '(')\"Ai respectively. Transforming i(t)x:i into coordinate Ci(l-l) can he realized by oi " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003042_s11668-021-01191-x-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003042_s11668-021-01191-x-Figure7-1.png", + "caption": "Fig. 7 Zoomed view", + "texts": [ + " The existing housing model had tensile stresses more than the permissible value for this aluminum alloy, as shown in following Fig. 4a and b. The maximum stresses at the mounting bosses were near about 283MPa, which is above the permissible limit of the material. A 0.67 mm displacement was Fe Si Mn Zn Cu Ti Mg Cr Al 0.094 0.69 0.034 0.001 0.23 0.03 0.85 0.06 98.011 also calculated. Similarly, all the modified models are subjected to finite element analysis, and the results are shown in Figs. 5 and 6. The results of the analysis of all the modified models are listed in Table 3. Figure 7 shows the zoomed model at the mounting bosses showing the thickness and the rib height for all the four modified models. From Table 3 it is observed that the existing and the modified model 1 exceeds the permissible stress value for the aluminum housing (Table 1). Models 2 and 3 exhibit lesser deflections and stresses. The modified model 4 gives the lowest stress and displacement values among all the models. From the results it is observed that, at the nearly same frequency range (387 to 392 Hz), the stress and the displacement values at the mounting bosses of EGR housing decreased from the existing for the existing model to modified model 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001343_icept.2007.4441500-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001343_icept.2007.4441500-Figure1-1.png", + "caption": "Fig. 1O Segment extensions in the FAB surface for the different VI/Vo values", + "texts": [ + " In the wire bonder, the constant velocity Vo is a bonding parameter used by assembly engineers to assess the size of the bonded ball. The impact time t, is also an adjustable bonding parameter. Different V1 values may be acquired by adjust the impact time. So the V1/V0 is also considered as an adjustable bonding parameter in this study. According to some correlative experimental data, the range of VIV2 value may be 0.1-0.25. During the whole bonding, the substrate is usually kept at 423-493K, and the temperature is named as bonding temperature T. 2 Modeling of Thermosonic Bonding 2.1 physical process Figure 1 is a typical schematic diagram of Au wire bonding on the Cu/Low-K wafer. It involves the capillary, 1-4244-1392-3/07/$25.00 (\u00a92007 IEEE bG ~~~9 9 -21-Constant bG -6.2xlO -5.3xlO Nm K 10 10 -2Constant CG 3.IxWO 3IlxlO Nm gas constant, and G is the shear modulus given by formula (3). Material constants are listed in Table 2 (Davies et al. [15]). G(T) = aG(T/Tm)2 + bG(T/Tm)+ CG (3) Parameter H CD T OR FA CA Value 30.4ptm 38.2tm 83.8ptm 12.7ptm 40 900 2.2 Strain Rate Equation During the thermo-sonic wire bonding process, the ultrasonic vibration will heat the bond-interface to very high temperature, the substrate is usually kept at 423-493K, and the FAB also keeps higher temperature" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003132_978-981-16-0119-4_20-Figure71.5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003132_978-981-16-0119-4_20-Figure71.5-1.png", + "caption": "Fig. 71.5 Consolidated upper triple clamp", + "texts": [ + " The next step is to consolidate the possible parts along with the upper triple clamp without eliminating the functionality of the assembly the parts need to be consolidated. So, out of seven components, the lower triple clamp, upper triple clamp and fork tubes can be considered as standard parts and without eliminating the functionality steering handle, cap, the base is combined with the upper triple clamp, and it can be made as a single part. The next step is to check the manufacturability of the consolidated part using Eq. (71.1). The consolidated design is shown in Fig. 71.5. By using the additive 886 J. Jayapal et al. manufacturing-enabled part count reduction (AM-PCR) followed in [9], seven parts are reduced to five. The details of mass properties and the modified complexity factor of the consolidated design are shown in Table 71.1. From Table 71.1, the MCF value of the consolidated design is more than 44, so the design is suitable for manufacturing using AM [6]. A throttle pedal design is selected to demonstrate the proposed framework and validate the effectiveness of the new method with the Part Consolidation Candidate Detection (PCCD) algorithm used in [8]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000812_icit.2008.4608539-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000812_icit.2008.4608539-Figure5-1.png", + "caption": "Fig. 5. Dimples with/without deformation", + "texts": [ + " In order to confirm this estimation, cutting condition with N=1 was introduced. The experiment successfully generated dimples with even shape as shown in Fig.4(b). B. Experiment 2 By increasing the depth of cut, an experiment with the following cutting conditions is conducted. Diameter of cutting tool R=0.1mm Clearance angle =15\u00b0 Number of cutting blade N=1 Shape of the cutting tool: Ball-end mill Depth of cut d= 0.02mm Inclination angle = 50 \u00b0 Feed rate c=0.1 mm/tooth Dimples generated are shown in Fig.5(a). It is observed that shapes of all the dimples are deformed in the right hand side. It is conceivable that the clearance surface may collide with the surface of the work piece. In order to confirm the presumption, clearance angle of the cutting tool is increased to =30\u00b0. A dimple obtained under the modified experimental condition is shown in Fig.5(b). The deformation of the dimples is not recognized. C. Experiment 3 In order to evaluate the effects of the inclination angle conditions are adjusted as follows: Diameter of cutting tool R=0.1mm Clearance angle =30\u00b0 Number of cutting blade N=1 Shape of the cutting tool: Ball-end mill Depth of cut d=0.01mm Inclination angle = 30\u00b0 or 45\u00b0 Feed rate c=0.1 mm/tooth Two cases with the inclination angles =30\u00b0 and 45\u00b0 are tested. Obtained dimples are shown in Fig.6 and Fig.7. In Fig.6, deformations can be recognized on the dimples in the encircled regions", + " This means the target metal surface collides with clearance face. Therefore, irregular deformation is caused. In order to suppress these unfavorable phenomena, desirable range of the inclination angle is considered. Fig. 6. Dimples with deformation ( =30\u00b0) Suppose, the following cutting conditions are denoted as Diameter of cutting tool is R mm Clearance angle : Number of cutting blade is N Shape of the cutting tool: Ball-end mill Depth of cut is d mm Feed rate is c mm/tooth where is selected enough big value to suppress the collision phenomena shown in Fig.5(b). The geometrical relation requires that the inclination angle should be achieved in the region shown in Fig.9. P1 is defined as a point on the cutting edge which is cN/2 apart from the trajectory tool center point P0. Therefore, inclination angle should satisfy the following equation. 1 1 180sin cos 1 2 cN d R R (1) Based on the geometrical relation, the maximum inclination angle can be derived from the following equation. 1 180cos 1 2 d R (2) x y The experimental results in the section 2 are discussed based on a numerical model", + " The distance between Line Ln and the cutting edge is 0.1R\u00d7n. Using the numerical model, the three-dimensional trajectory surfaces of the cutting edge and lines L1 through, L5 are calculated in two cases, =15\u00b0 and =30\u00b0. The intersections of the trajectory surfaces and the workpiece surface are obtained as shown in Fig.12. These simulation data show that cutting with a tool of clearance angle =15\u00b0 causes collisions with the workpiece surface. Furthermore, the area where the collision may occur is close to the location of deformation observed in Fig.5. B. Discussion on experiment 3 Intersection of three-dimensional trajectory surface of cutting edge and the metal surface is calculated with the tool inclination angle =20\u00b0. The intersections are shown in Fig.13. The results mean that the movements of the cutting edge cause the similar deformation. It should be noted that there exits two regions in the dimpled area. In one region the cutting edge precedes the clearance surface. In the other region the clearance surface precedes the cutting edge" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001454_tmag.2007.893143-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001454_tmag.2007.893143-Figure1-1.png", + "caption": "Fig. 1. Magnetic assembly and manipulation system illustration.", + "texts": [ + " One important issue, which seems to have been largely ignored, is the controllability of the parallel magnetic assembly and manipulation. In such applications, one is most interested in controlling a number of colloidal magnetic objects using the smallest possible number of external parameters. For example, it would be convenient to control positions and movements of several particles using only a uniform external magnetic field. In this paper, we consider the simplest system of interest containing two spherical magnetic microparticles in fluid and a permanently magnetized micromagnet, as illustrated in Fig. 1. The field produced by the micromagnet is approximated as the field due to a single isolated magnetic pole. The spherical microparticles are treated as linearly magnetizable dipoles. The evolution of their positions representing the internal state of the system is described by a set of coupled first-order differential equations. The external uniform magnetic field is viewed as the system\u2019s input. Since the relation between the state and the input in this system is nonlinear and its controllability is difficult to analyze, we focus on a simpler problem of local controllability as a starting point" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000610_bfb0119393-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000610_bfb0119393-Figure2-1.png", + "caption": "Figure 2. A planar system", + "texts": [ + " In the general case the additional tasks, corresponding to a given criterion, belong to a space of dimension (v - tt - 1)(v - tt - 2) and the criteria, corresponding to a given set of additional tasks, belong to a space of dimension , (v- l ) _ 1 - (v - i t - 1)it. Thus, in general, the space of solutions generated by 2 minimizing a quadratic criterion has greatest dimension than that generated by adding tasks. The first example consists of an a c a d e m i c horizontal double-pendulum mounted on the platform; then v = 5. In the f rame ~ the position of the point O3 is given by the Cartesian coordinates ~1 and ~s and the orientation of the endeffector by the angle ~3; then It = 3. Fur thermore we suppose tha t a = b = 0 (see Fig. 2). The mat r ix J (q ) is: cosq3 - -a l s in q34 -- as s in q345 - -a l s in q34 -- as s in q34~ --a2sinq345 ] sin q3 al cos q34 -{- as cos q345 al cos q34 -{- as cos q345 as cos q345 0 1 1 1 with: q34 = q3 + q4 and q345 ---- q3 -4- q4 + qsThe control which corresponds to the minimizat ion of the quadrat ic criterion such tha t Q = E, where E is the unit ma t r ix of order 4, is given by (10) with: al cos q34 al sin q34 , f - b 1 - -~ sin q3 ~ cosq3 al cos q, - ~ sin q3 ~ cos q3 sin q3 -- cos q3 with: q45 = q4 + qs, when q4 # ~:2\" when q4 = i ~ " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000813_pvp2008-61601-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000813_pvp2008-61601-Figure1-1.png", + "caption": "Figure 1: THE CWRU BEARING TEST RIG [8]", + "texts": [], + "surrounding_texts": [ + "McFadden and Toozhy [15] assume that the envelope spectrum contains peaks only at those frequencies that are harmonics of the characteristic inner race defect frequency BPFI surrounded by modulation sidebands at multiples of the shaft rotation frequency fr given by f = m BPFI\u00b1n fr (1) where m and n are integers. The inner race defect frequency can be related to the shaft rotation frequency by BPFI = z 2 (1+ Bd Pd cos\u03b2) fr = z( fr \u2212 fc) (2) Combining (1) and (2) gives f = mz( fr \u2212 fc)+n fr (3) which shows that some spectral lines also occur at integer multiples of fr\u2212 fc, the rotation of the shaft relative to the cage. Time synchronous average signals are obtained by convolving a given signal x(t) with a comb filter. The latter is a train of N impulses of amplitude 1/N, spaced at interval Tt = 1/ ft y(t) = c(t)\u2217 x(t) (4) where c(t) is given by c(t) = 1 N N\u22121 \u2211 n=0 \u03b4(t +nTt) (5) In the frequency domain, Eqn.(4) becomes Y ( f ) = C( f ) \u00b7X( f ) (6) where C( f ) = 1 N sin(\u03c0NTt f ) sin(\u03c0Tt f ) (7) For large N, only frequencies that are exactly multiples integers of the trigger frequency ft are passed by the filter. This model 3 nloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/28/2016 Ter was successfully applied to the condition monitoring of a rolling element bearing with many spalls on the inner race. However, its implementation is impractical in industrial applications. The first limitation comes from the fact that a very large number of averages N is required to completely eliminate unwanted frequency lines, limiting us to short time duration signal, and real time machine trending could be difficult to achieve. The second limitation is more problematic: we need to know the cage rotating frequency. Getting access to the bearing cage is almost impossible in many industrial applications. Facing this difficulties, we seek a more simplified procedure to obtain a good approximation of the real time synchronous averaging of raw signals." + ] + }, + { + "image_filename": "designv11_83_0002333_s00500-021-05640-5-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002333_s00500-021-05640-5-Figure2-1.png", + "caption": "Fig. 2 Finite element model of coil spring with mesh and boundary conditions for static analysis", + "texts": [ + " In the linear static analysis, the coil spring was meshed with hexahedron solid elements at a 2 mm mesh size. To have a good quality mesh, it is important to keep the aspect ratio, warpage and skewness of the elements below the values of 5, 15 and 40, respectively (Kong et al. 2019b). A static load of 3600 N equalling to a quarter of total car weight with three passengers was applied to the bottom of the coil (Putra et al. 2017a). In addition, a fixed boundary condition at the upper coil was also imposed on the coil spring. Figure 2 depicts the meshed geometrical model of the coil spring with the boundary conditions. A strain gauge was then placed on the hotspot to measure the real-time strain histories of the coil spring during the course of the road tests. The vibration signals from the road excitations represented by the vertical acceleration of the car wheel (Kong et al. 2019b) were then recorded with an accelerometer affixed to the lower suspension arm. The strain and acceleration signals with a time span of 80 s were simultaneously collected at a sampling frequency of 500 Hz to avoid excessive data loss normally experienced by those of the standard automotive applications (Putra et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002838_012021-Figure20-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002838_012021-Figure20-1.png", + "caption": "Figure 20. Strain Distribution in E-Glass", + "texts": [ + " Analysing Testing Result of E-Glass 3.5.1. Total Deformation The Max. and Min. Total Deformation in E-Glass is 0.23872 mm and 0 mm respectively shown in Figure 18. The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. 3.5.2. Stress Distribution The Max. and Min. Stress Distribution in E-Glass is 62.053 MPa and 0.37676 MPa respectively shown in Figure 19. 3.5.3. Strain Distribution The Max. And Min. Strain Distribution in E-Glass is 0.0008937 and 0.0000069443 respectively shown in Figure 20. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.6. Analysing Testing Result of Basalt Fiber 3.6.1. Total Deformation The Max. And Min. Total Deformation in Basalt Fiber is 0.19542 mm and 0 mm respectively shown in Figure 21. The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. 3.6.2. Stress Distribution The Max. And Min. Stress Distribution in Basalt Fiber is 62.804 MPa and 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002430_012006-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002430_012006-Figure6-1.png", + "caption": "Figure 6. Force analysis of Yaw chain Figure 7. Force analysis of Pitch chain", + "texts": [ + " In the figure, dm is the mass of the moving platform; zm is the mass of the adapter; en is the external moment of the moving platform; ef is the external force of the moving platform; il m is the mass of the MEMAT 2021 Journal of Physics: Conference Series 1820 (2021) 012006 IOP Publishing doi:10.1088/1742-6596/1820/1/012006 link i ; iqm is the mass of the slider i ; In view of the simple structure of the parallel mechanism, and the Newton-Euler method is intuitive and easy to understand, this article adopts the Newton-Euler method to establish the dynamic model of the parallel rotating mechanism. The force analysis of the yaw and pitch chains is shown in Figure 6 and 7. In the figures, lidf is the force vector of the moving platform acting on the yaw and pitch links; qilif is the force vector of the slider acting on the yaw and pitch links; iN is the force vector of the support shaft acting on the yaw and pitch sliders; xyef and xzef are the components of external force ef in the xoy and xoz planes; xyen and ezn are the components of the external moment en in the xoy and xoz planes. Based on the analysis of the forces on the yaw and pitch chains, the forces on the sliders and the links are decomposed as shown in Figures 8 and 9" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000180_6.2004-5021-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000180_6.2004-5021-Figure1-1.png", + "caption": "Figure 1. Satellite Formation Coordinate System", + "texts": [ + " The formation control problem being studied in this paper is a two satellite formation. The first satellite, the leader, is in a circular orbit around a spherical earth and for simplicity, is assumed to maintain that orbit in free flight. The second satellite is in a slightly elliptical orbit but remains close to the first when compared to the overall radii of their orbits around the earth. The difference between the two orbits is enough to create useful geometries for multi-satellite viewing of various targets. The coordinate system used in this paper is illustrated in Figure 1. The x-axis extends radially from the center of the Earth through the center of gravity of the leader satellite, the y-axis is along the tangential velocity vector and the z-axis forms the right-handed system. American Institute of Aeronautics and Astronautics 3 Assuming the second satellite is in a slightly elliptical orbit but remains close to the first when compared to the overall radii of their orbits around the earth, we can use the linearized Hill\u2019s Equations to describe the legal relative motion between a leader and follower satellite2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000524_s0076-6879(08)03415-0-Figure15.7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000524_s0076-6879(08)03415-0-Figure15.7-1.png", + "caption": "Figure 15.7 Fiber optic sensor probes: (A) bare fiber, with the jacket or buffer crosshatched; (B) gel probe, with the bare end of the fiber inserted in a small volume of gel polymerized in situ; and (C) a dialysis probe, with the dialysis membrane secured over the end by anO-ringmade of Tygon tubing.", + "texts": [ + " The fluorescent transducer (in our case a variant of carbonic anhydrase) must be polymeric or otherwise immobilized by attachment to a polymer or particle since small molecules will tend to diffuse away, at the same time small molecule or ionic analytes can diffuse in. At present we make two types, neither of which is particularly novel: dialysis probes and gel probes. The dialysis probes (C) essentially have a small ( 1 ml) chamber at the end of the fiber with the fluorescent transducer inside and an opening covered with a dialysis membrane (see Fig. 15.7). The gel probe (B) has the transducer material entrapped in a porous hydrogel at the distal end. The advantage of the dialysis probe is that it is relatively easy to assemble and fill, and it can be refilled if the protein no longer responds or becomes photobleached. However, its response time is likely to be slow due to the relatively long path for the analyte to diffuse through to the distal end face of the fiber itself, and the diameter of the probe (limited by the ring which holds the dialysis tubing in place) is inconveniently large for in vivo use in blood vessels or the brain" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002434_978-3-030-59864-8_13-Figure1.2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002434_978-3-030-59864-8_13-Figure1.2-1.png", + "caption": "Fig. 1.2 Illustration of the compression apparatus with optical access. (a) Isometric view. (b) Side view in cross section", + "texts": [ + " Particle size distributions were measured with a Beckman Coulter LS 13320 laser diffraction particle size analyzer with 50% of the particle size distribution equal to 0.473\u00a0\u00b1\u00a00.006\u00a0mm (V350) and 1.163\u00a0\u00b1\u00a00.128\u00a0mm (V1000). The particle shapes are roughly spherical and have internal porosity typically in form of a single void as shown in the images of Fig.\u00a01.1. Water content in the particles was measured by drying under vacuum at 105\u00a0\u00b0C for 24\u00a0h and was nominally constant at 4.5%. Literature values for MCC report a value of the elastic modulus of 7.5\u00a0GPa [11]. The MCC particles are compressed uniaxially in a load- controlled manner by the apparatus of Fig.\u00a01.2. The apparatus consists of a pneumatic cylinder (Bimba Flat-I, Model FOS-1251.5-4GLV) to apply force to the top of a confined powder sample. A support structure vertically aligns the axis of the pneumatic cylinder to the axis of the confined sample. The sample is confined by the rectangular walls of the compression die body and confined axially by an upper and lower ram. Three walls of the compression die body are formed by a 304 stainless steel channel with height of 3.81\u00a0\u00b1\u00a00.02\u00a0cm, wall thickness of 1", + " For most tests, a carbide tipped hand held engraving tool (like McMaster Carr Part No. 1613\u00a0T2) was used to vibrate the setup in order to settle the particles. The carbide tip was held against the plate the confiner rested on for a period of at least 30\u00a0s. Then, the upper ram is installed at the upper load cell (Transducer Techniques, Model THD-5\u00a0K-T-OPT-HT) and plunger of the pneumatic piston such that it was elevated above the top surface of the particulate sample. The downward motion of the upper ram was initiated and the distance between the ram-mounted flags (Fig.\u00a01.2b) decreased as measured by the laser micrometer (Micro-Epsilon, Model 2500). Pressurization of the pneumatic piston is controlled with a LabView program and a voltage-controlled pressure regulator (Proportion-Air Model QB1SSFEE500) resulting in a constant rate of load applied by the upper ram. Force on the particulate sample increased at the preprogramed loading rate until the maximum load was achieved. The force was applied at a rate of 0.22\u00a0kN/min resulting in a varying displacement rate during compression nominally between 0", + " The maximum applied load was determined by the pneumatic cylinder\u2019s maximum pressure rating of 1.4\u00a0MPa and was connected to bottled air. The LabView program controlled the maximum applied force to 6.7\u00a0kN.\u00a0During a test, the LabView program controlled the pneumatic cylinder pressurization and recorded the laser micrometer sensor and load cell data at a rate of 1\u00a0Hz. The maximum force was held for 5\u00a0min and then released. Figure 1.3 plots the compression curve data from the compression apparatus with optical access (Fig.\u00a01.2) with the data from a cylindrical compression apparatus [10]. The data from the different compression experiments show good agreement in terms of applied stress and strain (Fig.\u00a01.3a). When plotted in terms of relative density and applied stress in a linear-log plot (Fig.\u00a01.3b), the new data from the optical compression apparatus begins at a lower tapped density. This is consistent with prior testing and is a trend with the aspect ratio (initial bed height/length scale of confiner cross section) of the powder bed [10]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002581_icees51510.2021.9383730-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002581_icees51510.2021.9383730-Figure8-1.png", + "caption": "Fig 8. Contours o f Temperature o f BLDC Motor", + "texts": [ + "987 6 4 0.999 7 5 1.000 V. Re s u l t s a n d Di s c u s s i o n A. Contours o f Temperature o f Motor To evaluate the effectiveness of CFD was this studies primary aim in order to optimize the BLDC electric motor\u2019s thermal parameters. To have a simplest geometry, to minimize the manufacturing cost was the goal of the design. The cylindrical housing was the initial geometry and for contours of temperature, the ANSYS simulation is analyzed for rotor, stator, shaft and other BLDC motor components. Fig.8 shows the temperature contours of the BLDC motor. The contours of temperature illustrate the flux of heat of these parts from winding, rotor, bearing and stator. The temperature is due to the flux linkage of the stator conductor and rotor permanent magnet flux linkage. During higher load condition the flux linkage will be more, the speed of the rotor increases and obviously the temperature of the BLDC motor also increases. Fig 9 shows the temperature increase on the housing and stator at the posterior side of the motor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003042_s11668-021-01191-x-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003042_s11668-021-01191-x-Figure5-1.png", + "caption": "Fig. 5 a Principal stress plot for Modification 1, b principal stress plot in Modification2, c principal stress plot in Modification 3 and d principal stress plot in Modification 4", + "texts": [], + "surrounding_texts": [ + "The 3D modeling of EGR cooler system is meshed using the Hypermesh software using the solid 3D, four node and tetrahedron elements with three degrees of freedom. The tetrahedral element meshing of the EGR cooler housing is shown in Fig. 3. Tetrahedral elements are used here because they fit arbitrary shaped geometries very well with their simple computations. In comparison, the hexahedral meshes are more accurate with the number of elements, since one hexahedral equal to six tetrahedral elements. However, the tetrahedral elements are best to model complex geometry domain with little distortion of mesh. While meshing the model the conditions given are linear elastic, isotropic and temperature independent. The material properties inputted are listed in Table 1 for aluminum for housing, cast iron for bracket and steel for the bolt. The chemical composition of the aluminum used is listed in Table 2. The bolts are pre-tensioned at the load of 25735N. After giving the DOF and Load, Nastran solver is used to solve the problem." + ] + }, + { + "image_filename": "designv11_83_0002382_j.medengphy.2021.03.002-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002382_j.medengphy.2021.03.002-Figure1-1.png", + "caption": "Fig. 1. Mechanical configuration of the TARPIN system. Left panel: A plastic housing embedded in the socket during socket fabrication protects the release/relock mechanism. Right panel: Electronics, motor housing, and release button removed to show mechanism. The tether to the elastomeric liner wraps around the spool (tether not shown). The USB port (not shown) is on the back side of the electronics board at a similar height to the release button.", + "texts": [ + " ystem design The system was made up of: a mechanical system termed the ARPIN (To Auto-release & Relock a PIN) that executes locking pin elease and relock and records tether length changes over time; nd an instrumented ratcheting dial that executes socket panel reease and panel retightening and records dial angle changes over ime. System components are listed in Table 1 . .1. TARPIN release/relock mechanism The main component of the mechanism is a motor assembly hat drives two stainless steel gears connected to a stainless steel pool ( Fig. 1 ). The motor system selected provides sufficient torque nd speed to assist the donning (relock) process in a timely maner while still fitting inside an embedded housing in a prosthetic ocket ( Fig. 2 ). It records motor angle at a resolution of 1729 bits er degree at the output shaft. A tether extends between the spool and the distal end of the ser\u2019s prosthetic liner. It is connected to the prosthetic liner via a hreaded custom locking pin. When the user activates the mechaism via the draw button on the posterior distal socket ( Fig", + " o fully release the tether and doff the socket, the user presses he release button. If the system is previously in partial doff mode, hen a single button push will fully release the tether. Otherwise, he user holds the release button continuously. This action actiates the solenoid to allow the user to backdrive the motor and ull the limb out of the socket. .5. Manual operation The release/relock system allows the user to manually draw in he tether in case of emergency, power failure, or malfunction. n adaptor piece is mounted to the spool shaft ( Fig. 1 ), allowing removable crank to be attached through the anterior aspect of he socket. The crank allows the user to manually turn the spool, inding the cable to draw the tether into the prosthetic socket Appendix 2). Currently, the manual crank is not part of the mechnism described above and must be carried separately. . Characterization of sensitivity and bench testing .1. Release/relock Sensitivity of the motor encoder on the TARPIN is determined sing the step counts per revolution of the digital motor encoder nd the diameter of the spool winding the tether" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002578_978-3-030-54814-8_96-Figure12-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002578_978-3-030-54814-8_96-Figure12-1.png", + "caption": "Fig. 12 General view (device) of triplex drilling pump", + "texts": [], + "surrounding_texts": [ + "The pressure created by the drilling fluid is converted through pistons (diameter\nof the pump for two cycles (working-idle) in the range of 720 degrees of the crankshaft rotation (engine speed\u20141400 rpm, third gear) is shown in Fig. 14.", + "Simulation of Loading Modes of Drilling Pump Power Drive 681", + "The results of converting the pressure over the piston into torque on the pump crankshaft, taking into account the inertia forces of the rotating and reciprocatingmasses" + ] + }, + { + "image_filename": "designv11_83_0000629_9780470612286.ch1-Figure1.1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000629_9780470612286.ch1-Figure1.1-1.png", + "caption": "Figure 1.1. A simple open structure robot", + "texts": [ + " However, this method, developed for simple open structures, presents ambiguities when it is applied to closed or tree-structured robots. Hence, we recommend the notation of Khalil and Kleinfinger which enables the unified description of complex and serial structures of articulated mechanical systems [KHA 86]. A simple open structure consists of n+1 links noted C0, \u2026, Cn and of n joints. Link C0 indicates the robot base and link Cn, the link carrying the end-effector. Joint j connects link Cj to link Cj-1 (Figure 1.1). The method of description is based on the following rules and conventions: \u2013 the links are assumed to be perfectly rigid. They are connected by revolute or prismatic joints considered as being ideal (no mechanical clearance, no elasticity); \u2013 the frame Rj is fixed to link Cj; \u2013 axis zj is along the axis of joint j; \u2013 axis xj is along the common perpendicular with axes zj and zj+1. If axes zj and zj+1 are parallel or collinear, the choice of xj is not unique: considerations of symmetry or simplicity lead to a reasonable choice" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002262_tia.2021.3058113-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002262_tia.2021.3058113-Figure3-1.png", + "caption": "Fig. 3. Maximum force profiles for various bias field values.", + "texts": [ + " Note that this model assumes linear material properties and thus Fmax is expected to be within the linear region of the materials. Non-dimensional results in this paper are obtained by setting k1 = k2 = Bmax = 1. This means that control fields and currents can be used interchangeably\u2013see (2). It is shown in [5] that when \u03b6 is allowed to vary, each of the bearing\u2019s three poles can produce force independently, allowing the bearing to produce forces anywhere within the dashed hexagonal profile of Fig. 3. However, when \u03b6 is held to a fixed value (i.e., using magnets as the external flux source for the bias field, as is assumed in this paper), the maximum force in any particular direction is limited to one of the inner contours of Fig. 3. These contours each have a different force rating Fr, which is defined as the minimum force magnitude on the maximum force profile. An example rated force vector is labeled in Fig. 3 for \u03b6 = 1 3 . The design choice of \u03b6 clearly has significant implications for the force rating of bearings with otherwise equivalent geometry (equivalent Fmax and therefore Bmax values). Using an optimal bias field (\u03b6 = 0.569) can improve the three-pole bearing\u2019s force rating by 15.5% and decrease its Ampere-turn requirements by 23.4% (see [5]) compared to the bearing bias values of \u03b6 = 0 or \u03b6 = 1 2 that are typically used. Fig. 4 highlights this result by showing the normalized force rating (blue curve) and Ampere-turn requirements (orange curve) of the three-pole bearing as a function of the bias field \u03b6", + " 5a depicts these zero-force locations as the blue x\u2019s and orange o\u2019s for two different values of \u03b6. The solid and dashed bounding regions indicate the upper limits for i\u03b1 and i\u03b2 ; any coordinates outside of the bounding regions will cause an airgap to exceed Bmax. It can be seen that for the case of \u03b6 = 0.4, three of the four solutions require pushing the airgap fields past saturation to produce the zeroforce vector, and are therefore not valid solutions. As the commanded force vector is swept across a bearing\u2019s realizable force profile (see Fig. 3), this broad variation in solution values and feasibility can result in the proposed exact force vector regulator commanding abrupt changes (discontinuities) in the coil currents, even for smooth force profiles. An illustrative example of discontinuities can be found by studying the geometrically simple case of producing a variable magnitude force along the x-axis. In this case, ~ic = Bmax k2 [B\u2032 c,1,\u2212 B\u2032 c,1 2 ,\u2212B\u2032 c,1 2 ]T is the minimum L2 solution. 1Since Bmax = k2 = 1, this is equivalent to i\u03b1 and i\u03b2 of the control current space vector: SV(~ic) = i\u03b1 + ji\u03b2 = ic,1 + aic,2 + a2ic,3, see (6)", + " During the transition period between the two solutions, large force vector errors will result, possibly causing the system to go unstable. The stability implications of this are investigated in Section V-A for an example bearing. An analysis framework is now constructed to explain discontinuities that arise in the exact force regulator\u2019s commanded bearing currents (i.e., Fig. 5c). The framework applies to any three-pole bearing design (combination of \u03b6 and Fmax) for any arbitrary desired force vector within a bearing\u2019s realizable force profile (Fig. 3). The analysis framework draws from the mature body of literature on quartic polynomials and is based on evaluating the discriminant of (4). Relevant mathematical background on how the solutions to quartics relate to properties of their discriminant can be found in numerous publications and textbooks, such as [25], [26]. The discriminant \u2206 of (4) can be calculated as (10) in terms of the desired force components F \u2032 x and F \u2032 y . \u2206 = 2916F \u2032 2 y ( \u2212 F \u2032 4 x + 24F \u2032 3 x \u03b62 \u2212 2F \u2032 2 x F \u2032 2 y \u2212 162F \u2032 2 x \u03b64 \u2212 72F \u2032 xF \u20322 y \u03b62 \u2212 F \u20324 y \u2212 162F \u20322 y \u03b64 + 2187\u03b68 ) (10) The sign of the discriminant determines the types of roots (coil current solutions) that will result", + " A positive discriminant indicates that all four solutions will either be real (all coil current solutions are valid) or complex (no coil current solutions are valid). The case of a positive discriminant can be further refined by evaluating the polynomials (11), if both polynomials are less than zero, all four solutions will be real. Finally, a zero discriminant indicates that the polynomial has a multiple root. P = \u221224F \u2032 x \u2212 216\u03b62 D = \u2212144(F \u20322 x + F \u20322 y )\u2212 4320F \u2032 x\u03b6 2 \u2212 11664\u03b64 (11) Regions in the F \u2032 x-F \u2032 y plane with different discriminant signs are depicted in Fig. 6 with the maximum profile of the threepole bearing (Fig. 3) imposed on top. Equations (10) and (11) can be used to show that when a force vector is commanded within each of these regions, the exact force regulator will have a differing number of solutions as follows: 1) Blue shaded region: four valid current solutions, \u2206 > 0 and the polynomials of (11) are negative. 2) Outside the blue shaded region: two valid current solu- tions, \u2206 < 0. The solid blue lines indicate edge cases, where \u2206 = 0. Interestingly, the region with four valid solutions (blue region) is completely defined in terms of the two generalized bearing design parameters \u03b6 and Fmax", + " 6, where the analysis framework expects a current discontinuity to exist (the lowest L2 solution is infeasible in region 2, but viable in region 1). 2) Variable magnitude force at 40\u25e6: In Fig. 7, the exact force vector regulator\u2019s control currents are computed for a variable magnitude force at an angle of 40 degrees for two different bias values (\u03b6 = 1 4 and \u03b6 = 1 2 ). The two curves extend to different force magnitudes because the two bearings have different force capability curves per [5] and Fig. 3. Once the force magnitude reaches 0.2Fmax, the currents in the bearing with \u03b6 = 1 4 discontinuously jump to a new solution. This issue does not occur for the bearing biased at \u03b6 = 1 2 . The discontinuity can be understood from the analysis framework of Section III-B based on when the force vector transitions between discriminant regions. This is depicted in Fig. 8, where the blue region 1 from Fig. 6 has been drawn on top of each bearing\u2019s maximum force profile (orange). In the case of \u03b6 = 1 4 , as the commanded force vector (shown in red) increases in magnitude, it passes out of region 1 at the force magnitude of 0", + " The possibility of this approach was revealed in the case studies of Section III-C, where bearings with \u03b6 = 1 2 did not experience discontinuities. Design guidelines, in terms of values of \u03b6, are now developed to identify and evaluate bearings where the exact force vector regulator will not command discontinuous currents. Section III-B showed that current discontinuities will not result when a commanded force vector remains in region 1 of Fig. 6. Inspection of Fig. 6 shows that the blue region 1 grows with \u03b62, while both region 1 and the red hexagon (and therefore a bearing\u2019s maximum force profile\u2013see Fig. 3) are proportional to Fmax. Consequently, as \u03b6 increases, the blue region will grow relative to the force hexagon (and maximum force profile). This is depicted in Fig. 11 for various values of \u03b6. Key ranges for \u03b6 can be identified as follows: 1) \u03b6 = 0: no current discontinuities, as the bearing\u2019s force profile will always reside in region 2 (the same number of solutions are always feasible), provided that the solution selection algorithm presented in [14] is utilized; 2) 0 < \u03b6 < 0.287: current discontinuities will occur within the bearing\u2019s force profile as it spans regions 1 and 2; 3) 0", + " The motion regulator can be implemented as a PID controller following recommended practices for magnetic bearings, such as those provided in [4], [23]. The output of the outer loop PID controller is a force vector command that passes through a saturation block before being inverted into the i\u03b1 and i\u03b2 current commands of (6). The exact force vector regulator of Section III makes up the force inversion and saturation blocks of Fig. 12. 2Recall that the rated force profile is a circle with a radius equal to the inner radius of the maximum force profile\u2013see Fig. 3. Authorized licensed use limited to: Carleton University. Downloaded on June 04,2021 at 02:56:52 UTC from IEEE Xplore. Restrictions apply. 0093-9994 (c) 2021 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. As previously discussed, calculating the control currents for a desired force vector requires computing the roots of a fourth order polynomial in the form of a depressed quartic (4)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001420_ukricis.2008.4798952-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001420_ukricis.2008.4798952-Figure2-1.png", + "caption": "Fig. 2. Circuit diagram of a DC motor", + "texts": [ + " The control objective is to make the beam of the TRMS track a predetermine trajectory. Fig. 1 shows the TRMS considered in this investigation. The crude dynamic model supplied by the manufacturer has been improved [20] and the DC motors are simulated with respect to the corresponding equations [20]. The TRMS possesses two permanent magnet DC motors, one for the main and the other for the tail propelling. The motors are identical with different mechanical loads. The dynamic model of the main motor, as shown in Fig. 2, is presented in (1) to (5). ( )1av v av av av av di V E R i dt L = \u2212 \u2212 (1) vvavav kE \u03c9\u03d5= (2) ( )1v ev Lv mr v mr d T T B dt J \u03c9 \u03c9= \u2212 \u2212 (3) avvavev ikT \u03d5= (4) vvtvLv kT \u03c9\u03c9= (5) The dynamic model of the remaining parts of the system in vertical plane is described in (6) to (8). In (6), the first term denotes the torque of the propulsive force due to the main rotor, the second term refers to the torque of the friction force, and the torque of gravity force is shown in the third term. [ ] v vvvfricvvmv J CBAgTFl dt d \u03b1\u03b1\u03c9 sincos)()( , \u2212\u2212+\u2212 = \u03a9 (6) where, ttstr t lmm m A ) 2 ( ++= , mmsmr m lmm m B ) 2 ( ++= , ) 2 ( cbcbb b lml m C += if 0 ( ) if 0 fvp v v v v v fvn v v v k F k \u03c9 \u03c9 \u03c9 \u03c9 \u03c9 \u03c9 \u03c9 \u23a7 \u2265\u23aa= \u23a8 <\u23aa\u23a9 (7) v v dt d \u03a9= \u03b1 (8) An accurate model of friction has been considered to cover viscous, coulomb and static friction types" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002004_50015-9-Figure12.14-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002004_50015-9-Figure12.14-1.png", + "caption": "Figure 12.14 Snowboard binding utilizing: thermoplastic elastomer (Hytrel) \u2013 ankle strap A, spoiler B, ratchet strap G; nylon (Zytrel) \u2013 side frames D and H, base and disk F, top frame J; acetal homopolar (Delrin) \u2013 strap buckles C1 and C2 (courtesy of Fritschi Swiss Bindings AG and Du Pont UK Ltd).", + "texts": [ + " These release instantly under too severe twisting, while providing foot stability and control. An acetal (highly crystalline) polymer, which is tough, has a low glass transition temperature Tg and good fatigue resistance, is used for the locking bar, heel release lever, heel block and front swivel plate, while glass-reinforced nylon is used for the front block and two base plates. These polymers are tough at low temperatures and both UV and moisture resistant. Other release bindings employ titanium and plastic components. Figure 12.14 shows the variety of different materials used in the binding for snowboards. Finally, ski poles are tubular with high specific stiffness and toughness. They are usually made of CFRP or CFRP\u2013GRP hybrids. 12.3.6 Archery Every schoolboy having read, or watched, Robin Hood is familiar with the longbow, probably knows it was made from yew and that considerable force (35\u201370 kgf) was needed to draw it. When drawn, the strain energy stored in the \u2018string\u2019 and the two (upper and lower) limbs of the bow is transferred to the arrow, which accelerates up to a velocity of 50 m s\u22121" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002796_j.matpr.2021.04.285-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002796_j.matpr.2021.04.285-Figure3-1.png", + "caption": "Fig. 3 (continued)", + "texts": [ + " 2 (continued) Our main aim is to build a gear that incorporates mistakeproofing technology to reduce assembly errors in gear trains and gear boxes. While arresting the DOF in the axial direction, a curve-shaped cut on the face of the tooth surface has been achieved, and the same space has been extruded on the bottom land, which is perpendicular to the gear\u2019s axis. The removed area on the face of the tooth in the first gear meshes with the extrusion made in the bottom land on the second gear during meshing which is shown in Fig. 3 (a) & (b). At the same time, the first gear\u2019s extru- sion meshes with the second gear\u2019s removed region on the tooth face. One DOF in the axial direction of the gear was arrested, slippage in the gear was reduced, and misalignment due to gear shaft inclination was reduced with this type of arrangement. The groove cut into the tooth must be higher than the gears\u2019 contact line. The extrusion must be positioned under the contact line. On each tooth of the gear, an extrusion and groove were formed. To reduce noise and increase smoothness, the extrusion and cut can use either concave or convex profiles" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002838_012021-Figure23-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002838_012021-Figure23-1.png", + "caption": "Figure 23. Strain Distribution in Basalt Fiber", + "texts": [ + " The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. 3.6.2. Stress Distribution The Max. And Min. Stress Distribution in Basalt Fiber is 62.804 MPa and 0.37819 MPa respectively shown in Figure 22. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.6.3. Strain Distribution The Max. And Min. Strain Distribution in Basalt Fiber is 0.00074223 and 0.0000057032 respectively shown in Figure 23. 3.7. Analysing Testing Result of Carbon Fiber 3.7.1. Total Deformation The Max. And Min. Total Deformation in Carbon Fiber is 0.68523 mm and 0 mm respectively shown in Figure 24. The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.7.2. Stress Distribution The Max. And Min. Stress Distribution in Carbon Fiber is 97.378 MPa and 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002864_0954406221995852-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002864_0954406221995852-Figure8-1.png", + "caption": "Figure 8. Planar four-bar mechanism.", + "texts": [ + " Figure 6 displays an improper selection of quasi-coordinates leading to incorrect results. Similarly, Figure 7 shows the effect of singularity on the behavior of generalized coordinates of the system. In more complex systems, it is difficult to find a comprehensive quasi-space to keep the dynamic and kinematic equations from violation during the entire simulation time. As such, switching between different spaces is proposed. The second example of application is the planar four-bar mechanism (PFBM) which is represented in Figure 8. This multibody system consists of four rigid bodies, including the ground, and four ideal revolute joints. The dimensional and inertia properties of each body are listed in Table 3.28 Unlike the previous one, this example contains holonomic constraints to prove the performance of the proposed method in both holonomic and nonholonomic constrained systems. Maggi\u2019s equations, as with Kane equations, have been applied to holonomic systems by enforcing constraints only at the velocity level, which raises theoretical and practical issues with the resulting ODE" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001116_coginf.2008.4639193-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001116_coginf.2008.4639193-Figure6-1.png", + "caption": "Fig. 6. Representation of a (a) invalid and (b) valid configurations for two milestones.", + "texts": [ + " Number of milestones is the main parameter for three metrics: query time, path simplification time and number of sets we have to generate in order to find a path. Red lines come from original PRM and, as can be seen in scenes, contain redundant nodes. The blue line is the simplified path after removing redundant nodes. The first important parameter is the effect of number of milestones on number of sets of configurations to produce valid configurations for the path. A valid configuration is a configuration which neither the points nor the edges connecting these points have collision with obstacles. Fig 6 shows a typical example. In Fig. 6 (a), with only two milestones, the chance to fmd a path is low because the nature of generating milestones is random. So we stuck at milestone 1 because the line connecting milestones 1 and 2 is blocked by obstacle, hence this configuration is invalid. In Fig.6 (b) we generate two milestones again, but this time the placement of ..\u2022. . ... .. ... (a) .. .* \u2022 (a) *. *. Fig. 4. Scene2 (a) path resulted from our algorithm and (b) path resulted from We milestones is such that we can find a path without any collision. Here two sets of configuration to produced to reach a valid configuration. Now if we have 10 milestones we will be able to find a path with high chance and the result is having only one set of configuration which is valid. As can be seen from diagram, number of milestones has inverse relation with number of configuration sets" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000774_gt2007-28172-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000774_gt2007-28172-Figure1-1.png", + "caption": "Figure 1. Schematic of the journal bearing with diamond knurled stator surface. (a) z-cross section of the bearing. (b) Foot print of a knurl.", + "texts": [ + " In this paper we shall numerically investigate the flow structure and the pressure generated in a journal bearing with diamond knurled stator surface. Three numerical models will be developed and compared. The three-dimensional NavierStokes equation with realistic boundary conditions will be solved. The total forces on the shaft will be presented and compared to those from a smooth bearing of the same clearance. An axial cross section of the knurled journal bearing studied in this paper is given in Figure 1a. The foot print of the knurls is shown in Figure 1b. Figure 2a is a three dimensional view of the bearing. The rotor of the bearing has a smooth surface while the diamond knurls on the stator surface are in a staggered pattern. The radius of the shaft is R0=12.7 mm (0.5 ded From: https://proceedings.asmedigitalcollection.asme.org on 12/25/2018 Terms of Use: h inch). The clearance of the bearing, C=25.4 \u00b5m (0.001 inch). The dimensions of the knurl in the circumferential (AB in Figure 1a) and axial direction are Wc=0.794mm (0.03125 inch) and Wa=1.588 mm (0.0625 inch), respectively. The depth of the knurl (DE in Figure 1a) is Wd=0.396 mm (0.015625 inch). Figure 2a shows the three-dimensional embodiment of the whole bearing model. In this model, flow of lubricant in all the knurls and the fluid film are simulated. The rotor surface, the stator land surface, the sidewalls of all the knurls, and the bottom wall of the knurls, all have non-slipping wall condition. The surface of the rotor rotates at 5000 rpm while all other walls are stationary. On the axial end surfaces, where the fluid domain opens to the surroundings, uniform pressure conditions 2 Copyright \u00a9 2007 by ASME ttp://www" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000917_j.ijimpeng.2007.10.009-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000917_j.ijimpeng.2007.10.009-Figure2-1.png", + "caption": "Fig. 2. Illustration of variables used in defini", + "texts": [ + " Accordingly, the normal velocities on the inflow and target boundary surfaces are vz \u00bc 1; 0 r < \u00f01\u00fe b0\u00detan\u00f0q\u00de; z \u00bc b0; vz \u00bc 0; 0 r; z \u00bc 0; (21) respectively. For potential flow, the additional boundary conditions in Eq. (7) are satisfied. The free boundary has unknown shape r \u00bc R\u00f0z\u00de and particle velocity v \u00bc \u00f0vr ; vz\u00de. Therefore, the normal velocity n$v on this boundary is unknown. Expressions for this velocity are established as follows. At time t > 0, a particle X initially on the un-deformed boundary is found at the position x \u00bc x\u00f0X; t\u00de of the deformed boundary (Fig. 2(a)). In terms of the dimensionless variables X \u00bc X=vpt and x \u00bc x=vpt (Fig. 2(b)), there is the corresponding relationship x \u00bc x\u00f0X\u00de: (22) The dimensionless velocity of a surface particle X is v \u00bc 1 vp vx vt \u00bc v vt tx X=vpt \u00bc x vx vX $X: (23) In Appendix A, it is shown that the distance l along the deformed boundary from the leading edge X \u00bc 0 of this boundary to the surface point x\u00f0X\u00de (Fig. 2(b)), is jXj. As the vector X has a constant ng the properties of the free boundary. direction defined by the half apex angle q, Eq. (23) and the relation l \u00bc jXj gives v \u00bc x vx vl l: (24) As vx=vl \u00bc t is a tangent vector to the boundary r \u00bc R\u00f0z\u00de at x, this relation gives n$v \u00bc n$x. The components of n are nr \u00bc \u00f01\u00fe \u00f0R0\u00de2\u00de 1=2 and nz \u00bc R0\u00f01\u00fe \u00f0R0\u00de2\u00de 1=2, and those of x are xr \u00bc R\u00f0z\u00de and xz \u00bc z, which gives the normal velocity on the boundary n$v \u00bc R R0z 1\u00fe R0 2 1=2 : (25) The radius of the free boundary surface R\u00f0z\u00de must satisfy the condition of continuity of mass of the deformed volume of the cone and that of arc length of the free boundary (Fig. 2(b)). These conditions can be expressed as 2p Z b 0 Z R\u00f0z\u00de 0 rrdrdz \u00bc p 3 \u00f01\u00fe b\u00de3tan2\u00f0q\u00de; Z b 0 1\u00fe R0 2 1=2 dz \u00bc 1\u00fe b cos\u00f0q\u00de; \u00f026\u00de respectively, where b \u00bc b0 and r \u00bc 1 for incompressible flow. Further, the free boundary is assumed to be tangent to the target surface z \u00bc 0 and to the undisturbed cone surface at z \u00bc b0. These conditions can be expressed as R0\u00f00\u00de \u00bc N; R\u00f0b0\u00de \u00bc \u00f01\u00fe b0\u00detan\u00f0q\u00de; R0\u00f0b0\u00de \u00bc tan\u00f0q\u00de; (27) respectively. The shape of the free boundary of the deforming projectile is defined by the equation r \u00bc R\u00f0z\u00de", + " The stress and velocity fields in the deforming apex of a conical projectile in axi-symmetric flow on a flat, rigid and friction-free target surface differs from those of a cylindrical projectile under corresponding flow conditions. For an impacting cone, self-similar flow persists from the start of impact, whereas for a cylinder the flow becomes stationary after an initial phase of transient flow. The mass of the undisturbed cone is self-similarly redistributed so that particles at the apex will be at the circle with radius R0\u00f0t\u00de \u00bc R\u00f00; t\u00de as illustrated in Fig. 1. By Eq. (24) and Fig. 2, the radial velocity of the front of the displaced mass is found to be vr \u00bc R0vp vp\u00f0tan\u00f0q\u00de \u00fe 1=cos\u00f0q\u00de\u00de vp, if 0 < q < 90 . For an incompressible fluid, this can be seen in Figs. 3 and 4, where vr\u00f0R0;0\u00de \u00bc R0 > 1. This radial velocity differs from that of a cylindrical projectile, where the radial velocity on the target surface never exceeds vp. The discrepancy between the results of the selfsimilar model and those of the Autodyn simulations with projectile material D (incompressible fluid) observed in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001795_upec.2007.4469107-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001795_upec.2007.4469107-Figure2-1.png", + "caption": "Figure 2. (a) Standard DTC Control Scheme basic vector diagram. (b) Vector diagram showing the stator voltage vector V2 dq components", + "texts": [ + " The fuzzy stator flux controller loop provides the input Vx and the fuzzy torque controller loop provides the input Vy to the Space Vector Modulator (SVM). The SVM combines the Vx and Vy inputs with the estimated stator flux position and the converter input voltage to modulate the required output voltage vector in the Matrix Converter. UPEC 2007 - 1115 3 DTC BASED FUZZY CONTROLLER OPERATION PRINCIPLE The proposed system takes into account the DTC operating principle [2], and the overall Induction Motor behaviour and uses an AI controller that directly performs the motor control functions. Figure 2(a) shows the situation presented when the flux space vector is in sector 1, presenting the effect that each stator voltage vector, if applied in this situation, would have on the torque and flux. Figure 2(b) shows in detail the effect that the selection of a particular voltage vector, voltage vector 2 in this case, would have, taking into account the present stator flux space vector. In this situation the quadrature component of space voltage vector 2 will increase the torque, while its direct component will increase stator flux. The linkage between stator flux space vector position and the actual effect produced by the stator voltage vector are the reasons why it is necessary to decouple the torque and stator flux feedback loops, in such a way that the selection of the next space voltage vector to be applied to the machine is influenced in the quadrature component by the torque error and in the direct component by the flux error" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000030_acc.2006.1657170-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000030_acc.2006.1657170-Figure2-1.png", + "caption": "Fig. 2. The two-link robotic manipulator.", + "texts": [ + " wi,j g (k) indicates the weight coefficient from the neuron j in the gth layer to the neuron i in the (g + 1)th layer. Ui(k) is the input of NN. \u03b8 (i) g (k) is a positive constant for the threshold of the neuron i in the (g + 1)th layer. As the additive gain perturbations defined in the formula (14), the outputs of NN are set in the range of [\u22121.0, 1.0]. Finally, it should be noted that the inputs of NN are e(k+ 1) of (13) and the outputs of NN are N(k) of (15). In order to verify the effectiveness of proposed method, a numerical example is given. Consider a two-link robot manipulator in Fig. 2 with the system parameters as: link mass m1 = 10[kg], m2 = 4[kg], lengths l1 = l2 = 0.2[m], joint positions \u03b81[rad], \u03b82[rad]. The system matrices are given as follows, where T = 0.01. A = \u23a1\u23a2\u23a2\u23a3 1.00 0 0.01 0 0 1.00 0 0.01 \u22120.09 0 0.90 0 0 \u22120.07 0 0.86 \u23a4\u23a5\u23a5\u23a6 , B = \u23a1\u23a2\u23a2\u23a3 0 0 0 0 0.01 0 0 0.01 \u23a4\u23a5\u23a5\u23a6 , Dk = [ 3 1 2 1 1 3 1 2 ] , Ek = 1, N(k) = \u23a1\u23a2\u23a2\u23a3 N1(k) 0 0 0 0 N2(k) 0 0 0 0 N3(k) 0 0 0 0 N4(k) \u23a4\u23a5\u23a5\u23a6 . It should be noted that N1(k), N2(k), N3(k) and N4(k) are the output of NN. The initial conditions are \u03b81(0) = 2, \u03b82(0) = 2, \u03b8\u03071(0) = \u03b8\u03072(0) = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000828_3.43615-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000828_3.43615-Figure6-1.png", + "caption": "Fig. 6 Longitudinal variation of control force.", + "texts": [ + " The subcript \u2014 c on a force or moment will indicate the force or D ow nl oa de d by U N IV E R SI T Y O F M IC H IG A N o n Fe br ua ry 9 , 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /3 .4 36 15 (C0_;(C0_ Fig. 5 (see fig.3) Effect of e.g. location on trim angle of attack and control surface deflection. moment of the configuration less that of the control surface. The lift and moment or the configuration are then L = L_c + Lc M = M-c + I (2) where lc is the distance between the center of application of the control force and the point about which the moment is written. The force and moment system is shown in Fig. 6. The force and moment system less the control surface, transferred to the neutral point of the configuration less the control surface (NP)-.CJ is shown in Fig. 7 in nondimensional form. The nondimensional distance between the stated neutral point, and the location of the control surface force is designated as //. The lift coefficient due to the control surface CLc is a function of angle of attack and control surface deflection. The value of the force of the control surface must be such as to balance the moment about (ATP)_C" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001594_wisp.2007.4447620-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001594_wisp.2007.4447620-Figure2-1.png", + "caption": "Fig. 2 Hydraulic Cylinder transducers has also additional advantages:", + "texts": [ + " / _ _ _ A X_ _ X J Transducers usually employed to convert hydraulic energy C into electrical signals are the so called pressure transducers. ...... Main advantage of those transducers is that they are low cost, o 20 40 U0 80 1 00 1 20 1 40 easy to install, reliable and provide good precision. Forces in chambers A and B For our application GEMS 2200 RGGl00AA3UA (Gems Sensors & Controls) with an operating range between 0 - 100 bar were selected to be placed at the input of chambers A and B(see Fig. 2 and 3.) for each one of the vertical joints of the robot, to measure the differential pressure that exists between \u00b0_i ____________h_-_A_A _A_A the areas A and B, and, as a result, to measure the static force 0 20 40 60 80 1 0 1 20 1 40 that the rod is exerting by using the following equation: Fp To notice the effect in the force measurement in respect to * Temperature effects are minimised when friction the pressure difference Fp, the cylinder was subjected to a measurements are made. Temperature variations * * * * r 1 * 1 * * 1 1 - rr * 1 ~~~~~~~~~~~~~~~~~~~~~~~~~Friction mnodel for the vertical jointchanges dramatically the oil viscosity, but differential F o easu rem en in im ises th at effeet" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001001_aim.2008.4601625-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001001_aim.2008.4601625-Figure1-1.png", + "caption": "Fig. 1 A Basic 3 DOF leg.", + "texts": [ + " The paper is organized as follows: section 2 presents basic four legged mechanism and kinematics equations. Section 3 briefly introduces the gait generation types. Section 4 shows the simulation results. Section 5 summarizes the results and gives the conclusion. In general, for a machine to walk, at least 2 Degrees of Freedom (DOF) leg is required. One of them is for lift up and land movement of the leg and the other one is for swing the leg back and forth. Most of the quadruped robots have three DOF legs, because of more maneuvering capability. A basic three DOF mechanism is shown in Fig. 1. From Fig. 2, basic position vectors of each leg can be expressed in global coordinate frame: ),,( Bi o Bi o Bi o i o zyxB = ),,( Ti o Ti o Ti o i o zyxT = , 4,3,2,1=i (1) From Fig. 1, We have: 2 2 1 3 3 2 2cos( ) cos( ) ( ) ( )o o o o i i i i i Ti Bi Ti Bia a a x x y y\u03b1 \u03b1+ + = \u2212 + \u2212 978-1-4244-2495-5/08/$25.00 \u00a9 2008 IEEE. 1 Proceedings of the 2008 IEEE/ASME International Conference on Advanced Intelligent Mechatronics July 2 - 5, 2008, Xi'an, China 4 3 3 2 2sin( ) sin( )i i i i i ia a a h\u03b1 \u03b1+ + = (2) 1tan( ) o o Ti Bi o o Ti Bi y y x x \u03b1 \u03ba \u2212= \u2212 where 1=\u03ba for 2=i or 3, 1\u2212=\u03ba for 1=i or 4 and where ih denotes the initial height of i oB . Maximum lateral stretch of a leg can be expressed as; We use ADAMS for our simulations and a scene from this program can be seen from the Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000931_s10778-007-0080-0-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000931_s10778-007-0080-0-Figure1-1.png", + "caption": "Fig. 1", + "texts": [ + " The bifurcations of the equilibrium states of an inverted simple pendulum with angular and linear eccentricities of the follower force were analyzed in [4, 5]. The springs of the pendulum had characteristics of the same type: hard, soft, or linear. In what follows, we will analyze how the difference in the types of the spring characteristics at the upper end and at the point of suspension of an inverted simple pendulum affects the bifurcations and stability of its equilibrium states. 1. Problem Formulation. An inverted pendulum with the upper end attached to a spring is schematized in Fig. 1. Here Qc is the elastic force developed by the horizontal helical spring of stiffness c. This spring and the spiral spring at the joint O may be hard (h), soft (s), or linear (l). The moment of the viscoelastic forces about the point O is denoted by M1. The angle between the pendulum and the vertical is denoted by 1 ( 1 0 when the springs are relaxed). The mass of the material point \u00c01 is m1. The rod OA1 is imponderable and rigid, its length is l1. The follower force \u00d0 has linear eccentricity 0 and makes an angle 3 k 1 with the vertical, where is the angular eccentricity and k is the orientation parameter" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002381_j.matpr.2021.01.637-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002381_j.matpr.2021.01.637-Figure5-1.png", + "caption": "Fig. 5. Areas of application of forces.", + "texts": [ + " S pn \u00bc sin p 2 a sin 1 cosa R Ri pmcosa ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi x2 \u00fe y2 p cosa \u00f03\u00de where: a: Pression angle; R: Primitive radius; Ra: Head radius; Ri: Radius of any point taken on the profile with coordinate (x, y); x, y: Coordinates of any point on the tooth profile; m: Module of the tooth; P: Circular step S pn : Normalized position; pn: Basic normal pitch: pn=pmcosa; Z1, Z2: Number of teeth of wheel 1 and wheel 2 S 0 0 2=Pn andS 0 0 1=Pn: normalized positions on the action line of the start and end of real contact. The calculation of the forces on the profile of the tooth evolved through several steps listed in the above flowchart (Fig. 6). Contact forces can be applied onto a point, a line or a surface. SOLIDWORKS allowed only the application of forces on a surface. So small areas of a few tenths of a square millimeter were created to apply these forces. They were created as small so as not to deviate too much from the real conditions (Fig. 5). Five contact zones on the tooth profile where the forces will be applied have been selected. With regard to the imposed displacements, this involves embedding on the lower surface of the tooth as shown in Fig. 7. The mesh under SOLIDWORKS was generated using given meshing parameters and, Fig. 8 shows an example of a mesh used. Once the mesh is executed the calculation of stresses and deformations is launched. For the study of the influence of the torque, all the other parameters were fixed and the torque was varied (5 Nm, 10 Nm) referring to the torques chosen for the experimental tests carried out from 2", + " 11d). This discrepancy is explained by the fact that the wheel, having a higher number of teeth, has a larger radius. So for the same torque transmitted it undergoes less effort. It is noticed that the contact stresses are maximum at the primitive point (S = 0). This can be explained by the fact that the load distribution factor is maximum at this point (5). The bending stresses at the root have the same appearance (Fig. 11). Comparing the same stresses when the forces are applied at the head (Fig. 5) zone (1) and towards the root zone (5), the bending stresses are higher in the case where the point of contact is located at the head than towards the root because the load distribution factor is almost the same at these two points while the radius is greater at the head in (1) than at the root in (5). In addition, the bending stress increases as the tooth wears out, which makes sense since wear causes a loss of material which weakens the tooth. For example, for a tooth with no wear, the bending stress on the pinion level reaches a maximum of 6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002732_j.engfailanal.2021.105451-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002732_j.engfailanal.2021.105451-Figure2-1.png", + "caption": "Fig. 2. FEM model of bucket wheel excavator SchRs 630, boundary conditions and loads.", + "texts": [ + " It should be noted that the units stress units in this figure, and all similar figures hereinafter, are expressed in kN/cm2, which corresponds to 10 MPa, e.g. 1.52e01 kN/cm2 would be 152 MPa (see Fig. 1). Loading of the structure includes all the weights (dead load) and estimated workload in the form of vertical, lateral and frontal force (overall digging force), according to valid standards. Reactions in bearings on pylons are calculated based on kinematic models of bucket wheel excavator. Loading of pylons is shown in Fig. 2. The results of the numerical analysis, in the form of the deformation and stress field, are shown in Figs. 3 and 4. Displacement of bearing balls is shown in Fig. 5. It can be seen that load distribution by bearing circumference is uneven. A detailed view of the locations 1, 2, 3, 1\u2032, 2\u2032 and 3\u2032 can be seen in Fig. 6. Locations 1 and 1\u2032 correspond to the connection between the base and the boom structure, locations 2 and 2\u2032 are located near the centre of the base and locations 3 and 3\u2032 are located in the base supports (which explains why the displacements are lowest in these locations)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000534_s0065-2458(08)60221-1-Figure57-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000534_s0065-2458(08)60221-1-Figure57-1.png", + "caption": "FIG. 57. Description of a vortex chamher amplifier.", + "texts": [ + " This cylindrical chamber represents a very interesting nonlinear device : the pressure drop is not only a nonlinear function of the flow-rate; it depends also on the direction of the flow at the inlet (omission of the wedge is possible and would make it easier to speak of one inlet), and it is this second phenomenon which is made use of in the vortex chamber amplifier. Therefore, the necessary 225 action to clfect amplification is to provide a change of the direction of flow, and it makes no difference how this is done, It is then also clear that a great many combinations of amplifiers and nonlinear elements such as the above mentioned cylindrical unit are possible. They easily allow histability or tristability (the latter leading to the use of both senses of rotation) to be introduced. As far as an arrangement based on momentum control as shown in Fig. 57 is considered, an elementary description is possible. Following ref. 66, one starts with the tangential velocity v t a t the entrance just after mixing of power and control stream: IJ, is the velocity of the tangentially injected control fluid, Qc the control and 0, the power flow-rate. As there arc no gravit,y effects the Bernoulli's equation reduces to (58) P P pa + ; I!2 = p , + 2 1h2. 2 Neglecting the radial velocity components and observing that v.r = constant (59) (conservation of angular momentum) the following expression for the controllable part of the pressure drop inside the cylinder is obtained: 226 DIGITAL FLUID LOGIC ELEMENTS Introduction of thc control power allows to define a power amplification factor pp = -T----- " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001395_detc2008-49180-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001395_detc2008-49180-Figure1-1.png", + "caption": "Figure 1. A Stephenson six-bar linkage", + "texts": [ + " The concept of sheet and side of JRS offers an excellent model to explain and understand the interaction between loops in six-bar linkages [1, 8, 10, 34] and hence the formation of branch, sub-branch, and full rotatability of Stephenson type linkages. A six-bar linkage may have up to six branches [1] and each branch may have one or more sub-branches. In a sub-branch, an input value corresponds to one and only one linkage configuration. A branch with full rotatability contains only one sub-branch and in such a branch, each input value corresponds to a unique linkage configuration. A Stephenson six-bar linkage (Figure 1) contains a singleDOF four-bar loop A0ABB0A0 and a two-DOF five-bar loop A0ADCC0A0. The branch identification [1, 10] is briefly explained with the examples in Figure 2 and Figure 3. Consider the relationship (called I/O curve) between and, which are two joint variables in the four-bar loop (Figure 1). Figure 2 shows the I/O curves in the five-bar JRS (the shaded loaded From: https://proceedings.asmedigitalcollection.asme.org on 12/15/2018 Terms of Use area) of a Stephenson six-bar linkage. The I/O curve intersects the boundary of the JRS at branch points (labeled as 1, 2, ---, 6 in Figure 2). The JRS sheet has two sides. The branch points divide the I/O curve into segments and each segment of the I/O curve within the JRS represents a sub-branch. Thus, Figure 2 shows that the linkage has three branches represented by segments 1-2, 3-4, and 5-6", + " Thus, the discriminant [41] of the input and output equation should vanish, which leads to a polynomial equation in terms of the input parameter only. The following equations must be satisfied, 0),( oif , and 0)( i . (1) The existence of dead-center positions is affected by the choice of the input joint and is irrelevant to the choice of the output joint. Therefore, finding the dead-center positions for Stephenson linkages can be treated in three categories. (1) In the first category, the input is given through a joint in the four-bar loop (Figure 1). The dead center positions exhibited in the I/O curve of the four-bar loop and the branch point with the JRS boundary are the dead center positions of the Stephenson six-bar linkage. This case has been well treated [1]. In the other two categories, the input is given through a joint, such as C0 (or D) and C (Figure 6), not in the four-bar loop. (2) If the joint at C0 is used as the input joint, the singularity occurs when links A0A, B0B, and CD intersect at a common point [1]. In order to demonstrate the generality of the proposed method for any input condition disregarding the linkage structure type, the singularities will be determined through the input vs", + " (7) can be re-arranged and the discriminant similar to Eq. (9) can be obtained to find the 8 values at the dead center positions. It is noted that with Maple software [42], the discriminant of the lengthy equation and the numerator and denominator can be automatic obtained and solving the polynomial equation is straightforward. The extraneous roots of Equation (10) can be eliminated with Equation (11). Each real root of the polynomial equation represents a dead center position. Example 1: The dimensions of Stephenson six-bar linkages (Figure 1 and Figure 6) are given below. 5.31 a , 82.22 a , 0.43 a , 5.24 a , 0.55 a , 23.56 a , 2.37 a , 0.18 a , 0.59 a , 04.310 a , 87.36 , 0.70 , 0.46 , 67.43 . 4 Copyright \u00a9 2008 by ASME : http://www.asme.org/about-asme/terms-of-use Downloa The branch points and dead center positions when the input is given through 2 of the Stephenson linkage in Figure 1 are listed in Table 1 and shown in Figure 8 and Figure 9. If the input is given through 8 (or 7 ) of the Stephenson linkage in Figure 6, the resulting equations (Equation (10) and Equation (11)) for the dead center positions are a 66th and 6th degree polynomial in terms 8t of (or 7t ). The real solutions are the dead center positions. The resulting dead center positions with 8 and 7 as the input are listed in Tables 3 and 2 respectively and their corresponding points on the 3 vs. 2 curve are shown in Figure 8. One must note the corresponding angular transform relationship between these two inversions of Stephenson linkages, which are indicated in Tables 1, 2, and 3. The superscripts 1 and 6 on the angular displacements represent the corresponding values in Figure 1 and Figure 6 respectively with the same dimensions. A geared five-bar linkage (Figure 4) can also be considered as a multiloop linkage consisting of a five-bar loop and a gear loop. With an imposed gear train or a linear constraint, the five-bar linkage becomes a single degree of freedom mechanism. The following general gearing relationship [6] can be used to express the gear constrain ))(1()( 5054054 n (14) ded From: https://proceedings.asmedigitalcollection.asme.org on 12/15/2018 Terms of Use: where 4 and 5 are the joint variables at joints D and E, 40 and 50 are the joint displacements of the reference gear position, and n is the gear ratio", + " The following discussion on full rotatability takes the effect of using different input joints into consideration. 4.1 Input Given to a Link of the Four-bar Loop or Geared Train: The full rotatability of geared five-bar linkage (Figure 4) with the input given in the geared train has been discussed by Ting [2]. The linear input-output (I/O) constraint must lie within the JRS of the five-bar loop and a unit geared ratio is required [2] unless the host five-bar contains no uncertainty singularity (or dead center positions). For a branch of a Stephenson linkage (Figure 1) to have full rotatability, the following conditions must be satisfied. 1. The four-bar loop is a Class I chain and the input joint must connect the short link of the four-bar loop. This is to assure that the input link of the four-bar loop has full rotatability. 2. The I/O curve of the branch must stay within the JRS of the five-bar loop. In other words, no branch point exists in the branch. An example is shown in Figure 7, which shows that the linkage contains four branches. Branches 3 and 4, which contain branch points, do not have full rotatability", + " In such an input condition, once the branch is determined, a branch that contains no dead center position has full rotatability. This can be done by mapping the corresponding output value to the four-bar I/O curve or the linear gear constraint (Figure 8 and Figure 5). The full rotat Figure ability identification is illustrated with the following two examples. aded From: https://proceedings.asmedigitalcollection.asme.org on 12/15/2018 Terms of Us Stephenson six-bar linkage: With the same dimension as in Example 1, branch points 1, 2\u2026 8 with the input given through 2 of the Stephenson linkage in Figure 1 are listed in Table 1 and shown in Figure 8. The dead center positions with the input give through 7 or 8 of the Stephenson linkage in Figure6 are listed in Table 2 and Table 3 respectively (Figure 9). Branch identification [1, 10]: With the branch identification discussed early, as seen in Figure 8, the I/O curve of 2 vs. 3 is divided into segments by the JRS. There are eight branch points 1, 2\u2026 8. Each branch may contain 0 or 2 branch points. Thus, there are four branches in this Stephenson linkage" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002128_s13369-020-05100-6-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002128_s13369-020-05100-6-Figure2-1.png", + "caption": "Fig. 2 Show the compressive force applied in different positions", + "texts": [ + " 1 Sketches and photographs illustrate a the dimensions of the CADWD machine, b isometric view of the designer model, c isometric view of the designer model in ANSYS workbench program and d experimental CADWDM system In the designing step of the proposed CADWDM system, most of the problems encountered in the design process were taken in mind, i.e., the stress concentration in the manufactured parts and their impact on the accuracy of the machine. Therefore, a specific load was taken, which is equivalent to the weight of the deposition torches and put in different areas in the machine to identify the stresses produced as well as the resulting deformation. The load applied in the fixture is 20 N and in gantry is 100 N as shown in Fig. 2. The load has been applied in a different positions to simulate the system to find the system\u2019s ability to withstand the stress concentration areas and strain, in addition to the deformation in the system. The machine safety factor is 15. While the yield stress of the machine is 280 MPa, the allowable stress of the machine can be obtained by dividing the yield strength by the machine safety factor (280Mpa/15 18.67Mpa). Therefore, the device is in safe side design because the maximum stress obtained is 11" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002191_j.matpr.2020.12.112-Figure10-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002191_j.matpr.2020.12.112-Figure10-1.png", + "caption": "Fig. 10. Total deformation-3.", + "texts": [ + " This point bargains about the compelled part appearing of the posts and imitating them by contributing specific physical representations and limit situations to reenact the fascinating situation plans ended likely. The segment stayed demonstrated utilizing ANSYS 14.0 software Plan Modeler programming as a strong perfect and ANSYS 14.0 worktable was utilized aimed at the assess- ment in the static examination. The key work method of a Limited Component Examination system is tended to like in Fig. 1(A) Fig. 2 Fig. 3 Fig. 4 Fig. 5 Fig. 6 Fig. 7 Fig. 8 Fig. 9 Fig. 10. The stainless harden fabric property Bulk (q): 7950 kg/m3 Young\u2019s modulus (E):206000Mpa Poisson\u2019s ratio: 0.3 The CAD typical of the bar is fit hooked on a limited quantity of components utilizing ANSYS 14.0 software inherent lattice calculation. The pillar contact locale is fit and interlinked to empower estimation of the power collaboration between them limited component investigation or FEA representing to a genuine task as a \u2018\u2018work\u201d a progression of little, consistently formed tetrahedron associated components, as appeared in the above fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002646_s11517-021-02347-5-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002646_s11517-021-02347-5-Figure8-1.png", + "caption": "Fig. 8 Prostate intervention robot integrated with BVS in MR environment", + "texts": [ + " The camera coordinates of the needle tip were reconstructed automatically from the images acquired by the BVS. The BVS is totally custom-made, and the algorithm specifically for the needle tip camera coordinate estimation has the advantages of general, simple, and fast computation. To the best of our knowledge, the algorithm, Eqs. (7)\u2013(10), is a novel method in the needle tip camera coordinate estimation. In addition, the custom-made BVS can be modified into MR conditional easily for use in the practical environment. As Fig. 8 shows, the MR conditional cameras are mounted on the holder which is connected to the base of the robot before the intervention process. The calibration, image process, stereo matching, and camera coordinate estimation of the BVS are conducted outside the MR bore, and a series of BVS parameters are obtained and integrated into the robot control software. Then the position of the cameras and the base of the robot remain constant by the holder, and the robot is secured on the MRI scanner bed. The needle tip position is measured by theMR conditional BVS in the needle placement phase, and detected by MRI during needle penetration" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003132_978-981-16-0119-4_20-Figure19.8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003132_978-981-16-0119-4_20-Figure19.8-1.png", + "caption": "Fig. 19.8 Verification of interpreted design", + "texts": [], + "surrounding_texts": [ + "A topology optimization result provides the guidance for developing a frame. A frame representing \u2018material, geometry, and manufacturing\u2019 feasibility needs to be created (see Fig. 19.7). The material, for example, is a standard material normally used on bicycles (\u2018IS 3074:2005\u20196). Certain additional members are added to connect members laterally. These are positioned in a manner that they do not interfere with other aggregates in the product.\nFor virtual verification, see Figs. 19.8, 19.9, and 19.10. The FE workbench within \u2018FreeCAD\u2019 is used. For meshing \u2018Netgen\u2019 and for solving \u2018CalculiX\u2019 are used (embedded inside the workbench). Two aspects Viz. \u2018stiffness\u2019 and \u2018propensity to yield\u2019 are assessed. The results are reasonable for the given conditions.\nThe \u2018VonMises\u2019 stresses provide a \u2018better correlation with experimental behavior than \u2018Tresca\u2019 yield criterion\u20197 and has been used to assess the developed stresses.\n6Bureau of Indian Standards. See: https://bis.gov.in. 7Teaching and learning packages of the \u2018University of Cambridge\u2019. See: https://www.doitpoms.ac. uk/tlplib/metal-forming-1/yield_criteria.php.", + "Amanufacturing team needs a drawing and specifications communicating the design (to the extent it impacts\u2014functional performance) (see Figs. 19.11 and 19.12).\nThe framework described above leverages technology to drive efficient design.", + "238 S. K. Mukherjee" + ] + }, + { + "image_filename": "designv11_83_0001848_robot.2007.363783-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001848_robot.2007.363783-Figure6-1.png", + "caption": "Fig. 6: Schematic of the manipulator correspond to the second category of the singular point", + "texts": [ + " In the forward kinematics solution, the direction of the normal vector of the moving platform must be found by the cross product of left B Z )( 5 and right B Z )( 5 that is: =m left B Z )( 5 \u00d7 right B Z )( 5 (37) If left B Z )( 5 and right B Z )( 5 are parallel then their cross product will be vanish. So Eq. (9) is always true and the only constraint equation for the two unknowns Rf , is Eq. (10). Since the numbers of the unknowns are more than the number of equations there exist many solutions for Rf , . In this case if the actuated joints locked that are direction of left and right legs are fixed, the platform could take different configuration. A schematic of the second category of singular point for the introduced mechanism is shown in Fig. 6. In this article forward and inverse pose of the novel Spherically Actuated Manipulator introduced in [1], were presented. In the inverse pose solution, active joint variables were calculated with no need for the evaluation of the passive joint variables. It was shown that there are 16 possible answers for the inverse pose problem. A closed form solution for the forward pose of the mechanism was obtained. Also it was verified that the maximum number of the solutions for the forward problem are four" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002342_j.colsurfa.2020.126012-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002342_j.colsurfa.2020.126012-Figure1-1.png", + "caption": "Fig. 1. The sample used for the forced wetting experiments. The lengths are not to scale. The dotted frame shows the observed corner of the sample.", + "texts": [ + " The experiments were performed in the same experimental setup used and described in [14,15]; hence, only a short description will be given here. The setup consists of an electro-pneumatic linear drive, dipping samples vertically into a liquid pool. This process is recorded by an imaging system in order to observe the wetting behavior and in particular to capture the instantaneous rivulet height. Two different sets of experiments with finite rivulet rise heights were performed in this study. The first set used a stainless steel sample, as pictured in Fig. 1, and silicone oil as a working fluid. This experiment was designed to examine the forced wetting case, when the sample is dipped into the bulk liquid. A second set of experiments was performed in which high contact angles were desired, yielding a finite rivulet height, even for the static situation, i.e. the sample does not move vertically into the bulk liquid. For this experiment the combination of a sample made of plastic and a supersaturated tenside solution was used to achieve the desired contact angle. These two experiments are described in more detail below. The first sample, used for the forced wetting measurements, is milled from 1.4404 stainless steel and sketched in Fig. 1; it has already been used in [15]. The bulk liquid pool consists of 20 cSt silicone oil (polydimethylsiloxane, brand: ELBESIL SILIKONO\u0308L B), which has a capillary length l\u03c3(= \u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305 \u03c3/(\u03c1g) \u221a ) of approximately 1.48 mm, calculated using the surface tension \u03c3, density \u03c1 and acceleration of gravity g, and can be assumed to wet stainless steel perfectly, i.e. it exhibits a very low contact angle. During these experiments the stainless steel sample was immersed into the pool with different constant velocities ranging from 0", + " Surface Evolver is a finite element tool that is used to minimize the energy functional E = \u222e Al \u03c3dA + \u222e Al \u03c1 z2 2 g\u2192\u22c5 n\u2192dA (1) by conjugate gradient and gradient descent methods subjected to constraints. In Eq. (1) E is the energy of the system, \u03c1 is the density of the liquid and Al the area of its liquid gas interface. g\u2192 has only one component, i.e. the z-component. In the present case an L-shaped crosssectional geometry of the liquid surface at z = 0 is chosen, as it is depicted in Fig. 2, wetting the walls of a corner (see Fig. 1) with vertex 1 only being allowed to move up and down the corner. The width and length of the two legs of the L-shape geometry, i.e. edges A and F, are chosen to be equal and of length 50 mm. The liquid has a constant contact angle at the walls of the corner (edges A and F) and the three- F. Gerlach et al. Colloids and Surfaces A: Physicochemical and Engineering Aspects 616 (2021) 126012 phase contact line is allowed to move freely. However the outer two vertices 2 and 6 at the wall, 50 mm away from the corner edge, are set to be at the capillary height l\u03c3 during the entire iteration process" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002432_tasc.2021.3064529-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002432_tasc.2021.3064529-Figure6-1.png", + "caption": "Fig. 6. Double-layered Dewar.", + "texts": [ + " 5 shows, for segmented HTS-LFSM, the secondary pole pitch should also be around twice the primary pole pitch or larger, or there would be insufficient room to establish the magnetic circuit. Considering if the pole pitch of secondary is too large, the utilization of room will be influenced, the most suitable primary and secondary pole pitch ratio should be around 1/2. Based on (3) and (4), sets of feasible pole pitch ratios can be obtained. Considering the former analysis of magnetic circuits, two suitable pole pitch ratios for the HTS-LFSM are 5/12 and 7/12. Fig. 6 illustrated the double-layered Dewar of HTS-LFSM. In this paper, every HTS coils is housed in an individual Dewar. Compared with the conditions where all the HTS coils and iron core are housed in one Dewar together, this system guarantees For the double-layered Dewar, the inner Dewar is fixed in the outer Dewar using poly tetra fluoroethylene. Between them exists a vacuum to provide heat insulation. The HTS coil is placed in the inner Dewar and liquid nitrogen is kept flowing in the inner Dewar, so that the heat from HTS coil and other sources can be brought away continuously" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001127_msec2007-31137-Figure11-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001127_msec2007-31137-Figure11-1.png", + "caption": "Figure 11. Strain prediction from AA6111 hemming simulations using the GMSS hardening law", + "texts": [ + " The excellent agreement between experimental and simulation results for both the true stress-strain determination and the comparison of loaddeflection response proves that the numerical procedure used in GMSS is error free (rather than a validation of GMSS method). GMSS can be used to provide engineers with accurate hardening behavior that is useful in the simulation of high strain deformation processes. For example, AA6111 hemming simulations showed sizable (i.e., 4%-13%) differences of strain outputs due to two different hardening law inputs, i.e., GMSS predicted hardening law vs. the power law. As shown in Figure 11, at exterior hemline (region A), the hemming simulation using the GMSS hardening law predicted the maximum principal (tensile) strain as 57%. Simulation was also carried out using the power hardening law (with K=463 and n=0.253 in nK\u03b5\u03c3 = ) model, with the maximum principal strain output as 53%. Similarly, at the interior hemline (region B), hemming simulation using GMSS hardening law predicted the minimum (compressive) strain as -90%, versus the -77% strain prediction using the power law. 6 Copyright \u00a9 2007 by ASME Copyright \u00a9 2007 by Gene al Motors Corporation ms of Use: http://www" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002020_acemp.2007.4510512-Figure23-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002020_acemp.2007.4510512-Figure23-1.png", + "caption": "Fig. 23 Distribution of magnetic flux density.", + "texts": [], + "surrounding_texts": [ + "3 Magnetic Characteristic Analysis of Three-Phase Transformer Figure 13 shows the analysis model of three-phase 5- legs construct ure transformer. The used material is the grain-oriented electrical steel sheet. Figure 14 shows the distribution of magnetic flux density. Furthermore, the inclination angle B from rolling direction of the vector B is shown in Fig. 15 in order to evaluate the structure of t he three-phase 5 legs typed transformer.", + "Figure 16 shows the distribution of distortion factor of magnetic flux waveform. The eddy current generates by distorting the magnetic flux waveform. The evaluation of waveform distortion gives large effect for the dynamic analysis. Furthermore, the magnetic power loss under the rotational flux is larger than under alternating flux. Figure 17 and 18 show the distribution of magnetic power loss by the dynamic E&S model and by the static E&S model, respectively. By using E&S model, magnetic power loss can be analysed directly. The comparison with the result by the conventional method (static E&S model) was examined in order to investigate features of the result by the dynamic analysis with dynamic E&S model, as shown in Fig. 19. The static E&S model has greatly evaluated the loss from actual result, since the effect of the distorted waveform is disregarded. It is found that magnetic field is suppressed by the effect of the eddy current. The static analysis excessively analyses the magnetic field.\n4 Magnetic Characteristic Analysis of Permanent Magnet Motor Figure 20 shows the concentrated flux typed surface permanent magnet motor (CSPM motor), which was developed for high density machine by our group and have been obtained by Oita University as the patent. The gap field of this motor has the 1.5 times of he magnetization of permanent magnet.\nFigure 21 shows the magnetic flux distribution. From this result it was found that the magnetic path becomes short, therefore the inductance increases as this motor.", + "Figs. 22 and 23 show the distribution of the magnetic field strength and the magnetic flux density, respectively. As same as procedure of the vector magnetic characteristic analysis, the distribution of magnetic power loss is obtained as shown in Fig. 24. The magnetic power loss increases in the place that both vectors B and H are large and the axis ratio in rotational magnetic flux is large. It is found that the major magnetic power loss is being generated at the tooth division of the stator core.\n5 Conclusions\nIn this paper we proposed the dynamic integrated typed E&S model for analyzing the eddy current effect under distorted magnetic flux conditions. And furthermore it was clarified how the effect of eddy current influence the vector magnetic characteristic.\nThe usefulness of dynamic vector magnetic characteristic analysis by using dynamic E&S model is shown in the following. (1) This method can analyse the effect of the eddy current which arises by the distorted flux waveform. (2) This method can consider the dynamic vector magnetic property and can analyse it. (3) The analysis of magnetic power loss (iron loss) can be directly carried out the magnetic characteristic analysis of the electrical machines.\nREFERENCES [1] H. Shimoji, M. Enokizono,\u201dE&S2 Model for Vector Magnetic Hysteresis Property\u201d, Journal of Magnetism and Magnetic Materials, 254, 290-292, 2003 [2] M. Enokizono, H. Shimoji, T. Horibe, \u201dLoss Evaluation of Induction Motor by using Hysteresis E&S2 Model\u201d, IEEE Transactions on Magnetics, Vol.38, No.5, pp.23792381,(2002) [3] S. Urata, H. Shimoji, T. Todaka, M. Enokizono \u201cMeasurement of two-dimensional vector magnetic properties on frequency dependence of electrical steel sheet\u201d, International Journal of Applied Electromagnetics and Machanics, 20, pp.155-162, 2004. [4] M. Enokizono, S. Urata, \u201cDynamic Vector MagnetoHysteretic E&S Model and Magnetic Characteristic Analysis\u201d, COMPUMAG 07, Aachen in Germany, 2007" + ] + }, + { + "image_filename": "designv11_83_0000153_6.2005-6156-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000153_6.2005-6156-Figure3-1.png", + "caption": "Figure 3, Definition of Indicator Surface in phase plan", + "texts": [ + " Now Su is proposed as follows: ( ) ++++\u22c5\u2212= u m u tM Tu L T S dduektef tg bPeu ),( )( 1)(sgn (16) Therefore \u2212\u2212\u2212\u2212\u2212++ + +\u22c5\u2264 u m u tM L T L u u m u t T M T ddu tg ek tg tefdd tg ektef ubPeW )()( ),( )( ),( American Institute of Aeronautics and Astronautics 6 0 )()()( ),( )( ),( \u2264 \u2212+\u2212\u22c5\u2264 tg ek tg ek tg tef tg tef bPeW L TT L u T According to Eq. (13), the last inequality implies 0 which is required to be asymptotically convergent toward an assigned goal frame < g >. By denoting with el the generalized error (position and orientation) of frame < l > with respect to < g >, let us define as x\u0307l := \u03b3el ; \u03b3 > 0 (48) the velocity reference signal that, once applied to < l >, would guarantee < l > itself to be asymptotically convergent to < g >" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002328_012035-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002328_012035-Figure1-1.png", + "caption": "Figure 1. Geometric configuration of origami patterned square tube.", + "texts": [ + " Series: Materials Science and Engineering 1062 (2021) 012035 IOP Publishing doi:10.1088/1757-899X/1062/1/012035 \ud835\udc36\ud835\udc39\ud835\udc38 = \ud835\udc40\ud835\udc52\ud835\udc4e\ud835\udc5b \ud835\udc39\ud835\udc5c\ud835\udc5f\ud835\udc50\ud835\udc52 \ud835\udc43\ud835\udc52\ud835\udc4e\ud835\udc58 \ud835\udc39\ud835\udc5c\ud835\udc5f\ud835\udc50\ud835\udc52 (1) Specific Energy Absorption (SEA) is the ratio of energy absorbed divided by the total mass of the structure. It allows a direct comparison of similarly shaped structures made from different materials. \ud835\udc46\ud835\udc38\ud835\udc34 = \ud835\udc38\ud835\udc5b\ud835\udc52\ud835\udc5f\ud835\udc54\ud835\udc66 \ud835\udc34\ud835\udc4f\ud835\udc60\ud835\udc5c\ud835\udc5f\ud835\udc5d\ud835\udc61\ud835\udc56\ud835\udc5c\ud835\udc5b \ud835\udc40\ud835\udc4e\ud835\udc60\ud835\udc60 (2) Based on literature review, an origami patterned square tube which has yet to be studied is proposed. The tube geometric configuration is illustrated in Figure 1. The function of the origami pattern is to allow the tube to fail in a predetermined and stable manner. The 3D CAD model of the origami patterned tube was created in Abaqus CAE preprocessor. The tube was modelled as surfaces where it was partitioned into several regions to represent the origami pattern. Also partitioning the surfaces enabled more uniform elements to be created during the subsequent meshing process. Table 1 shows the dimensions of the plain tube and origami patterned tube. Table 2 shows the detail parameters for the different origami pattern arrangements on the square tube" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002641_j.robot.2021.103783-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002641_j.robot.2021.103783-Figure1-1.png", + "caption": "Fig. 1. Diagram of a general redundantly actuated PM.", + "texts": [ + " Following the Introduction, displacement coordination equaions of the redundantly actuated PMs will be established in ection 2. In Section 3, the principle of force distribution in three ifferent control methodologies will be revealed. In Section 4, the est platform based on the 2RPR + P PM will be constructed, and he experimental test and analysis will be performed in Section 4. n Section 5 the application of three control methodologies are iscussed, and finally conclusions are drawn. . Establishment of displacement coordination equations Fig. 1 illustrates that a kind of redundantly actuated PMs ave n DOFs, consisting of m actuated limbs and one (or no) constrained limb, where m > n. Therefore, they belong to redundantly actuated PMs. It is assumed that the stiffness of the moving platform is higher than that of the limbs. Then, the elastic deformation of the moving platform is negligible. When only elastic deformations of limbs of the active overconstrained PMs are considered, the coordination correlation between the redundant and non-redundant limbs has two parts, including the displacement of the actuator di(i = 1, 2, " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000130_2005-01-2911-Figure17-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000130_2005-01-2911-Figure17-1.png", + "caption": "Figure 17: Trigon modular robotic construction system", + "texts": [ + " The Cubolding units use electric screw mechanisms for the linear prismatic actuator triangles that are particularly susceptible to wear and tear from dust and particles. In planetary surface constructions, the scaffolding units would need to be carefully designed to protect these mechanisms from wind-borne dust. The term \u201cTrigon\u201d comes from \u201ctriangle + polygon\u201d. The self-constructing Trigons are similar to the robotic Cubolding units in that they have the capacity to climb previously assembled portions of the structure. The system consists of triangular or square panels that when assembled can create trusses and other structural elements (Figure 17). Computer controlled revolute actuators at the panel edges affect precision motion relative to the panel body, allowing it to swing end over end along the completed structure. As the free end is swung around and approaches the structure surface, the connectors are moved inward. Once the connectors enter the clasping area of the existing structure, an outward pressure is applied and a firm connection is established. At this point the previous fixed end can now release, and the end over end motion is repeated until the panel finds its own place in the uncompleted structure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001522_cce.2008.4578933-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001522_cce.2008.4578933-Figure4-1.png", + "caption": "Figure 4. LOS technique", + "texts": [], + "surrounding_texts": [ + "Coordinates of way-points (xk,yk) for a desired trajectory in the earth-fixed reference frame can be selected from the system database and are based on the manoeuvring patterns, such as IMO search patterns for maritime search and rescue mission at sea. The database consists of \u2022 Desired way-point positions: wpt.pos: {(x0,y0), (x1,y1), ..., (xk,yk)} (1) \u2022 Desired speeds between way-points: wpt.speed: Ud = {u0, u1, u2, ..., uk} (2) \u2022 Desired heading angles: wpt.heading: d\u03c8 = { d1\u03c8 , d2\u03c8 , d3\u03c8 , ..., dk\u03c8 } (3) The desired heading angle d\u03c8 is calculated by the LOS technique as follows: k+1 dk k 1 y yatan2 x x+ \u2212\u03c8 = \u2212 (4) When the ship is moving along the desired trajectory, a switching mechanism for selecting the next way-point is necessary. The next way-point (xk+1,yk+1) is selected when the ship lies within a circle of acceptance with a radius R0 around the current waypoint (xk,yk) satisfying ( ) ( )2 2 2 k k 0x x y y R\u2212 + \u2212 \u2264 (5) The value of R0 is often chosen as two ship lengths, i.e. R0 = 2Lpp in [2][9]. A reference trajectory generator using a vessel simulator is constructed. The vessel model used in this paper is of Nomoto\u2019s first-order model with forward speed dynamics and described as follows: ( )d d dx U cos= \u03c8 (6) ( )d d dy U sin= \u03c8 (7) where (xd,yd) is the desired position, Ud > 0 is the desired speed and \u03c8d is the desired heading. The forward speed dynamics is ( ) 2 x d w d d xm m U 0.5 C AU\u2212 + \u03c1 = \u03c4 (8) where \u03c1w is the density of sea water, Cd is the drag coefficient, A is the projected cross-sectional area of the submerged hull of ship in the x-direction, and (m \u2013 mx) is the mass included hydrodynamic added mass. The course dynamics is chosen as d dr\u03c8 = (9) d d rTr r K+ = \u03b4 (10) where T and K are ship manoeuvrability indices, rd is the desired yaw rate and \u03b4r is rudder angle. The guidance system has two inputs, thrust \u03c4x and rudder angle \u03b4r. The guidance controllers can be chosen as PI and/or PID types, see [2][9]. When the ship goes along the desired trajectory, the reference heading angle can be adjusted by the exponential decay technique as shown in Figure 5. See detailed information in [9][15][9]. Heading and position errors when the ship is moving along the desired trajectory are calculated as follows 1 d d 2 d d 3 d e (x x)cos (y y) sin e (x x) sin (y y)cos e \u2212 \u03c8 + \u2212 \u03c8 = = \u2212 \u2212 \u03c8 + \u2212 \u03c8 \u03c8 \u2212 \u03c8 e (11) where e1 = path tangential tracking error e2 = cross-track error (normal to path) e3 = heading error If the rudder-roll damping controller is switched on the vector of errors including roll error (e4) becomes 1 d d 2 d d 3 d 4 e (x x)cos (y y) sin e (x x) sin (y y) cos e e 0 \u2212 \u03c8 + \u2212 \u03c8 \u2212 \u2212 \u03c8 + \u2212 \u03c8 = = \u03c8 \u2212 \u03c8 \u2212 \u03c6 e (12) If the speed controller is on the speed error will be calculated 5 de U U= \u2212 (13)" + ] + }, + { + "image_filename": "designv11_83_0000146_bf02329037-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000146_bf02329037-Figure3-1.png", + "caption": "Fig. 3--Graphical construction to obtain the response of the system", + "texts": [ + " Subst i tut ing eq (7) in the above equation, we obtain: - - ~ = F s - - m~o2 C I C F D B B where P _-- Ps ~--- For BD -- A 2 ) n '~ +n4~ Xa+P=Ps (II) BC -- AD m/pLa , mass rat io FL2/EI , dimensionless ampl i tude of exci t ing force FsLZ/EI, dimensionless spring force at peak value of deflection ]Yl ~ yo , P8 = a x X . lY[ < yo, P~ = ,~eXa + ( ,~ - ,~z)u where al ~ k l L Z / E I dimensionless constant of the bil inear spring ,~2 ~ k2L3/EI dimensionless constant of the bil inear spring Yo ~ y o / L The va lue of Xa de termined by eq (11) ensures the approximate val id i ty of eq (6). The solution of eq Experimental Mechanics [ 9.65 (11) c a n b e o b t a i n e d as i l l u s t r a t e d in Fig. 3. T h e r i g h t - h a n d s ide of eq (11) e x p r e s s e s t h e s p r i n g fo rce a n d t h e l e f t - h a n d s ide r e p r e s e n t s a s t r a i g h t l i ne w i t h o r d i n a t e i n t e r c e p t P a n d s lope S e q u a l to t a n -1 ( B D - - A ~ ) ~a -t- ~4~ . T h e i n t e r s e c t i o n of t h i s l i n e BC -- AD w i t h t h e s p r i n g c h a r a c t e r i s t i c g ives t h e v a l u e s of Xa w h i c h sa t is f ies e q (11) . F r o m Fig. 3, i t c a n b e o b - s e r v e d t h a t i f t h e s lope S is s u c h t h a t : Sa ( s lope of l i ne AD) > S > $2 ( s lope of l i ne AC) t h e r e w i l l b e t h r e e v a l u e s of Xa w h i c h can s a t i s fy eq (11) , s lope S m a y fa l l i n t h e r a n g e ($2, Sa) fo r a n in f in i t e n u m b e r of f r e q u e n c y r anges . T h e s e f r e q u e n c y r a n g e s a r e b o u n d b y t h e n a t u r a l f r e q u e n c i e s of a c a n t i l e v e r b e a m w i t h a l i n e a r s p r i n g s u p p o r t a n d a c o n c e n t r a t e d m a s s a t t h e f r e e end", + " P $1 = al - - - - (12a) Yo P $ 2 = a l -t- y - -~ (12b) Sa = a2 (12c) T h e a m p l i t u d e of e n d de f l ec t ion of t h e b e a m for d i f f e r e n t r a n g e s of s lope S can b e d e t e r m i n e d as fo l lows : (1) S --~ $1 : o n e s o l u t i o n P Xa -- (13a) (~1 - - S ) (2) $1 < S < $2 : o n e so lu t i on P + (~2 - - ~ I ) Y o Xa = (13b) (~2 - S ) (3) $1 ~ S ~ $2 : t h r e e so lu t i ons P Z a - - (,~1 - S) P + ('~2 - - a l ) Y o X . = (13c) ( ~ s - - S) P - - (,~2 - - ~ 1 ) Y o X a (~2 - - S) E q u a t i o n s (13) , (10) a n d (9) d e t e r m i n e t h e a p - p r o x i m a t e h a r m o n i c r e s p o n s e of t h e sys t em. I t c a n be n o t i c e d f r o m Fig. 3 t h a t o n l y one so lu t i on wi l l ex i s t for a l l v a l u e s of S if t h e fo rce P is m o r e t h ~ n i n - d i c a t e d b y p o i n t A'. P o i n t A ' is o b t a i n e d b y t h e i n t e r - sec t ion of t h e l i ne GB' w i t h t h e fo rce axis . F o r o b t a i n i n g t h e b a c k b o n e c u r v e s c o r r e s p o n d i n g to f r e e v i b r a t i o n s , P = o a n d in Fig. 3 p o i n t A wi l l co inc ide w i h p o i n t O. T h i s i n d i c a t e s t h a t f r ee v i b r a - t i ons w i l l ex i s t o n l y in t h e f r e q u e n c y r a n g e s for w h i c h al --~ S ~ ~s. F o r S = ~1, t h e a m p l i t u d e of e n d de f l ec t ion w i l l b e 0 to Yo, a n d S = as w i l l c o r r e s p o n d to Xa e q u a l to inf in i ty . F o r al < S < as, t h e a m p l i - t u d e w i l l b e b e t w e e n Yo a n d inf in i ty . T h e m o d a l c o n - f i g u r a t i o n c o r r e s p o n d i n g to a n y f r e q u e n c y of f r ee v i b r a t i o n is g i v e n b y eq (9 ) " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002731_s40799-021-00471-3-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002731_s40799-021-00471-3-Figure4-1.png", + "caption": "Fig. 4 Displacement in relation to the initial position", + "texts": [ + " One surface of the cam will be loaded when the force rises from 0 to +F, and the other surface will be loaded on return from +F to 0. Obviously, the highest stresses are in the surface of contact with the radial bearing and are 311 MPa. It is important that the maximum stresses in this area do not reach more than (0.65 \u2219 \u03c3y), where \u03c3y is the yield stress of the material from which the cam is made, given that the area subjected to maximum stresses will work at stress by compression fatigue after a pulsating cycle. Fig. 4 shows the displacement map where it can be seen that in the area of action of the radial bearings, the displacement is 0.1 mm under the action of the imposed pressing force of 18 kN. Due to the way in which the cam is constructed, contact between the radial bearing and one of the required cam guide surfaces will not be lost. Also illustrated in Fig. 4 is the position of the sides of the cam after moving under the action of the pressure force. In this figure, the displacement is increased in scale in relation to the real situation in which the maximum displacement is 0.1 mm. The fatigue testing takes place in the elastic domain until accentuated degradation of the material begins. When the loading occurs, the bearings on the cam vertically move on distance v (also distance AB in Fig. 5) and rotate, with respect to the central axis, at an angle of\u03c6 (arc BC in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001848_robot.2007.363783-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001848_robot.2007.363783-Figure4-1.png", + "caption": "Fig. 4: The direction of h, n and m", + "texts": [ + " (14), (since both represent leftX )( 6 3 ) the desired ff 21 ,\u03c6\u03c6 could be determined as below: = = 3 2 1 n n n n knownCompletely Bleft B left RXX 3 66 3 )()( = = \u2212 ff f ff CC S CS 21 2 21 \u03c6\u03c6 \u03c6 \u03c6\u03c6 (16) So: =f1\u03c6 atan2 ( ), 31 nn \u2212 (17) and f2\u03c6 is: =f2\u03c6 atan2 ( )/, 132 fcnn \u03c6\u2212 if \u03c0\u03c6 orf 01 = (18) =f2\u03c6 atan2 ( )/, 112 fSnn \u03c6 others (19) Now TB P can be computed as below: TB P = TTTTTT P B 65 6 4 5 3 4 0 30 (20) )(: ),(:, )(: )(: )(: 6 21 5 6 4 5 3 4 0 3 0 knownisionconfiguratthebecauseknownT dcalcualetearebacuaseknownTT calcualtedisbecauseknownT knownareinputsbecauseknownT knownisionconfiguratthebecauseKnownT P ff f B \u03c6\u03c6 Finally, since Eq. (17) gives two possible answers for f1\u03c6 , for each pair of ( f , R ) there are two possible answers for the forward pose solution. Moreover, Eq. (11) can lead to two answers for f thus two pairs of ( f , R ) can be acceptable and as a result there exit four possible solutions for the forward pose problem. VI. INVERSE RATE KINEMATICS In inverse rate Kinematics, having the linear velocity of point G shown in Fig. 4 as well as the angular velocity of the moving platform, all \u03b8 have to be evaluated. Consider the following definitions: GARqzlSzz /3 ,\u02c6,\u02c6\u02c6 \u2032=== where, 3z\u0302 is shown in Fig. 2, l is the length of the link and GAR /\u2032 is the position vector of point A\u2032w.r.t point G . The velocity of point A\u2032 in Fig. 4 is: qtS P \u00d7+= \u03c9 (21) where t and P\u03c9 are the linear velocity of point G and angular velocity of the moving platform, respectively and are known. Also S from the leg view is: SslS L Z \u00d7++= \u03c9 \u03c9\u03c9 )(\u02c6 (22a) where L\u03c9 is the angular velocity of the leg, Z\u03c9 is the component of L\u03c9 along 3z and\u03c9 is composed of the components of L\u03c9 which are in the plane that has 3z as the normal vector. Since Z\u03c9 and S have the same direction there cross product will be vanished. Therefore Eq. (22a) is: SslS \u00d7+= \u03c9\u02c6 (22b) Taking the dot and cross product of Eqs", + " Equaling the angular velocity of the cross symbol of the universal joint, from the platform and leg\u2019s views, we have: hznp l \u03c9 \u03c9 \u03c9\u03c9\u03c9\u03c9 ++=+ (24) In which n\u03c9 is the relative angular velocity of the cross symbol of the universal joint w.r.t the platform and is in the direction of n ; h\u03c9 is the relative angular velocity of the cross symbol of the universal joint w.r.t the leg which is in the direction of h and at last Z\u03c9 is the component of the leg's angular velocity in the leg\u2019s direction. Defining k as: hnk \u00d7= (25) where n and h are shown in Fig. 4 and taking the dot product of the Eq. (24) with k leads to: zp kkk \u03c9\u03c9\u03c9 \u22c5+\u22c5=\u22c5 (26) Therefore Z\u03c9 which is the component of the leg\u2019s angular velocity along itself is equal to: s sk k p z \u02c6 \u02c6 )( \u22c5 \u2212\u22c5 = \u03c9\u03c9 \u03c9 (27) Thus the angular velocity of the leg is ( \u03c9\u03c9\u03c9 += zl ): l Sss zkl Ssk zk k p l \u00d7+ \u22c5 \u00d7\u22c5+ \u22c5 \u22c5 = \u02c6\u02c6 \u02c6 1\u02c6 \u02c6 \u03c9 \u03c9 (28) Eq. (28) in the matrix format is: [ ] [ ] [ ]Tpl tSpvFpv \u03c9\u03c9 = (29) where ( ) s lzh hz l sFpv T ~\u0302 \u02c6\u02c6 \u02c6\u02c6~\u0302 \u22c5 \u2212= (30) ( ) ( )qs lzh hz l qs zh hzFsv TT ~~\u0302 \u02c6\u02c6 \u02c6\u02c6~~\u0302 \u02c6\u02c6 \u02c6\u02c6 \u22c5 +\u2212 \u22c5 = (31) RULE I The rule below was used in deriving Eqs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000239_iecon.2006.347435-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000239_iecon.2006.347435-Figure3-1.png", + "caption": "Fig. 3. (a)Method of detect cross (b)Detected cross", + "texts": [ + " The degree of circularity takes the value of 0 \u2264 e \u2264 1, and a perfect circle has 1. Then we set the suitable value of the threshold and detect the connection region with the bigger degree of circularity than threshold. This paper set threshold 0.6 in order to allow some distortion of the circle of the mark by rotation of the mark. Next we detect the cross mark on the search region which is the circle region as mentioned. The search line which is the square region with 1/3 of the diameter of a circle is prepared, and in Fig. 3, the search line is dashed line. Detection of the cross mark is conducted by using the cross candidate points which are set of a black pixel on the search line. This paper assumes that (i) there are four cross candidate points on the search line, (ii) there are two pairs if the angle which the cross candidate point and the search center make, and (iii)there are four cross candidate points that the black pixel continues to the search center. It assumes that the region which fulfills these assumptions is the cross and if the cross is detected, we get center of the search region (uk, vk) as the mark position on the image" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001400_6.2008-7171-Figure10-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001400_6.2008-7171-Figure10-1.png", + "caption": "Figure 10 Instantaneous attitude error", + "texts": [ + " (13) The high frequency switching is filtered out using the equivalent control 1 ~ ~\u0302 + = s v v p p p \u03c4 , (14) where p\u03c4 is the filter time constant to be tuned. Finally, the resulting designs of the control terms are pppp pp Kvu u \u03c3 \u03c8 ~~\u0302~ ~~ 1 0 += = , (15) where assures exponential convergence of the sliding variable to zero. 0>pK B. Attitude Controller Design Applying inverse rotation kinematics, the desired instantaneous axis and angle of rotation are defined by computing the cross product between current thrust force vector F and the desired acceleration vector \u03b3~ normalized to unit length (Figure 10). \u03b3 \u03b3 ~ ~ ~ ~ ~ \u00d7= F Fud . (16) The direction of unit vector gives the desired axis of rotation. Therefore, the actual attitude error angle isdu \u239f\u239f \u239f \u23a0 \u239e \u239c\u239c \u239c \u239d \u239b \u00d7= \u2212 \u03b3 \u03b3\u03b5\u03b8 ~ ~ ~ ~ cos~ 1 F F . (17) Equations (18) express the attitude, angular velocity and angular acceleration errors, respectively \u03c9\u03c9\u03c9\u03b5\u03b5 \u03c9\u03c9\u03c9\u03b5\u03b5 \u03b3 \u03b3\u03b8\u03b8\u03b5 \u03b8 \u03b8 \u03b8 ~~~~~ ~~~~~ ~ ~ ~ ~ ~ ~ cos~~~ 1 \u2212=\u2212== \u2212=\u2212== \u22c5 \u239f\u239f \u239f \u23a0 \u239e \u239c\u239c \u239c \u239d \u239b \u00d7=\u2212= \u2212 cw cw d u d c u u F F . (18) Following similar sliding mode observer techniques, a sliding variable of second order is designed \u222b++= dtkkk ipv \u03b8\u03b8\u03b8\u03b8 \u03b5\u03b5\u03b5\u03c3 ~~~~ 222 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001658_6.2007-2229-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001658_6.2007-2229-Figure1-1.png", + "caption": "Figure 1: Composite Materials", + "texts": [ + " Compared with aluminum, the composite materials used in the advanced program will offer tremendous advantages to aircraft designers. These lighter weight materials use less fuel, resist fatigue, and prevent corrosion. Composite materials are not new to aircraft structures; since the mid-1960s, military aircraft have driven the development of advanced materials, but today an increasing share of new commercial airplane development is devoted to composites. For example, composites will make up approximately 50% of the weight of the Boeing Advanced Dreamliner, a sharp increase from the 12% figure for the Boeing 777. See Figure 1. II. The Partnership: Academic\u2013Industry-Community Addressing the competency shortfalls identified by Boeing Commercial Aircraft Structures Engineering group, the Learning, Training, and Development group approached the University of Washington, through the departments of Aeronautics and Astronautics and Mechanical Engineering to assist in the development of a composite curriculum that could be delivered within higher education to prospective engineers and technology practitioners. Central to this industry\u2013academic partnership was the development of new coursework to enhance student learning both at the higher education and at industrial in-service levels" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002462_s40430-021-02915-8-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002462_s40430-021-02915-8-Figure7-1.png", + "caption": "Fig. 7 Wavelet response of the velocity induced by the nematodes considering different stroke periods. The images are represent the least (top) to the most efficient (bottom)", + "texts": [ + "8T the asymmetric upstroke (UPS) and downstroke (DWS), swimming velocity profiles appear substantially different. In particular, the waveforms appear trimodal instead of bimodal, and this has been observed in different modes of interpolation. We observed that the beginning of the motion, when the filament is still fighting inertia, the symmetric motion provides better propulsion. This behavior is observed until the third swimming stroke. However, as inertia is overcome, the presence of higher frequencies and vortices induces best swimming performance. Figure\u00a07 shows the wavelet transform and more specifically Fig.\u00a07d shows a greater amount Journal of the Brazilian Society of Mechanical Sciences and Engineering (2021) 43:207 1 3 207 Page 12 of 13 of energy in longer periods. Interestingly, the large amount of energy generated at the beginning so that the filament can overcome the inertia in these conditions, also improves its efficiency in swimming. It is important to emphasize that this would require a lot of energy from a biomechanical system and therefore it may not be the ideal model. On the other hand, Fig.\u00a07c presents a good alternative. It is also noted in Fig.\u00a08 that again, the higher coherence is present below 8\u00a0ms and in specific positions. In this case, the vigorous start is important to overcome the reversibility due to low Reynolds number, making the microorganisms able to swim greater distances by applying greater energy at the beginning. Motivated by the growing interest in active suspensions, we conducted a study regarding the propulsion of a flexible filament emulating C. elegans motion. We were particularly interested in understanding how the variation of parameters of the motion of an orientational flexible filament changes its propulsion" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001652_bf03177400-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001652_bf03177400-Figure5-1.png", + "caption": "Fig. 5. CRD nozzle welds and cracks [8].", + "texts": [], + "surrounding_texts": [ + "The manipulator performs auto-welding on the sloped wall of the CRD nozzle as shown in Fig. S. The welding trajectory can be represented by a set of points; however, the motion of the manipulator has some inevitable basic tracking error due to inter polation errors. It was determined from experimental trials that the tracking error in GTAW auto-welding should be less than 1.0 mm. A welding trajectory with 0.3 mm maximum interpolation error was generated by dividing the track into suitably small increments. To obtain uniform weld condition requires a con stant linear speed of the welding torch. In this study, the tip speed of torch was maintained constant by controlling the input speed of the manipulator." + ] + }, + { + "image_filename": "designv11_83_0003272_acc50511.2021.9483178-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003272_acc50511.2021.9483178-Figure2-1.png", + "caption": "Fig. 2. CAD representation of FIP under consideration, fasteners and encoder not shown", + "texts": [ + " The differential equations relating the pendulum angle and the voltage applied to the motor are developed and 978-1-6654-4197-1/$31.00 \u00a92021 AACC 1730 Authorized licensed use limited to: University of Glasgow. Downloaded on August 22,2021 at 19:25:57 UTC from IEEE Xplore. Restrictions apply. the resulting transfer function is presented. Derivations of a flywheel-inverted-pendulum mathematical model can be found in numerous other works [17] but since a substantial number are incorrect, the transfer function is derived afresh here. A depiction of the plant is given in Fig. 2. It comprises an electric motor, optical encoder (not shown), 3D-printed components,2 bearings, and fasteners (also not shown). The motor is a Maxon brushed 90 watt model. The armature of this motor provides sufficient rotational inertia so an additional flywheel is not needed. Neither the motor-shaft position nor its speed are measured. The angle of the pendulum rod is sensed with a Maxon five hundred counts-per-turn optical quadrature encoder. The two base parts and the pendulum rod are 3D printed. They are connected through a pair of 608ZZ ball bearings. The shaft protruding from the base part on the right in Fig. 2 is part of the printed pendulum rod and is what the encoder connects to. The displacement parameters for the FIP are shown in Fig. 3. The angular displacement of the pendulum rod from vertical is denoted with \u03b81. The displacement of the motor shaft from a position aligned with the pendulum rod is denoted with \u03b82. The distance between the pendulum axis and the motor shaft is given as l1 and the distance between the pendulum axis and the center of mass of the rod is lC1. Applying Newton\u2019s second law for rotation to the pendulum rod and combining it with the results of applying Newton\u2019s second law (for translation) to the motor armature yields (m1lC1 +m2l1) (gS1)\u2212b1\u03b8\u03071 +b2\u03b8\u03072\u2212\u03c4m = ( J1 +m2l 2 1 ) \u03b8\u03081 (1) where mi is the mass of the i-th link (in this case the pendulum rod and the motor), g is the gravitational acceleration 2CAD files are available at https://github" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000815_s1068371208060060-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000815_s1068371208060060-Figure5-1.png", + "caption": "Fig. 5. Constitutive vectors in the sector 0 \u2264 \u03b8 \u2264 . \u03c0 3 --", + "texts": [ + " V* 2 3 -- v a0 v b0e j 2\u03c0 3 ------ v c0e j 4\u03c0 3 ------ + +\u239d \u23a0 \u239b \u239e ,= VECTOR PWM IMPLEMENTATION IN A THREE-PHASE, THREE-LEVEL RECTIFIER 319 INITIAL STATEMENTS OF VECTOR PWM IMPLEMENTATION IN A THREE-LEVEL RECTIFIER Different states are considered for the rectifier elements that result in a total of 25 combinations, by means of which 19 space vectors of power means are implemented (Fig. 4). The vector PWM may be synthesized similarly to that shown in [4] for the two-level rectifier. In general, a constitutive vector may be formally represented as follows by a linear combination of the three nearest (forming) generalized vectors ( , , ): where \u03c4i, j, k is the weight coefficient (percentage) of the forming vectors. For example, to synthesize a constitutive vector for sector 1, 0 \u2264 \u03b8 \u2264 (Figure 5), a combination and three adjacent forming vectors may be engaged from the same triangle where constitutive vector is located. The absolute values of the forming vector are shown in Table 1. For triangle 2: (1) Expression (1) is the necessary basis on which to solve the problem, which also requires that weight coefficients \u03c41, \u03c42, and \u03c43 be determined. Different solution methods exist, the selection of which is usually governed by the hardware implementation of the control system. Here, we engage the method presented in [4], for which purpose it is expedient to perform a transformation to coordinates \u03b1, \u03b2 as follows: (2) (3) The equation (4) can serve as the third equation to the system of equations (2), (3)", + "+ += Similarly, the constitutive vector weight coefficient may be determined for all of the triangles. Table 2 presents the weight coefficient expressions for the sector 0 \u2264 \u03b8 \u2264 . BOUNDARY CONDITIONS FOR THE CONSTITUTIVE VECTOR To synthesize the constitutive vector, the number of forming vectors is primarily governed by the location \u03c0 3 -- 320 RUSSIAN ELECTRICAL ENGINEERING Vol. 79 No. 6 2008 BROVANOV, KHARITONOV of the constitutive vector on the plane of one of the triangles formed by the 25 system states. For example, consider the sector formed by triangles 1\u20132\u20133\u20134 (Fig. 5); constitutive vector may be located within any of these triangles. For the convenience of vector PWM implementation, divide the hexahedron formed by the current state combinations into six 60\u00b0 sectors. Then, the problem is reduced to the following stages of synthesis: (1) The number of the sector in which the constitutive vector is located must be determined; (2) The triangle where the constitutive vector is located at the very moment of the synthesis of the rectifier input voltages must be determined. Consider the sector that includes the triangles 1\u20132\u2013 3\u20134 (Fig. 5). The triangles shown are demarcated by straight lines l, m, and n. Triangle 2 is demarcated from the triangle 3 by the l line, whereas triangle 3 from triangles 1 and 4 are defined by lines m and n, respectively. Thus, the straight demarcation lines may provide certain conditions that help to determine the triangle where the constitutive vector will cross. For example, the condition for the end of the constitutive vector to transfer from triangle 2 to triangle 3 is for straight line l to cross", + " Because the location of the constitutive vector (characterized by a length and an angle) is the subject of our study, the boundary, V* i.e., straight line l, should be described in the polar coordinate system. If l line limits the constitutive vector length, then this condition may be written in the form where |p| is the vector absolute value, the vector extends from the O point up to the l line under the right angle, and \u03b1 is the angle between the OP polar axis and the p vector. The absolute value of | | is easily determined from Fig. 5 and is equal to ; thus, \u03b1 = . Hence, the boundary value of the constitutive vector length between the second and the third triangles is equal to Based on the expression obtained, the two following conditions may be formulated to determine the constitutive vector location within triangle 2: Similarly, the constitutive vector location within triangle 1 takes the form where The condition for the location of the constitutive vector within triangle 3 takes the form where The condition for the location of the constitutive vector within triangle 4 has the form V* p \u03b8 \u03b1\u2013( )cos ------------------------- ,= p p p 1 2 3 --------- \u03c0 6 -- Vb23* 1 2 3 \u03b8 \u03c0 6 --+\u239d \u23a0 \u239b \u239ecos ------------------------------------- 1 3 3 \u03b8cos \u03b8sin\u2013( ) ------------------------------------------------" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002276_s12206-021-0114-2-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002276_s12206-021-0114-2-Figure2-1.png", + "caption": "Fig. 2. Models and sections of a premium connection: (a) box part; (b) pin part.", + "texts": [], + "surrounding_texts": [ + "The analysis process is summarized in Secs. 2 and 3. The 2- D axisymmetric model was used to analyze the premium con- nection system owing to its shorter analysis time and simplistic nature in comparison to a 3-D model. For tensile and compressive forces, the analysis results of the 3-D model and 2-D axisymmetric model were in agreement. The 3-D modeling-based analysis was required to analyze situations where bending occurs [13]. Fig. 5 shows the geometry of an API standard premium connection. Fig. 6 describes the parameters applied to the box. The box length is 234.95 mm, with an outer and inner diameter of 153.67 mm and 143.26 mm, respectively. In Fig. 7, the box thread is magnified to illustrate the detailed geometric parameters of the thread. The thread thickness is 4.72 mm with a space length of 7.58 mm. The pitch is 15.25 mm and the thread length is 7.04 mm. Fig. 8 describes the parameters related to the pin. The outside diameter and thickness of the pin is 139.7 mm and 10.54 mm, respectively. Fig. 9 shows a magnified image of the pin thread and summarizes the thread geometry parameters. The space length, pitch, and thread length are 7.25 mm, 15.25 mm, and 7.97 mm, respectively. Fig. 10 describes the parameters related to the thread. The upper (stabbing flank) and lower corner radius (load flank) are both 0.2 mm." + ] + }, + { + "image_filename": "designv11_83_0001753_cca.2007.4389347-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001753_cca.2007.4389347-Figure3-1.png", + "caption": "Fig. 3. Gravity forces and propulsive force in the vertical plane", + "texts": [ + " The dynamic model as supplied by the manufacturer has been improved in this study and the DC motors are simulated with respect to the corresponding equations. The TRMS possesses two permanent magnet DC motors; one for the main and the other for the tail propelling. The motors are identical with different mechanical loads. The mathematical model of the main motor, as shown in Fig. 2, is presented in (1) to (5). The mathematical model of the remaining parts of the system in vertical plane is described in (6) to (8) (see Fig. 3). dt di LiREU av avavavavv ++= (1) vvavav kE \u03c9\u03d5= (2) vmr v mrLvev B dt d JTT \u03c9 \u03c9 ++= (3) avvavev ikT \u03d5= (4) vvtvLv kT \u03c9\u03c9= (5) In (6) the first term denotes the torque of the propulsive force due to the main rotor, the second term refers to the friction torque that covers all viscous, coulomb and static frictions, and the torque of gravity force is shown in the third term. [ ] v vvvfricvvmv J CBAgTFl dt d \u03b1\u03b1\u03c9 sincos)()( , \u2212\u2212+\u2212 = \u2126 (6) where, ttstr t lmm m A ) 2 ( ++= , mmsmr m lmm m B ) 2 ( ++= , ) 2 ( cbcbb b lml m C += \u00af \u00ae < \u2265 = 0 0 )( vvvfvn vvvfvp vv fork fork F \u03c9\u03c9\u03c9 \u03c9\u03c9\u03c9 \u03c9 (7) v v dt d \u2126= \u03b1 (8) It is noted that there are differences between the input voltage levels in the MATLAB/Simulink environment and the motor terminal voltages, and that the relationship between these two sets of values is nonlinear" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001187_imece2007-43017-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001187_imece2007-43017-Figure2-1.png", + "caption": "Fig. 2 Coordinate systems for the universal face-hobbing hypoid gear generator", + "texts": [ + " In the face-milling cutter head, the inner and outer lade groups are evenly arranged on two respective concentric ircles; therefore, in the above equation, only the profile angle nd the curvature and cutter radii need be taken into account. 2 Copyright \u00a9 2007 by ASME shx?url=/data/conferences/imece2007/71496/ on 02/27/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use 2. Mathematical Model of a Universal Face-Hobbing Hypoid Gear Generator A mathematical model of a universal face-hobbing hypoid gear generator for spiral bevel and hypoid gears was established in Ref. [9]. As shown as Fig. 2, this machine is a virtual cradletype machine having tilt, a cradle, work-gear support mechanisms, and so on, so that it can simulate existing facemilling and face-hobbing cutting systems with or without AFM motions. The coordinate systems ( , , ) t t t t S x y z and 1 1 1 1 ( , , )S x y z are rigidly connected to the cutter head and the work gear, respectively. The transformation matrices t S to 1 S yield the following surface locus for the cutting tool in coordinate system 1 S : ( ) ( ) ( ) ( ) 1 1 1 1( , , , )= ( ) ( ; ) ( ) ( ) , , , , , , , U U U U c f fa i c at t i c R m m u u i j S E A B \u03b2 \u03c6 \u03c6 \u03c6 \u03b6 \u03c6 \u03b2 \u03b6 \u03b8 \u03b3 \u22c5 \u22c5 \u22c5 = \u2206 \u2206 M M Mr r (2) where \u03b2 and 1 \u03c6 are the rotation angles of the cutter and work gear, respectively, and u is the variable of the blade edge" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001462_isam.2007.4288476-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001462_isam.2007.4288476-Figure2-1.png", + "caption": "Fig. 2. 2D notations for Cylinder/Plan interactions", + "texts": [ + " CYLINDER/PRISM ANALOGY AND DIFFERENCE This section aims at defining the meniscus shape equation, and calculating the volume of liquid, for both geometries interactions : prism and cylinder interacting with a plane. A. Meniscus shape Let us describe some notations used in Fig. 1, \u03b81 and \u03b82 are the contact angles between liquid and, respectively, the plane and the prism (or the cylinder), z is the distance between the object and the plane, \u03c6 represents the aperture angle for the prism and the immersion angle for the cylinder (fig. 2), h is the immersion height, \u03b1 is the sum of both angles \u03c6 and \u03b82, x1 and x2 are positions of the contact line with liquid. Both interaction models presented here below are based on a simplification of the Laplace equation giving the pressure difference across the liquid-vapor interface pin \u2212 pout as a function of the surface tension \u03b3 and the meniscus curvature H [5] : 2\u03b3H = pin \u2212 pout (1) which can be rewritten into : \u03b3( 1 R1 + 1 R2 ) = pin \u2212 pout (2) where ( 1 R1 + 1 R2 ) represent the double of the mean curvature H", + " In equation 34 (energetical method), the term factor of z z+h can be expressed as: ( \u03b2 \u2212 cos \u03b81 tan \u03c6 \u2212 cos \u03b82 sin \u03c6 ) 1 1 + \u00b5 tan \u03c6 = \u2212 cos \u03b81 + cos\u03b1 tan \u03c6 (36) Equation 34 can be rewritten into 1 2L\u03b3 dW dz = cos \u03b81 tan \u03c6 + cos \u03b82 sin \u03c6 \u2212 z z + h cos \u03b81 + cos\u03b1 tan \u03c6 (37) By substracting and adding sin\u03b1 to the latter equation, the expression of force can be found : 1 2L\u03b3 dW dz = h h + z ( cos \u03b81 + cos\u03b1 tan \u03c6 ) + sin\u03b1 For the prism/plan interaction, both methods are identical, it can also be shown numerically for the interaction between cylinder and plan. IV. REAL CASE: CYLINDER/PLAN INTERACTION A. Laplace approach Using expression 15 and cylinder parameters (see Fig. 2), a mathematical relation similar to 16 can be found between V and h. Unfortunately, h cannot be found analytically2 and a numerical algorithm is needed to perform \u03c6 calculus. The capillary force expression is deduced from relation 19 and is given by : F = 2L\u03b3 ( R cos \u03b81 + cos\u03b1 z + R(1 \u2212 cos \u03c6) sin \u03c6 + sin\u03b1 ) (38) 2that is the reason why this relation has been studied, because it has an analytical solution in the case of prism/plane interaction. B. Numerical Equivalence By applying the energetical method, an expression of W similar to eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000676_978-1-4020-8829-2_9-Figure9-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000676_978-1-4020-8829-2_9-Figure9-1.png", + "caption": "Fig. 9. Optimal designs for the isotropy maximization of the Delta robot", + "texts": [ + " The corresponding optimization results are reported in Table 1 and lead to an optimal isotropy index of 91.36% from the first starting point. With five optimization variables, the solution is obviously improved and the average isotropy over the cube vertices reaches 98.98%. The optimization results can be found in Table 2. The SQP algorithm needs 54 evaluations of the objective function. It should be noted that the isotropy index does not depend on the radii themselves but rather on their ratio which tends to unit value as shown in Table 2. Initial and optimal designs are sketched in Fig. 9. The Hunt platform (see Fig. 10) has three position and three orientation degrees of freedom which require to normalize Jf before computing \u03ba, each time the parameters change, i.e. at each call of the objective function. As explained above, this involves an additional optimization parameter: the characteristic length LC . The nine other parameters of this optimization problem (see Fig. 10) are the legs lengths LI and LS, the characteristic radii of the platform RP and of the base RB, the gauge H between adjoining actuators on the base, the angle \u03b1 (around a vertical axis), followed by angle \u03b2 (around an horizontal axis), angle \u03c8, and finally the vertical distance zc between the base and the center of the desired workspace volume (small cube in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000883_optim.2008.4602378-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000883_optim.2008.4602378-Figure3-1.png", + "caption": "Figure 3. The magnetic field spectra within the machine with respect to the (Od) axis, to the left, and (Oq)-axis, to the right respectively.", + "texts": [ + " Rated Angular Speed rpm 2840/420 4. Pair-Pole Number - 2/12 5. Starting-Capacitor \u03bcF 14 The self and mutual inductances of the machine were not available on data sheets. Therefore motor's inductances have been calculated with the FEM method based on the geometry of the machine [6], [7]. The software environment used for this purpose was Maxwell 2D. The spectra of the magnetic induction field within the machine with respect to the (Od) and (Oq) axes, computed with the FEM method, are presented onto Fig. 3. In Fig. 5 and Fig. 6 the distributions of the magnitude of B-vector onto the induced surface are also depicted.. As seen onto these plots, the distribution of the B magnitude onto the motor's air gap approaches a sinusoidal shape on both axes, however the magnitude on the two axes flux linkages differs. Based on the FEM analysis, the inductances of the machine have been calculated; the results are presented in Table III. TABLE III. THE EQUIVALENT SELF INDUCTANCE AND MUTUAL INDUCTANCE OF THE MACHINE MSP 311 CALCULATED WITH FEM METHOD No Denomination Symbo l Units Value 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002796_j.matpr.2021.04.285-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002796_j.matpr.2021.04.285-Figure1-1.png", + "caption": "Fig. 1. Existing Spur Gear Design.", + "texts": [ + " By using ISO and JGMA, different stresses were calculated and compared with obtained FEM results [9]. Yilmaz Can et al., They compared experimental findings to velocity field results and discovered a better upper bound solution. In addition, the fatigue activity of gear tooth was investigated [10]. The teeth of the Spur gears are usually parallel to axes, so when the two gears are engaged, there is no axial thrust due to tooth loads; contact extends the entire width on a line parallel to the rotation axes Fig. 1. At high speeds, the load is suddenly applied, resulting in a high impact load and unnecessary noise. Spur gears have a lower power transmission capability than helical gears because they have a smaller contact area. The friction on the spur gears is also rising. Noise and vibration are primarily caused by frequent variations in tooth stiffness. Spur gears with a high contact ratio may be used to eliminate or minimize tooth stiffness variation. When the axes of spur gears are not parallel to each other during assembly, static and dynamic stresses in the teeth of gears occur" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003132_978-981-16-0119-4_20-Figure19.1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003132_978-981-16-0119-4_20-Figure19.1-1.png", + "caption": "Fig. 19.1 Layout of a bicycle", + "texts": [ + " Bathe[17]: \u201cThis method is used for linear analysis of solids and structures that starts with idealization of the system, discretization of the system into a set of springs which follows the Hooke\u2019s law, formulation of the equilibrium equation and then finding the solution to the equation.\u201d From \u2018Lecture-1\u2019 of an online course available under the Creative Commons License BY-BC-SA. Link: https://youtu.be/oNqSzzycRhw. design of a bicycle (based on a design from the 1880s3) has chain-driven traction on the rear-wheel-sprocket, and both wheels are of a similar rolling radius. The layout of the bicycle (see Fig. 19.1) involves: starting with a particular size of wheels; relative positioning of the saddle, the pedal and handlebar; that are dependent on anthropometrics as well as desired handling characteristics. Patterson et al. [8] have reported that trail and the \u2018moment of inertia\u2019 are important factors to determine the handling qualities of a bicycle, when designed. Within the presented framework, the focus is to develop a design proposal for the frame; as other aggregates such the sprockets, pedals, and pedal cranks can be selected off-the-shelf" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001431_12.774989-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001431_12.774989-Figure8-1.png", + "caption": "Fig. 8 Analysis of obstacle climbing", + "texts": [ + " Modeling the resistance force due to obstacle climbing in soft soils Proc. of SPIE Vol. 6795 67954R-4 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 06/17/2016 Terms of Use: http://spiedigitallibrary.org/ss/TermsOfUse.aspx and compliant obstacles is an extremely involved process, here the obstacle is assumed to be discrete and incompliant, a manageable configuration equation of obstacle resistance can be derived from the equations of static equilibrium at the moment that the front wheel get up as shown in Fig. 8. The total resistance on front wheel due to obstacle climbing is: cos sinOF F FR N a N a\u00b5= \u2212 \uff085\uff09 From the equations of static equilibrium: 0 : cos sinY F R FF N N N W\u00b5 \u03b1 \u03b1\u03a3 = + + = \uff086\uff09 0 : sin cos 0X F F RF N N DP\u00b5 \u03b1 \u03b1\u03a3 = \u2212 + = \uff087\uff09 0 : / 2 0A F R RM N r DP r N L WL\u00b5\u03a3 = + \u2212 + = \uff088\uff09 The resistance ROF can be obtained as: (cos sin ) 2( cos sin cos sinOF WL a aR r r a r a uL a L a \u00b5 \u00b5 \u00b5 \u2212 = + \u2212 + + \uff089\uff09 Where \u00b5\u03b1 is adhesion coefficient of wheel and obstacle, the contact angle \u03b1 can be expressed in terms of the obstacle height and wheel radius r: arcsin[( ) / ]r h r\u03b1 = \u2212 \uff0810\uff09 Assuming the obstacle height 0.05m and adhesion coefficient 0.25, then the resistance will be function of wheel radius r as shown in Fig. 9: The curve in Fig. 9 shows that as the wheel diameter increasing, the obstacle resistance will decrease. From Eqs. (6)-(8) for the system shown in Fig. 8 , the normal force on front and rear wheel, NF and NR, the drawbar pull force of rear wheel, DPR, can be computed as: 2( cos sin cos sin )F WLN r r r L L a\u00b5 \u03b1 \u00b5 \u03b1 \u00b5 \u03b1 = + \u2212 + + (11) ( cos sin ) 2( cos sin cos sin )R WLN W r r r L L a \u00b5 \u03b1 \u03b1 \u00b5 \u03b1 \u00b5 \u03b1 \u00b5 \u03b1 + = \u2212 + \u2212 + + (12) (cos sin ) 2( cos sin cos sin )R WLDP r r r L L a \u03b1 \u00b5 \u03b1 \u00b5 \u03b1 \u00b5 \u03b1 \u00b5 \u03b1 + = + \u2212 + + (13) If the wheel radius is definite, NF, NR and DPR are function of obstacle height h, but hmax cannot be solved analytically in Eqs. (11)-(13) due to complex nature of the equations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000332_50003-4-Figure3.30-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000332_50003-4-Figure3.30-1.png", + "caption": "Fig. 3.30. The model running at large spin (turning and equivalent camber) at a relatively small (a,b) or large positive or negative slip angle (c).", + "texts": [ + " The next item to be addressed is the response to large spin in the presence of side slip. Figures 3.30a,b refer to this situation. Large spin with two sliding ranges occurs when the following two conditions are fulfilled. THEORY OF STEADY-STATE SLIP FORCE AND MOMENT GENERATION 129 1 and tanlal _< a l ~ l - 1 (>0) (3.77) Igol _> go/- aO 0 y y If the second condition is not satisfied we have a relatively large slip angle and the equations (3.74,3.75) hold again. This situation is illustrated in Fig.3.30c. For the development of the equations for the deflections we refer to Fig.3.30a with the camber equivalent graph. First, the distances y will be established and then the deflections V~o will be added to obtain the actual deflections v. Adhesion occurs in between the two sliding ranges. The straight line runs parallel to the speed vector and touches the boundary Ym~xR 9 The tangent point forms the first transition point from sliding to adhesion. More to the rear, the straight line intersects the other boundary YmaxL\" With the following two quantities introduced A 1 - a l go l - 1 , A 2 - alfpl + if,, (3", + "28 where the level of camber correspond to 'small' spin. It can be observed that in accordance with Fig.3.25 the force at zero side slip first increases with increasing spin and then decays. As was the case with smaller spin for the case where spin and side slip have the same sign, the slip angle where the peak side force is reached becomes larger. When the signs of both slip components have opposite signs, the level of side slip where the force saturates may become very large. As can be seen from Fig.3.30b the deflection pattern becomes more anti-symmetric when with positive spin the slip angle is negative. This explains the fact that at higher levels of spin the torque attains its maximum at larger slip angles with a sign opposite to that of the spin. The observation concerning the peak side force, of course, also holds for the slip angle where the torque reduces to zero. Spin , l o n g i t u d i n a l and side slip, the width effect The width of the contact patch has a considerable effect on the torque and indirectly on the side force because of the consumption of some of the friction by the longitudinal forces involved" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002408_tmag.2021.3064023-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002408_tmag.2021.3064023-Figure1-1.png", + "caption": "Fig. 1. Topology of the proposed machine. (a) 3-D diagram. (b) Cross-sectional view and direction of dc current injection.", + "texts": [ + " In the design of high-speed machines, the lowest possible pole pair number is usually selected to reduce the fundamental frequency and the induced loss. In this article, the number of stator/rotor slots are selected as 6/2, and the armature winding pole pair is 1. Besides, in the proposed machine, doubly salient structure is adopted similar to that of SRM, which has the advantages of high reliability, robust, and good heat dispersion [7]. And the doubly salient structure is amenable for high-speed operations. The 3-D diagram and cross-section view and direction of dc current injection of the proposed 6/2 HSVRM are shown in Fig. 1. 0018-9464 \u00a9 2021 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See https://www.ieee.org/publications/rights/index.html for more information. Authorized licensed use limited to: University of Prince Edward Island. Downloaded on June 03,2021 at 12:32:51 UTC from IEEE Xplore. Restrictions apply. However, the design of low-slot-pole HSVRM brings the shortcoming of overlong end length. Fig. 2 presents the configurations of three types of windings. In the machines that use lap winding, as presented in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002262_tia.2021.3058113-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002262_tia.2021.3058113-Figure2-1.png", + "caption": "Fig. 2. Combined radial-axial magnetic bearing that uses permanent magnets to generate an external bias field for the three-pole bearing.", + "texts": [ + "org/publications_standards/publications/rights/index.html for more information. This paper assumes the three pole bearing structure depicted in Fig. 1 is operated by the conventional three phase inverter typical of electric motor drives. The coils are connected in a Wye configuration, requiring that their currents sum to zero. In addition to airgap \u201ccontrol\u201d fields created by the coil currents, this paper assumes that an external flux source provides a bias field to the airgap. An example configuration is shown in Fig. 2, which depicts a combined radial-axial magnetic bearing (CRAMB) utilizing axially-magnetized permanent magnets to generate the bias flux. The \u201cradial stage\u201d is implemented as the three-pole bearing of Fig. 1. This particular topology is discussed at length in [24]. Other common geometry configurations to realize this are reviewed in a previous work by the authors [5], which also developed a generalized force model for the three-pole bearing and proposed an \u201cexact solution\u201d to calculate the currents necessary to create a desired force vector", + " The saturation block returns the minimum of the commanded input magnitude and the pre-computed maximum force profile for the specified force angle. The output magnitude and force angle are converted back into x and y components for use in the current lookup tables of the \u201cForce Inverse\u201d block. V. VALIDATION The proposed exact force vector regulator is now validated and compared to the conventional linear approach by studying the control implementation of an example bearing through both simulation and experimental results. The experimental results are obtained from a prototype CRAMB (see Fig. 2) Authorized licensed use limited to: Carleton University. Downloaded on June 04,2021 at 02:56:52 UTC from IEEE Xplore. Restrictions apply. 0093-9994 (c) 2021 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. that implements the externally biased three-pole bearing as its radial stage (shown in Fig. 15). A mechanical swivel bearing allows the shaft end opposite the magnetic bearing to spin and pivot but constrains it from displacing radially" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002560_tec.2021.3069096-Figure15-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002560_tec.2021.3069096-Figure15-1.png", + "caption": "Fig. 15. Flux Density of the rotor and the stator of the RB section. (a) Rated fundamental current operation (b) Twice of the rated current (c) Three times of rated current. observed from Fig.15 that the maximum flux density at peak current operation is 0.8T in the rotor and 0.85T in the stator iron. This is far lower than that of the saturation flux density of the iron material M235-35A which is around 1.25 T. It can be claimed that the RB of the machine is \u201cimmune\u201d to saturation effect. This benefits from the optimization achievements. As a result, the air-gap between the RB and the stator is relatively large. The flux blocking plastic ring between the RB and the PMSM rotor prevents the leakage magnet flux from entering into the RB. Most importantly, the big difference from that of a conventional PMSM machine is that the flux in the RB section is only generated by stator winding excitation, the overall flux-", + "texts": [ + " The gap is greater than that of the main air-gap length. In the manufacturing process, a non-magnetic plastic ring with the thickness of 1.5 mm is interlaid between the rotor end and the PM rotor. Fig. 14 shows the axial flux distribution of the whole rotor. the fluxes shooting through the rotor end are significantly reduced with the non-magnetic rotor ring. Another concern is related to the iron core saturation of the RB as the torque producing current is increased for higher torque production during machine acceleration. As shown in Fig. 15, the flux density of the rotor and stator iron core is plotted when machine is operated at rated current, twice and three times of the rated current respectively. The simulation points are selected when the rotor rotated to the time instant where maximum flux conduction is achieved, i.e. where the rotor teeth are exactly aligned to the flux conducted stator teeth. This is aimed at analyzing the maximum possible flux density of the RB machine section and the influence of the saturation caused by increased fundamental torque current" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002446_s11106-021-00183-8-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002446_s11106-021-00183-8-Figure1-1.png", + "caption": "Fig. 1. Schematic diagram of the 3D printer", + "texts": [ + " Hence, the SLM method is highly relevant for the production of parts from the Inconel 718 alloy. This is one of the few commercially attractive additive manufacturing techniques that can be used to obtain parts from nickel-based alloys with a porosity closed to zero [4\u20136]. The SLM technology uses a localized and focused laser beam for melting powder particles to form a liquid metal pool of micron sizes, which subsequently solidifies at high speed, forming a layer. The process functional diagram for the 3D printer is presented in Fig. 1. Such equipment allows creating component parts of complex geometry in layers through melting using a digital 3D model as source information. The thickness of the layer varied between 15 and 150 m depending on the material. Ytterbium fiber lasers with a power of 200 to 1000 W were used to melt metal in powder form. The radiation focuses on the desired place to form the contour of the workpiece using high-speed drive mirrors [7, 8]. The chamber is filled with inert gas (nitrogen or argon) to prevent the undesirable oxidation during production" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002945_978-3-030-69178-3_15-Figure5.5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002945_978-3-030-69178-3_15-Figure5.5-1.png", + "caption": "Fig. 5.5 Schematic view of a robot arm with collision avoidance zones around one of its links, marked by do1 and do2", + "texts": [ + " Alternatively, if soft task priorities are preferred, additional objectives with a manually specified weight w\u03bb can be used: w\u03bb\u2016e\u03bb\u20162 = w\u03bb(q\u0308 Qe\u03bb q\u0308 + Pe\u03bb q\u0308). Weights in different orders of magnitude can be chosen to enable a hierarchy of soft task priorities. However, such soft task priorities cannot strictly achieve the error dynamics, e.g., (5.7), due to the non-zero task residuals (5.12). Therefore, in case of critical situations such as collision avoidance, constraints are needed. Suppose there are m obstacles around a robot with n links, and m pairs of a constraint and a task with different proximity zones allocated to each link, as shown in Fig. 5.5. Performing continuous distance checking between the robot and an obstacle, the pair of nearest points pr ,po \u2208 R 3 belonging to the robot and the obstacle, respectively, is obtained.When the obstacle is closer than a threshold d \u2264 do1, the inequality constraint SiJ fi q\u0308 + Ce fi \u2264 0 \u21d0\u21d2 e\u0308 j + Kdj e\u0307i + Kpiei \u2264 0 (5.13) enforces the error dynamics e\u0308 + Kd e\u0307 + Kpe = 0, where e = Sinod, and no denotes the normal defined by pr \u2212 po. If the distance is defined as d(u,q) = \u2016pr \u2212 po\u20162, differentiating d(u,q) leads to the Jacobian J fi : J fi = \u2202d \u2202q = 2(pr \u2212 po) \u2202pr \u2202q " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002935_978-3-030-77102-7_10-Figure9-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002935_978-3-030-77102-7_10-Figure9-1.png", + "caption": "Fig. 9. The last iteration", + "texts": [ + " Using these tiles, EER and PER on transitions of TATPN2 as in Fig. 7, we get the swastik pattern. In TATPN2, when the transitions t1t2t3 fires, we get the pattern in Fig. 8(a) and again when the transition t4 fires we get Fig. 8(c) and again when the transition t5 fires, we get Fig. 8(d) and when the transition t6 fires, we get the swastik pattern in Fig. 8(e) and when the transition t7 fires, we get the swastik pattern in Fig. 8(f) and when t8 fires, it removes the token Q from P8 and it uses the rule \u03bb \u2192 Q(u, r, d, l) and deposits the pattern Fig. 9 in P9. When t9 fires we get the swastik pattern as in Fig. 10. We can generate different sizes of this pattern as the number of times t8 and t9 fires. Triangular Array Token Petri Net and P System 85 In this section, we develop a P system called Triangular Array Token Petri Net P System (TATPNPS), which uses Elementary Evolution Rules (EER) and Parallel Evolution Rules (PER) introduced in Sect. 3 as evolution rules in its regions and has labeled triangular arrays as objects. Definition 9. A Triangular Array Token Petri Net P System (TATPNPS) is \u03c0 = (\u03a3T , \u03bc, P1, P2, \u00b7 \u00b7 \u00b7 Pm, (R1, T1), (R2, T2) \u00b7 \u00b7 \u00b7 (Rm, Tm), i0) where \u03a3T is a finite set of labeled triangular tiles; \u03bc is a membrane structure with m membranes, labeled in a one- to-one way with 1, 2, \u00b7 \u00b7 \u00b7 m; P1, P2, \u00b7 \u00b7 \u00b7 Pm are finite sets of triangular picture patterns over \u03a3T associated with the m regions of \u03bc, i0 is the output membrane which is an elementary membrane and R1, R2, \u00b7 \u00b7 \u00b7 Rm are finite sets of evolution rules namely EER and PER associated with the m regions of \u03bc", + " P1 = , P2 = P3 = P4 = P5 = P6 = \u03c6 R1 = {(\u03bb \u2192 C(rd), here), (\u03bb \u2192 A(ld), here), (\u03bb \u2192 D(lu), in)} R2 = {(\u03bb \u2192 [E(u);H(r);F (d);G(l)], here), (\u03bb \u2192 [F (lud);G(rud);E(rdd);H(ldd)], in)} R3 = {(\u03bb \u2192 (H(c2);F (c4);G(c6);E(c8)), in)} R4 = {(\u03bb \u2192 (G(rud);E(rdd);H(ldd);F (lud)), in))} R5 = {(\u03bb \u2192 T4(u, r, d, l), here), (\u03bb \u2192 T4(c2, c4, c6, c8), here), (\u03bb \u2192 T4(c2, c4, c6, c8), out), (\u03bb \u2192 T5(u, r, d, l), here), (\u03bb \u2192 T5(c2, c4, c6, c8), here), (\u03bb \u2192 T5(c2, c4, c6, c8), out)} R6 = \u03c6. The picture derivation is shown in the following table: Region (i) Input Rule (Ri) Output (Ti) 1 (\u03bb \u2192 C(rd), here) 1 (\u03bb \u2192 A(ld), here) Fig. 8(b) 1 Fig. 8(b) (\u03bb \u2192 D(lu), in) Fig. 8(a) 2 Fig. 8(a) (\u03bb \u2192 E(u);H(r); Fig. 8(c) F (d);G(l), here) 2 Fig. 8(c) (\u03bb \u2192 F (lud);G(rud); Fig. 8(d) E(rdd);H(ldd), in) 3 Fig. 8(d) (\u03bb \u2192 H(c2);F (c4); Fig. 8(e) G(c6);E(c8), in) 4 Fig. 8(c) (\u03bb \u2192 G(rud);E(rdd); Fig. 8(f) H(ldd);F (lud), in) 5 Fig. 8(f) (\u03bb \u2192 T4(u, r, d, l), here) Fig. 9 5 Fig. 9 (\u03bb \u2192 T4(c2, c4, c6, c8), here) Fig. 10 5 Fig. 10 (\u03bb \u2192 T5(u, r, d, l), here) Fig. 10 is (\u03bb \u2192 T5(c2, c4, c6, c8), here) magnified (\u03bb \u2192 T4(u, r, d, l), here) Fig. 10 is (\u03bb \u2192 T4(c2, c4, c6, c8), out) magnified (\u03bb \u2192 T5(u, r, d, l), here) Fig. 10 is (\u03bb \u2192 T5(c2, c4, c6, c8), out) magnified 6 Magnified R6 = \u03c6 Magnified Fig. 10 Fig. 10 swastik pattern Thus TATPNPS \u03a01 generates a family of swastik patterns. Theorem 2. TPLm(TATPNPS) \u2229 L(TATPN) = \u03c6. Triangular Array Token Petri Net generates Swastik patterns as given in Theorem 1", + " \u03b8 = { # # # a , # # a a , # # a # , # a # a , a a a \u2666 , a a \u2666 \u2666 , a a \u2666 a , a # a # , a \u2666 a \u2666 , \u2666 \u2666 \u2666 a , \u2666 \u2666 a a , \u2666 \u2666 a \u2666 , \u2666 a \u2666 a , a \u2666 a a , \u2666 a a a , \u2666 a \u2666 \u2666 , a a \u2666 \u2666 , a \u2666 \u2666 \u2666 , a a # # , a # # # } Hence PAL-LOC is properly contained in PATPNS. Example 7. Consider a PATPNS C = (P, T, I,O), generating a local partial array language given in Example 5. Let S\u2666 = # # # b , Q1 = [ # b ] , Q2 = [ # # ] , Q3 = [ # B # ] , B \u2208 \u0393 1\u00d7n p , where B = a (\u2666)r b, Q4 = (#)m. PATPNS generating the 2nd member of the partial array language namely # # # # # # # b b b b # # a \u2666 \u2666 b # # a \u2666 \u2666 b # # a a a b # # # # # # # is given in Fig. 9 as an example. Theorem 5. PATPNS is closed under projection. We consider a partial array token Petri Net structure C = (P, T, I,O) generating the partial array language PL. Let \u03c0 : \u0393p \u2192 \u03a3p be a projetion such that \u03c0(a) = \u03b1, a \u2208 \u0393p, \u03b1 \u2208 \u03a3p. Without loss of generality \u0393p \u2229 \u03a3p = \u03c6. We can construct a PATPNS C \u2032 = (P \u2032, T \u2032, I \u2032, O\u2032) such that PL(C \u2032) = PL, where T \u2032 = {t\u20321, t \u2032 2, . . . t \u2032 k}, t\u2032i = {A\u2032 | Q\u2032 i, A \u2032 \u2212 Q\u2032 i/A \u2032 ij = (\u03c0(A))ij , (Q\u2032 i)rs = (\u03c0(Qi))rs}, Partial Array Token Petri Net and P System 151 1 \u2264 i \u2264 k, A,Qi \u2208 \u0393 \u2217\u2217 p , A\u2032, Q\u2032 i \u2208 \u03a3\u2217\u2217 p ", + " The Output console displays the following information: Current iteration, the action of the agent in the form: (index, (posX, posY ), value, action) \u2013 for example: (1, (1, 5), 4449, \u2018A\u2019) = agent 1 at the position (1,5) updates Alpha on the blackboard with value 4449, (4, (2, 6), \u2018L\u2019, \u2018m\u2019) = agent 4 at the position (2,6) is moving to position Left (Fig. 8). The simulation terminates when the termination criterion, the iteration reaches the maximal value, or when no more delta agent can move. In the Fig. 9, On Numerical 2D P Colonies with the Blackboard and GWO 175 there can be seen that agent one has found the best solution \u2013 number 4449, and that agent 5 is dead (it is in the final configuration). The simulation of the GWO algorithm by the numerical 2D P colony with the blackboard works, and it gives the desired results, it finds the optimal solution However, one must say that the original GWO algorithm works faster. The main issue is, that the wolves in the GWO are \u201cgetting faster\u201d as the number of the iterations of the algorithm increases, but the agents of the P colony can move only one step further in each derivation step" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000468_20050703-6-cz-1902.01340-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000468_20050703-6-cz-1902.01340-Figure3-1.png", + "caption": "Fig. 3. Constraint based on relationship between obstacle and leg position", + "texts": [ + " [ xs q \u2264 xl q ] \u21d2 [ xs q \u2264 Gx \u2264 xl p ] (3) [ xs q > xl q ] \u21d2 [ xl q \u2264 Gx \u2264 xs p ] (4) Here, Gx is a value in which the center of gravity of the biped robot is projected to the X-axis coordinates, xs q, xs p is a value in which the position with the tiptoe of the idling leg is projected to the X-axis coordinates, and xl q, xl p is a value in which the position with the tiptoe of the saddle support is projected to the X-axis coordinates. Moreover, the idling leg and the saddle support must change depending in the state of mode explained in the foregoing paragraph. 2.2.3. Constraint Condition from Obstable Next, the constraint condition when the biped robot steps over and exceeds the obstacle is led. Figure 3 shows the appearance when the idling leg steps over and exceeds the obstacle. (a) shows the appearance before the obstacle is stepped over and exceeded. It is necessary to carry the tiptoe of the idling leg of the biped robot to the upper part of vestibular ganglion of the obstacle after this while avoiding contact with the obstacle. There is no problem if the height of tiptoe Q of the idling leg is higher than that of ground when Q on the right side from upper right O of the obstacle. However, the idling leg should exist in the over of O if Q is at the left of O", + " Such a constraint conditions are expressed as follows [ xp \u2265 xr ] \u21d2 [ zp \u2265 zr ] (7) [ xp < xr ] \u21d2 [ h(x f , z f , \u03c6) \u2265 0 ] (8) where, h(x f , z f , \u03c6) = ( xr \u2212 x f ) sin \u03c6 \u2212 (zr \u2212 z f ) cos \u03c6 \u2212 r 2 . The current constraint conditions were a physical constraints that accompanied the stepping over excess of the biped robot. However, the constraint concerning the torque and the velocity exists generally in the motor that drives the biped robot. Then, the equation of motion expressed by the position and the angle is expressed in the following first order models based on the knee joint of the biped robot in Fig. 3. x(k + 1) = Ax(k) + Bu(k) (9) x(k) = [ x f (k), z f (k), \u03c6(k) ]T u(k) = [ ux f (k), uz f (k), u\u03c6(k) ]T A = ( 1 \u2212 \u2206t \u03c4 ) I3 B = ( \u2206t \u03c4 ) I3 The constraint concerning the torque and the velocity of the biped robot is expressed by putting the limitation on the input deflection every sampling time intervals. Then, the constraint condition is expressed as follows |u(k) \u2212 u(k \u2212 1)| \u2264 [ Mx f , Mz f , M\u03c6 ]T . (10) Here, Mx f , Mz f andM\u03c6 are the maximum movable amount of knee joint to direction of X-axis, the maximum movable amount of knee joint to direction of Z-axis, the maximum rotatable amount of knee joint leg, respectivery", + " Figure 6 shows the appearance expressed in animation based on the simulation result. In Fig. 4, the solid line is a center of gravity position, the dotted line is an upper bound at the center of gravity position, and the broken line is a lower bound at one. The constraint requirement expressed by equation (3) and (4) is met from the center of gravity position is always in the bound pair. Figure 5 (a-1) and (a-2) show the appearance of the constraint condition expressed by (5) and (8) equation when the leg shown in the bold line of Fig. 3 is an idling leg, and (b-1) and (b-2) show the constraint condition expressed by (5) and (8) equation when the leg shown by the thin line of Fig. 3 is an idling leg. Besides, the dash line in figure is a supplementary line to see whether to meet the constraint condition requirement. It is understood that it is zq > 0 according to (a-1) when xq is at the right of point O the obstacle, which means point O equal 0.67[m] or more. This shows constraint condition (5), and means it meets the requirement. Moreover, it is understood that it is g(x f , z f , \u03c6) > 0 when xq is at the left of point O the obstacle, which means point O is less than 0.67[m]", + " It is understood that it is zp > zr according to (a-2) when xp is at the left of point R the obstacle, which means point R equal 0.47[m] or more. This shows constraint condition (7), and means it meets the requirement. Moreover, it is understood that it is h(x f , z f , \u03c6) > 0 when xp is at the left of point R the obstacle, which means point R is less than 0.47[m]. This shows constraint condition (8), and means it meets the requirement. The constraint condition (5) and (8) is also filled from (b-1) and (b-2) for the leg expressed by thin line of Fig. 3. Figure 6 shows the animation which is the stepping over excess of the obstacle by the biped robot. Then it divides into mode1 to mode4 shown in Fig. 1. It can be confirmed that the biped robot steps over and has exceeded the obstacle by the lead of the constraint condition, and the best generation of the control input to the robot that meets the requirement from Fig. 6 visually. The achievement of the stepping over excess of the obstacle by the biped robot was able to be confirmed by simulation work" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000501_978-0-585-35228-2_7-Figure7.12-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000501_978-0-585-35228-2_7-Figure7.12-1.png", + "caption": "Figure 7.12", + "texts": [ + " Most kinetic models use the rigid body segment representation of the locomotor system, but must also define the distribution of the mass within the segments and the points about which moments are calculated. The majority of commercial gait analysis systems use multilink segment models which have either spherical or hinge joints connecting the segments. Thus in these types of models the joint rotation centres are fixed with respect to the coordinate systems embedded in each segment. Each segment may be considered as a free body (Figure 7.12). The linear accelerations of the centre of mass, the angular orientation, velocities and accelerations are known from the optometric measurements, their subsequent processing and use of the kinematic model. The resultant force and moment acting at the distal end of the segment are assumed to be known (from analysis of the distal segment). Considering the six equations of dynamic equilibrium it is possible to determine the reaction forces and moments acting at the proximal joint. Free-body diagram of a segment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002838_012021-Figure9-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002838_012021-Figure9-1.png", + "caption": "Figure 9. Total Deformation in Structural Steel", + "texts": [ + " Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.1.2. Stress Distribution The Max. and Min. Stress Distribution in Al 6061 T6 is 63.7 MPa and 0.3898 MPa respectively shown in Figure 7. 3.1.3. Strain Distribution The Max. and Min. Strain Distribution in Al 6061 T6 is 0.00096883 and 0.0000076477 respectively shown in Figure 8. 3.2. Analysing Testing Result of Structural Steel 3.2.1. Total Deformation The Max. and Min. Total Deformation in Structural Steel is 0.18644 mm and 0 mm respectively shown in Figure 9. The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.2.2. Stress Distribution The Max. and Min. Stress Distribution in Structural Steel is 183.73 MPa and 1.8481 MPa respectively shown in Figure 10. 3.2.3. Strain Distribution The Max. and Min. Strain Distribution in Structural Steel is 0.0010678 and 0.000012423 respectively shown in Figure 11" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000469_50009-6-Figure8.32-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000469_50009-6-Figure8.32-1.png", + "caption": "Figure 8.32. Lateral view of a manatee swimming, tracings of body, limb, and tail movements. Arrows indicate direction of movement. (From Hartman, 1979.)", + "texts": [ + " Nearly all of our information on sirenian swimming comes from Hartman\u2019s (1979) work on the West Indian manatee. Compared to cetaceans, manatees are poor swimmers and are unable to reach or sustain high speeds. According to Hartman (1979) movement is initiated from a stationary position by an upswing of the tail followed by a downswing, repeated until undulatory movement is established. Each stroke of the tail displaces the body vertically, the degree of pitching increasing with the power stroke (Figure 8.32). The tail also serves as a rudder. Cruising animals can bank, steer, and roll by means of the tail alone. The use of the flippers in locomotion differs somewhat from cetaceans. While cruising, the flippers of adult manatees are held motionless at the sides. Juvenile manatees have been reported to swim exclusively with their flippers (Moore, 1956, 1957). Flippers (either independently or simultaneously) are normally used only for precise maneuvering and for corrective movements to stabilize and orient the animal while it is feeding" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003259_978-981-16-1769-0_32-Figure18-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003259_978-981-16-1769-0_32-Figure18-1.png", + "caption": "Fig. 18 Proposed layout", + "texts": [ + " Hence, with all ergonomic considerations, the new design of the fixture is proposed. The solid model of the fixture with all modifications is shown in Fig. 17. The new modified fixture has been designed at the height of 900 mm which is considering the anthropometry height of standing posture. The fixture is so designed that it can be tilted to any angle by the difference of 10 degrees which will eliminate the unnecessary kneeling angles and would improve neck position. Ergonomic Workstation Design for Welding Operation\u2014A Case Study 239 Figure 18 shows the proposed design of the layout of the welding workstation of turntable which consists of material storage, material handling devices, the position of the fixture, and the position of the finished job storage station. Overhead fixtures with rollers are proposed in the modified design to hold and convey the welding torch with exhaust gases hood for ventilation. This can be provided which would improve the working condition of the workers. All the industries are now concerned about the performance improvement of the shop floor operations through macro and micro modifications in their processes", + " The deformation in equal sense loading is followed by torsional effects which is also observed during the experiments. The maximum stress and strain at the fold along with maximum free end displacement, obtained from FEA, are validated as shown in Table 5. After the snap, although experimentally and analytically the strain at the fold is noted to be increasing, it is not seen in the FEA (see Figs. 16, 17). The reason behind this is the excessive and unrealistic X deformation due to the momentum gained after the snap, which shifts the fold away from the fixed end (see Fig. 18). This shows the need for additional constraints to capture the post-snap behavior accurately. Behavior Study of Tape Springs for Space Deployment Applications 361 \u2022 Before the snap has occurred in the opposite sense loading, for the same amount of applied load, lower stresses and strains at the fold are noted. Also, the load at which the snap and fold formation occurs is higher in the opposite sense. Thus, opposite sense loading results in higher deployed stiffness and lesser deformations at the same loads (prior to the snap), reducing the effects of disturbances in the deployed state" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002874_s10846-021-01410-5-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002874_s10846-021-01410-5-Figure4-1.png", + "caption": "Fig. 4 Illustration of pre-grasp object angle (\u03b8pre) and post-grasp object angle (\u03b8post), relative to the horizontal axis of the image (horizontal). Two examples of grasp outcome are shown", + "texts": [ + " The orientation error score OE compares the rotational change of the object: OE \u00bc 1\u2212 j \u03b8post\u2212\u03b8pre j 180 \u00f02\u00de Stepper motor NEMA 23 SY57STH76-2804A [72] \u2022 Holding torque: 19 kg-cm \u2022 200 steps per revolution where \u03b8pre is the major orientation of the object pre-grasp and \u03b8post is the major orientation of the object post-grasp. \u03b8pre and \u03b8post are measured between the horizontal axis of the image and the major axis of the object before and after a grasp attempt, respectively, and range from [0\u00b0, 180\u00b0]. This metric scores the orientational change of an object introduced by the executed pick-and-place action. It should be noted that it does not respond to translational error\u2014as illustrated in Fig. 4. OE ranges from [0, 1]. The relationship between this score and the angular difference between pre- and post-images is linear. As the angular change tends toward 180\u00b0,OE tends linearly toward 0. If no orientation change is measured, OE produces a value of 1. The proposed metrics may be of interest to other similar works for improvement and as an objective baseline for comparison between methodologies\u2014since they are not dependent on the system or object test pool. OS and OE measurement is relatively simple, requiring only a single top-down RGB camera and basic DIP processing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001131_icsma.2008.4505635-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001131_icsma.2008.4505635-Figure5-1.png", + "caption": "Fig. 5 DMP method and image method Since the image charge is along the line of the applied charge vector with the different distance, the position of the image charge can be expressed in terms of the ratio of lengths AR from (14) and unit vector of", + "texts": [ + " The magnetic charge with the strength m at P(O, acoso, asino) can be expressed in the xyz local frame and the potential on the spherical surface is given by 1 m 4)7=- 2 +2 _aco +/-2 +2 _aR cos0 To satisfy (12), unknown distance and strength a, for the image charge in (13) can be expressed by a=R2 a or a=a where a = R21 a can be chosen since the ima charge should be outside of the sphere. Similarly, the strength of the image charge m c be given by m = -Rmla Given the parameters in (14) and (15), the sca potential function TD inside circle can be expressed follow: For 0 < r < R region: mr 1 Rla x2+(y a)2 Z2 VX2+(y (R2 a))2 z2 From (16), three practical cases shown in Fig. 5 considered here to illustrate the combination of 1 DMP and image method since an electromagne actuator generally consists of rotor, magnet and stator. Case 1 (iron rotor): PM is outside the sphere Case 2 (iron stator shell): PM is inside sphere shell Case 3: Combination ofboth Case 1 and Case 2 Case 1 considers a pair of the DMP model for permanent magnet with the strength m and ro boundary of a radius rR in Fig. 5(a). In (16), 1 position of each source/sink (xl, yl, z1) can be express in the spherical coordinate as follows: -X1i- Cos 01 Cos01 Y[ = r[ sin 0 coss _Zhy - sitn where r, = 2+ y2 + Z2 ;0j1= tan-l (y/ xl); and 0 = cos-1 (zj1 rl ) (13) the applied charge from (17). Fxil LCos 01 Cos 1 ,an Yl = ~~~~A,r, sin 01 cos 01,an Lz iR sin (15) where AR =rR ir' ;rR is the ra [lar (subscript R indicates the rotor). as From (15), the strength of the inr given in terms of AR by mR =-m AR (18) idius of the rotor nage charge is also (19) In Case 2, the image charge of the stator boundary with the outer radius rs in Fig. 5(b) can be expressed in terms of the ratio As = r / r, as follows: -x-l - cos O1 sin 1 l = Ass sin 01 sin 1 (20) _Z1_ S Csco01 Similarly, ms = -Asm (21) where the subscript S indicates the stator. Case 3 includes both the iron rotor and stator. Since the strength and position of image charge of each case are found, Case 3 can be obtained using the principle of superposition of the solution in Cases 1 and 2. The potential functions including each boundary can be expressed as summing each potential function of the corresponding image charges" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002747_13506501211016108-Figure13-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002747_13506501211016108-Figure13-1.png", + "caption": "Figure 13. Temperature field in the textured journal bearing (19 grooves, height of 5 \u03bcm, variable pressure) compared with the untextured journal bearing.", + "texts": [ + " The remaining lubricant fluid in the cavitation area was thrown over the wall of the bearing and the oil vapor bubbles stayed closer to the shaft, which was due to the difference in density. A critical influence on the bearing material was the amount of heat generated. In the solid parts of the bearing system, the mass movement of a rotating fluid caused a heat transfer convection mechanism when the heated fluid moved away from the source of heat and carried energy with it. The temperature distribution across the solid parts is shown in Figure 13 by using a color contour that predicted the intensity and distribution of temperature in the textured and untextured journal bearings for different lubricant feed pressures. In this figure, according to the right-angled coordinate system, which was in full compliance with the coordinate system presented in Figure 1, the location of the cavitation area (converging region) was specified. Blue showed the minimum value and Red showed the maximum value of the journal bearing temperature under thermal equilibrium conditions. The maximum values of temperature over the shaft and bearing surfaces occurred near the inlet lubricant feed pressure in the x\u2013y-plane located in the upper half of the cylindrical bearing (between 0\u25e6 and 175\u25e6). In the qualitative comparison, the angular range of the surface areas with the highest temperature (red) near the inlet lubricant feed pressure was reduced for the textured bearings compared with plain bearings (Figure 13(a) to (c)). The results showed that the maximum value of temperature was higher and the minimum value of temperature had a small change for the untextured bearing system compared with the textured one, which showed better heat loss to the surrounding fluid medium and increased thermal performance of solid parts in the textured bearing. The presence of a textured surface on the solid bearing caused a further increase in heat transfer from the solid part of the bearing to the lubricant fluid in the fluid\u2013solid interface region", + " As a result, the solid part of the textured bearing operated at a lower temperature and the number of high-temperature points on the bearing structure and the shaft was reduced compared with the untextured journal bearing. Meanwhile, by increasing the lubricant feed pressure, the maximum temperature value increased for the textured and untextured bearing systems. The heat exchange at the textured bearing walls to the fluid is an important factor for predicting the heat transfer coefficient in the bearing system arrangement. According to the results presented in Figure 13, the percentage increase in the amount of total heat transfer of the textured and untextured bearing surfaces at different lubricant feed pressures is given in Figure 14 concerning the lubricant feed temperatures. Comparisons were made based on the increase in lubricant feed temperature relative to the temperature of 310 K. The percentage increase in the amount of total heat transfer monotonically increased with the lubricant feed temperature factor for the textured and untextured bearing surfaces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001413_j.jsv.2008.02.056-Figure11-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001413_j.jsv.2008.02.056-Figure11-1.png", + "caption": "Fig. 11. Loading and meridional stress assumptions with axial restraint at R1.", + "texts": [ + " (75e) For ring element N, we have D10 \u00bc d10 \u00fe d1T \u00feHB2d1H 002 \u00fe TB4d1T 004 , (76a) D20 \u00bc d20 \u00fe d2T \u00feHB2d2H 002 \u00fe TB4d2T 004 , (76b) D30 \u00bc d30 \u00fe d3T \u00feHB2d3H 002 \u00fe TB4d3T 004 , (76c) D40 \u00bc d40 \u00fe d4T \u00feHB2d4H 002 \u00fe TB4d4T 004 , (76d) ARTICLE IN PRESS T.A. Smith / Journal of Sound and Vibration 318 (2008) 428\u2013460450 DV0 \u00bc dV0 \u00fe dV T \u00feHB2dV H 002 \u00fe TB4dV T 004 , (76e) where the double primes on H2 and T4 denote unit values per unit of circumferential length. The loadings and meridional stress diagrams for pressure and axial forces with axial restraint of the ring element at the edge of the ring associated with the radius R1 are shown in Fig. 11. The force Vp1 shown in Fig. 11 is given by V p1 \u00bc p\u00f0R2 2 R2 1\u00de 2R1 . (77) By using Eqs. (57) to define Hp T and Vp T and considering only the forces due to the pressure loading p in Fig. 11, the torque Tp, considered positive for clockwise rotation of the cross section, is found to be TT p \u00bc V T p \u00f0R1 R3\u00de HT p \u00f0R3 R\u00de cot a \u00bc p\u00bd\u00f0R2 2 R2 1\u00de\u00f0R1 R3\u00de h2 \u00f0R1 \u00fe R2\u00de\u00f0R3 R\u00de=\u00f0R2 R1\u00de. Thus, Tp \u00bc TT p =2pR; resulting in Tp \u00bc p\u00f0R2 2 R2 1\u00de\u00f0R1 R3\u00de 2R ph2 \u00f0R1 \u00fe R2\u00de\u00f0R3 R\u00de 2R\u00f0R2 R1\u00de , (78) ARTICLE IN PRESS T.A. Smith / Journal of Sound and Vibration 318 (2008) 428\u2013460 451 and the total torque T01 due to combined pressure loadings p and axial load L, by using Eq. (59), is found to be T01 \u00bc Tp \u00fe TL \u00bc p\u00f0R2 2 R2 1\u00de\u00f0R1 R3\u00de 2R ph2 \u00f0R1 \u00fe R2\u00de\u00f0R3 R\u00de 2R\u00f0R2 R1\u00de \u00fe LR1\u00f0R2 R1\u00de R " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002102_tla.2021.9451234-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002102_tla.2021.9451234-Figure1-1.png", + "caption": "Fig. 1. Prot\u00f3tipo quadrirrotor constru\u00eddo.", + "texts": [ + " Apresenta-se ainda, as contribui\u00e7\u00f5es sintonizadas ZNAQ, ZNAQC e ZNAQCV, com suas respectivas valida\u00e7\u00f5es. O prot\u00f3tipo foi constru\u00eddo com uma estrutura S500, com quatro bra\u00e7os em configura\u00e7\u00e3o em X, sendo esta a configura\u00e7\u00e3o dominante em prot\u00f3tipos cargueiros. Os bra\u00e7os s\u00e3o unidos atrav\u00e9s de um conjunto central, que tamb\u00e9m suporta os demais componentes do quadrirrotor, como a controladora de voo, a bateria e o r\u00e1dio receptor, podendo tamb\u00e9m suportar a instala\u00e7\u00e3o de outros acess\u00f3rios, como pulverizadores de saneantes. O prot\u00f3tipo constru\u00eddo \u00e9 apresentado na Fig. 1 Os propulsores s\u00e3o compostos individualmente por um motor brushless MT2216-810 KV e uma h\u00e9lice 10x4,5\u201d, com for\u00e7a de empuxo equivalente a 0,95 kg. O controlador eletr\u00f4nico de velocidade dos motores \u00e9 o Rsky 40A opto acoplado. A controladora de voo \u00e9 a NM Lite. A bateria \u00e9 de 4 c\u00e9lulas Lipo, com 5000 mA e 14,8 V. A comunica\u00e7\u00e3o \u00e9 composta por um r\u00e1dio transmissor modelo FS-i6S operando 10 Giga Hertz com o r\u00e1dio receptor FS-iA10B. O prot\u00f3tipo quadrirrotor foi validado atrav\u00e9s de voo experimental, alcan\u00e7ando efetiva capacidade de voo e realiza\u00e7\u00e3o de manobras, com resultados satisfat\u00f3rios" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001001_aim.2008.4601625-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001001_aim.2008.4601625-Figure2-1.png", + "caption": "Fig. 2 Simple Quadruped Robot Mechanism.", + "texts": [ + " The paper is organized as follows: section 2 presents basic four legged mechanism and kinematics equations. Section 3 briefly introduces the gait generation types. Section 4 shows the simulation results. Section 5 summarizes the results and gives the conclusion. In general, for a machine to walk, at least 2 Degrees of Freedom (DOF) leg is required. One of them is for lift up and land movement of the leg and the other one is for swing the leg back and forth. Most of the quadruped robots have three DOF legs, because of more maneuvering capability. A basic three DOF mechanism is shown in Fig. 1. From Fig. 2, basic position vectors of each leg can be expressed in global coordinate frame: ),,( Bi o Bi o Bi o i o zyxB = ),,( Ti o Ti o Ti o i o zyxT = , 4,3,2,1=i (1) From Fig. 1, We have: 2 2 1 3 3 2 2cos( ) cos( ) ( ) ( )o o o o i i i i i Ti Bi Ti Bia a a x x y y\u03b1 \u03b1+ + = \u2212 + \u2212 978-1-4244-2495-5/08/$25.00 \u00a9 2008 IEEE. 1 Proceedings of the 2008 IEEE/ASME International Conference on Advanced Intelligent Mechatronics July 2 - 5, 2008, Xi'an, China 4 3 3 2 2sin( ) sin( )i i i i i ia a a h\u03b1 \u03b1+ + = (2) 1tan( ) o o Ti Bi o o Ti Bi y y x x \u03b1 \u03ba \u2212= \u2212 where 1=\u03ba for 2=i or 3, 1\u2212=\u03ba for 1=i or 4 and where ih denotes the initial height of i oB " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002838_012021-Figure22-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002838_012021-Figure22-1.png", + "caption": "Figure 22.Stress Distribution in Basalt Fiber", + "texts": [ + " Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.6. Analysing Testing Result of Basalt Fiber 3.6.1. Total Deformation The Max. And Min. Total Deformation in Basalt Fiber is 0.19542 mm and 0 mm respectively shown in Figure 21. The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. 3.6.2. Stress Distribution The Max. And Min. Stress Distribution in Basalt Fiber is 62.804 MPa and 0.37819 MPa respectively shown in Figure 22. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.6.3. Strain Distribution The Max. And Min. Strain Distribution in Basalt Fiber is 0.00074223 and 0.0000057032 respectively shown in Figure 23. 3.7. Analysing Testing Result of Carbon Fiber 3.7.1. Total Deformation The Max. And Min. Total Deformation in Carbon Fiber is 0.68523 mm and 0 mm respectively shown in Figure 24. The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000961_s11668-007-9014-8-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000961_s11668-007-9014-8-Figure1-1.png", + "caption": "Fig. 1 Failed camshaft and crankshaft gears. Note the broken and missing teeth", + "texts": [ + " The hardness profiles were used to determine the depth of the nitrided layer. According to the Chinese standard (GB 1543-89) [1], the depth where the hardness values measured are equal to the value of core hardness +HV50 is defined as the depth of the nitrided layer. Z. Yu (&) X. Xu Electromechanics and Material Engineering College, Dalian Maritime University, Dalian 116026, P.R. China e-mail: zhiweiyu@newmail.dlmu.edu.cn X. Xu e-mail: xxiaolei@dlmu.edu.cn Visual Observations The failed camshaft and crankshaft gears are shown in Fig. 1. The turning direction of the crankshaft and drive and driven relationship between the two gears establish the turning directions of both gears. For the failed camshaft gear, adjacent teeth fracture appears in four regions (marked 1, 2, 3, and 4 in Fig. 1, 2). In each region, more than one tooth fractured (for example, labeled 4-1, 4-2, 4-3, 4-4 in Fig. 2), and plastic deformation of adjacent teeth appears in four regions (marked 5, 6, 7, and 8 in Fig. 1, 3). In each region, more than one tooth face exhibits severe plastic deformation. For the crankshaft gear, four adjacent teeth fracture regions (marked A, B, C, and D in Fig. 1, 4) and two adjacent teeth are plastically deformed (regions marked E and F in Fig. 1, 5). It is worth noting that some metal fragments were imbedded in the grooves between teeth in two regions of plastic deformation on the crankshaft gear and in one region on the camshaft gear. The length of the fragments corresponds to the face width. When the two gears were engaged according to their turning directions, the orientation of plastic deformation bands on the two failed gears was in correspondence with each other as shown in Fig. 6. Such an appearance may result from fractured teeth crushing onto the tooth faces and being embedded into the groove between teeth" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002804_j.matpr.2021.04.568-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002804_j.matpr.2021.04.568-Figure5-1.png", + "caption": "Fig. 5. Fatigue analysis of E-Glass/Epoxy.", + "texts": [], + "surrounding_texts": [ + "CAD models are designed in CATIA which contains special tools for the generation of traditional surfaces which will then be transformed into solid models, using traditional and composite monoleaf spring materials. The measurements of a spring leaf of a TATA SUMO vehicle are used for the design of the mono-leaf spring. Table1 displays the mono leaf spring configuration parameters and the planned mono leaf spring CAD model, as shown in Fig. 1. Finite Element Analysis (FEA) is a computational method for deconstruction into very small elements of a complex structure. In the simulated world, ANSYS offers an affordable way to examine the success of goods or processes. Digital prototyping is called this method of product growth. Users will iterate different scenarios using simulated prototyping techniques to refine the software even before the production is launched. This allows the probability and expense of failed designs to be reduced. In this study, a model was developed that was imported into the ANSYS workstation and FEA analysis. Meshing is the mechanism where the entity is divided into very small pieces called components. Elements. It is sometimes called a piece by piece. The leaf spring model is here meshed with a 10 mm brick mesh part scale. The front end is restricted and the rear end only in Y and Z is restricted; translational movement is permitted in X direction. Conditions of loading include applying a force on the middle of the spring of the leaf upward on the base of the leaf spring. The loading range is between 1000 N and 5000 N. The meshed model and limit and conditions of loading of a leaf spring are seen in Figs. 2 and 3." + ] + }, + { + "image_filename": "designv11_83_0000534_s0065-2458(08)60221-1-Figure19-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000534_s0065-2458(08)60221-1-Figure19-1.png", + "caption": "FIG. 19. Basic ball element. [From ref. 18.3", + "texts": [ + " Down to this size the curves of Fig. 18 give a good description of the influence of miniaturization. No publication reporting commercial use of valves for doing logic is known. There is a general feeling that one single technique will not be applied before a much larger field comprising other moving and nonmoving part approaches is explored. 2.1 .W Ball Elements A very interesting approach to logic devices using moving parts works with balls. Its obvious simplicity probably cannot be exceeded. Figure 19 shows the basic arrangement : A ball moves freely, but not loosely, in a cylindrical housing having four tubular connections A , B, X , and Y. The best way to explain its operation is to consider the element incorporated in a complete circuit as shown, for instance, in Fig. 20. The terminals A and B are supplied through separate 187 H. H. GLAETTLI l /x FIQ. 20. Bistable circuit with ball element. [From ref. 18.1 flow restricting orifices RA and RB by a common pneumatic supply at a pressure pa. The radial connections X and Y are open to the atmosphere", + " Therefore, a compromise between response time and stability becomes necessary for a given pressure, a fact that is true for all bistable elements. An alternative possibility of control exists by applying trigger pulses to X or Y , respectively, rather than closing one of these connections. This avoids the leakage path around the ball as a way to introduce the energy required to start switching. The fact that a control signal of a finite power is necessary can be disregarded. Pressure gains of up to 100 are reported to be feasible [18]. An interesting variant of the element in Fig. 19 is shown in a symbolic representation in Fig. 21. A third connection Z allows a third stable position, a very rare fact in the world of memory elements (no commercial use of tristable elements is known at this time). The same method of obtaining bi- or tristability with the aid of resist189 ances can also be applied to valves. They, however, can then no longer be considered as purely static switching means under these circumstances, and tristability requires a more complicated geometry for valves than for ball elements", + " 12 : for dc operation the output signal obtained at A1 is in phase with the input signal fed to B1 and Yl, and the output signal of the second element obtained at Yz is inverted with reference to the input signal present a t A2 and X 2 . The two elements, however, are not as symmetrical as the two spool valves: the pressure at A1 varies between pa 191 and a certain minimum value, whereas the pressure a t Yz varies between 0 and a maximum pressure lower than p, . To a certain extent, the ball elements shown in Fig. 25 represent a variation of the basic element shown in Fig. 19 and Fig. 20. Leakage around the ball is no longer required, and only one seat must be provided since the element is used as a three terminal device. Input and output paths are well separated as far as fluid flow is considered; proper relations, however, must exist between the pressure levels. As there is fluid flow only in one position, and as no steady control flow is required, the element may be called semistatic. what higher than that a t Az, X I , and A1, respectively. This docs not affect appreciably the first ball in its number two position, as long as connection 21 is open" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000602_978-3-540-74764-2_52-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000602_978-3-540-74764-2_52-Figure2-1.png", + "caption": "Fig. 2. XYZ-Positioning Sensor for the AFM-tool", + "texts": [ + " The output voltage depends directly on the intensity of the external magnetic field, which is tuned to transfer a power output of 330 mW / 3.3 V from the coil power pack. The local position sensor is based on an AFM scanner mounted on a rotor high motion positioning actuator. The AFM sensor consists of one position scanner and one cantilever. The scanner is made of 4 PZT stack actuators that permits movements with 3 DOF (x,y,z) and the cantilever contains a force sensor based on a piezoresistance (see Fig. 2). The AFM tool consists of three main components: 1.) the AFM probe with the integrated piezoresistance, 2.) an AFM holder for easy probe exchange and 3.) the XYZ scan stage. A rotational drive has been developed, evaluated and redesigned to be a suitable interface between robot and tool. Five rotational drives with specific rotors for each tool have been built (see Fig. 3). Furthermore, the rapid prototyping technique for making multilayer piezoceramics, developed in the beginning of the project, has been utilised to make the drivers for the grippers" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002783_09544062211012724-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002783_09544062211012724-Figure1-1.png", + "caption": "Figure 1. 3-RPR PM with clearances and loads.", + "texts": [ + " The advantage of the proposed method is clearly presented using the comparison with the existing method. Moreover, the effect of employed forces on singularities is also analyzed. Finally, we conclude the paper in the last section. Deviation arisen from clearances and loads A 3-RPR parallel mechanism (PM) consists of a manipulator and 3 identical RPR limbs evenly distributed around the manipulator. Generally, the prismatic joint of the RPR limb is the actuator. A 3-RPR PM with multiple joint clearances under loads is modeled, as shown in Figure 1. In Figure 1, the locked prismatic joints are represented by linear elastic elements and denoted as ki\u00f0i \u00bc 1; 2; 3\u00de. The manipulator is a triangle (A1A2A3) in shape and the revolute joints B1, B2, and B3 are fixed to the base. A body coordinate system O1 x1y1 is attached to the center of the manipulator to indicate the manipulator pose in the global system O \u2013 xy. Initially, the two systems coincide. The revolute clearance, e.g. the clearance of joint A1 as shown in Figure 1, is modeled by a vector denoted as rc from the center of the bearing to the center of the pin. If no penetration is considered, the vector rc yields k rc k r1 r2 (1) where operator k k computes the Euclidean norm of a vector, r1 and r2 are the radii of the bearing and the pin, respectively. Once a load W \u00bc \u00bdF;M > represented in the global system is exerted at the center point O1, the manipulator will be translated by \u00bdex; ey and rotated by h to a new pose indicated by Oe xeye. The deviation \u00bdex; ey; h is the error we concern" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002022_mmvip.2007.4430742-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002022_mmvip.2007.4430742-Figure2-1.png", + "caption": "Fig. 2. Geometry sketch map of calculating the output fiber luminous flux.", + "texts": [ + " According to the definition [4], [ ])arcsin(tantan NAT == \u03b8 , we know that (1) when )2/( Tad < , dTa 2> , the light power coupling into the output fibre is zero; (2) when )2/()2( Trad +> , the emanated light of the output fibre and the image of the input fibre is intersected in the light-cone bottom, the section area is a constant \u03c0r\u00b2,the end area of this light-cone is 2)2( rdT +\u03c0 , therefore, the gap coefficient in this scope is [ ]2)2/(' rdTr +=\u03b7 ; (3) When a/2T \u2264 d \u2264 (a+2r) /2T, the luminous flux coupling to output fiber is determined by the two parts overlapping area, which is the bottom part of the light-cone formed by the image of input fibre emanated light and the output fiber, the overlapping part as shown in Fig. 2. The overlapping area can be accurately calculated by using the gamma function, or using the linear approximation method, which the intersecting edge of the light-cone end area and the output fiber end surface is approximated by using beeline. If the \u03b4 is the distance between the light-cone edge and the output fiber, in this similar premise, the percentage of the light-cone irradiated surface of the output fiber end face can be given by the simple geometric analysis as. )]}1(sin[arccos)1()1({1 rrr arc \u03b4\u03b4\u03b4 \u03c0 \u03b2 \u2212\u2212\u2212\u2212= (1) This formula is a very useful function, whose curve is shown in Fig.3. The value of r \u03b4 can be calculated by the geometric relationship in Fig. 2 as follow. r dT r \u03b1\u03b4 \u2212= 2 (2) Supposed that the reflector is dark absorption, the optical power coupling efficiency of the two fibres is the ratio of the overlapping light spot area and the light-cone area. The input optical power percentage received by output fibre is the following as. 20 ) 2 (' rdT r P P i + === \u03b2\u03b2\u03b7\u03b7 (3) where \u03b7 is the coupling efficiency. The curve line of the step type fiber coupling efficiency \u03b7 and reflector location d can be available by the above relationships, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002208_012043-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002208_012043-Figure5-1.png", + "caption": "Figure 5. Shows diagrams and photos of drive lubricators", + "texts": [], + "surrounding_texts": [ + "Issues related to the wear and tear of the wheel sets flanges have always been and do remain relevant. Reducing the wear rate of the working surfaces of the wheels flanges to the level and below the level of the wear rate of the surfaces of the rolling circles makes it possible to reduce the number of wheel turning due to the \"thin\" flange and, accordingly, to increase the number of turns on the \"rolling\" of the tread . The latest technological scheme helps to reduce turning costs by 3\u20134 times and increase the resource of wheel sets by 4\u20135 times. Traditionally, the solution to problems associated with reducing the wear rate of the wheel sets flanges was carried out by supplying consistent antifriction lubricants to the contact zone of the wheel flange with the rail. However, the positive effect of lubrication of the wheel flanges is accompanied by negative consequences, such as oiling of the rolling stock and elements of the permanent way, getting of the rolling circle of the locomotive wheels into the contact zone and, as a consequence, the appearance of two-sided \"flat wheels\". As a result of complex theoretical experimental studies, a technological scheme was developed to protect the working surfaces of the wheel flanges from wear by applying a protective metal film on them . This solution is based on a slight excess of the sliding speed of the working surface of the flange relative to the rail (no more than 0.2%) (Figure 3, point 2) compared to the sliding speed of the rolling surface of the tread. Dynamics of Technical Systems (DTS 2020) IOP Conf. Series: Materials Science and Engineering 1029 (2021) 012043 IOP Publishing doi:10.1088/1757-899X/1029/1/012043 Thus, in the two-point contact, an additional frictional connection is created, which helps to increase the tractive effort of the locomotive by 20\u201325% when driving in a traction mode (Figure 4). Currently, there are two schemes for lubricating locomotive wheel set flanges with solid anti-friction materials: driven and non-driven. Dynamics of Technical Systems (DTS 2020) IOP Conf. Series: Materials Science and Engineering 1029 (2021) 012043 IOP Publishing doi:10.1088/1757-899X/1029/1/012043 In some cases, for example, with limited geometrical spaces for the installation of flange lubrication systems (FLS) or grease rail lubrication systems (GRLS) or the need to provide a long service life with a single refueling, non-powered structures are the best option to protect against wear of flanges. The scheme for calculating the values arising in the contact zone of lubricating rods (briquettes) or metal-clad rods is presented below." + ] + }, + { + "image_filename": "designv11_83_0002945_978-3-030-69178-3_15-Figure12.9-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002945_978-3-030-69178-3_15-Figure12.9-1.png", + "caption": "Fig. 12.9 Parameterisation scheme (process macro) for the robotic screwing process [53]", + "texts": [ + " As this setup is intended to be integrated into an HRC workcell including a robot system, a hand\u2013eye calibration problem needs to be solved as elaborated in [53]. Additionally, [53] describes an approach to perform an inertia compensation of the externally unloaded instrumented tool. This is required to sense the process forces only. Using the Newton\u2013Euler equations, the forces, torques and wrench of the tool centre point (TCP) can be calculated. Depending on the application, the retrieved process parameters of the instrumented tool need to be applied to a robotic skill definition, based on a process macro. Considering a screwing process, Fig. 12.9 depicts the scheme of parameterisation necessary to replicate the process using a robot system, where S and E describe the start and end positions of the screwing process, O the pre-position offset. F and T, describe the screwing force and final torque, respectively. Considering a system integration with the skill-based XRob robot programming framework (Sect. 12.3.1), 300 S. C. Akkaladevi et al. the retrieved process parameters (instrumented tool) need to be provided, e.g., via a REST-API to the programming system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000021_acc.2005.1470116-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000021_acc.2005.1470116-Figure1-1.png", + "caption": "Fig. 1. Gyroscope structure: polarization coils (1 to 8) excitation coils (1,3,5,7), rebalancing coils (2,4,6,8)", + "texts": [ + " The gyroscope consists of an elastic body such that one of its resonant modes is excited to constant amplitude vibrations (drive mode). Inducing rate about a particular body-fixed axis excites a different resonant mode into vibration (output mode). Generally, the rate by which the energy transfers from the first to the second mode is a measure of the induced rate. In an open-loop mode, the amplitude of the second mode is a measure to the rate, while in a closed-loop mode the rebalancing force allows the measure. For this application, the vibrating gyroscope employs inductance sensing for the second mode vibrations. Fig. 1 shows the structure of the gyroscope. It is composed of an elastic hollow cylinder, and a set of coils distributed around the cylinder. A permanent magnetic field is induced by a d.c. current exciting the polarization coils. By applying an alternating current to the excitation coils, a first mechanical mode of the elastic cylinder is excited into vibrations Fig. 2. The oscillation of the cylinder is an elliptical motion, that vibrates at the frequency of the magnetic field. If the cylinder is subjected to revolutions around z-axis with an angular rate \u2126(t), the elliptical motion is driven with the cylinder, with an opposite and lower speed \u03d5(t) = \u2212kc 2 \u2126(t), where kc depends on the gyroscope geometry (see Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002761_5.0049972-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002761_5.0049972-Figure1-1.png", + "caption": "FIGURE 1. The robot position variation from global to local frame.", + "texts": [ + " 2341, 020028-1\u2013020028-8; https://doi.org/10.1063/5.0049972 Published by AIP Publishing. 978-0-7354-4095-1/$30.00 020028-1 automation. It can be observed that, there lies a research gap in the literature regarding design of a robust navigational controller that can adapt to environmental scenarios and generate a collision-free smooth trajectory for effective navigation of a robotic agent. Here, a SNN based navigational strategy has been designed for optimized path fusion of a mobile robot in an environment. Figure 1 depicts the motion of the mobile robot in a global platform setting [10]. The velocity of the robot towards the target is represented as velocity (v). The local position of the robot (x, y) is changed to global position (Xnew; Ynew) using transformation from local angle (\u03b8) to global angle (\u03b8 ). During this transformation, the angular velocity (Av) also experiences a change from time \u2018t\u2019 to \u2018t+\u03b4t\u2019. The velocity of robot\u2019s left and right wheels can be expressed using equation 1, as: ( ) ( ) 2 2 l v r v l l v A R or v A R= \u2212 = + (1) In this equation, \u201cAv\u201d represents the angular velocity of the wheels and \u2018R\u2019 represents the radius measured from the Instantaneous Centre (ICC) to the local centre point of the robot as depicted in Figure 1. The values of \u2018R\u2019 and \u201cAv\u201d are represented as per the following equations 2 and 3. ( ) ( ) 2 l r r l l R v v v v= + \u2212 (2) If the wheels of the robot are rotated as per \u2018ICC\u2019, then the time interval for the local reference frame changes from \u2018t\u2019 to \u2018\u03b4t\u2019. Hence, local pose of the robot is updated on the global pose with an update in the position. In a similar fashion, angle (\u03b8) is updated to ( \u03b8 ) as per the following equation 4; vA t = + (4) \u201cICC\u201d can be calculated as per the following equation 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003016_j.jfranklin.2021.06.001-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003016_j.jfranklin.2021.06.001-Figure1-1.png", + "caption": "Fig. 1. The relevant coordinate frames.", + "texts": [ + " The relevant coordinate frames The definitions of relevant coordinate frames are given as follows: 1) The ECI coordinate frame F I \u2014O \u2212 X I Y I Z I : The definition is well known (For details, please see the work of [24] ). 2) The LOS coordinate frame F ls \u2014S \u2212 x ls y ls z ls : Define the center of mass of the service spacecraft as its origin, the axis x ls is directed from the origin to target, the axis y ls is on the X I \u2212 Y I plane and normal to the axis x ls , the axis z ls completes the right-handed coordinate system, and the angle \u03b5 between the axes x ls and X I should satisfy the condition Eq. (2) . As shown in Fig. 1 , R sr and R ta are the position vectors from the Earth to the service pacecraft and target respectively, r = R ta \u2212 R sr represents the relative position vector from he service spacecraft to target, and the axis x \u0301is the projection of x ls onto the X I \u2212 Y I plane. bviously, the relative position vector r can be expressed by a set of spherical coordinates = [ q i ] 3 = [ \u03c1 \u03b2 \u03b5 ] T , with \u03c1 = \u2016 r \u2016 being the relative distance, \u03b2 being the azimuthal ngle measured from the axis X I to x \u0301, and \u03b5 being the inclination angle measured from the xis x ls to x \u0301" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002731_s40799-021-00471-3-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002731_s40799-021-00471-3-Figure3-1.png", + "caption": "Fig. 3 The von Mises stress map", + "texts": [ + " The return to zero is due to the elastic stresses accumulated by the test piece. The universal test machine should be set so as to keep the variation in force fixed, regardless of the degree of degradation of the material in the specimen. A finite element analysis of the cam used in the fatigue and torsion test device adaptable to the universal test machine was performed. The cam is used to transform the translational movement of the central shaft into a translational and rotational movement. Figure 3 shows the model used and the von Mises stress map. The end sections are totally fixed, and in the area where the radial bearing makes contact, the cam is loaded with a surface pressure of 200 N/mm2. Each surface with an area of 90 mm2 will result in a total pressing force of 18 kN. The pressure was applied simultaneously on both surfaces of the cam channel, although, in reality, the two surfaces on which a radial bearing makes contact are loaded successively. One surface of the cam will be loaded when the force rises from 0 to +F, and the other surface will be loaded on return from +F to 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001450_aim.2008.4601693-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001450_aim.2008.4601693-Figure2-1.png", + "caption": "Fig. 2. Twin screw mechanism", + "texts": [ + " This method can avoid an obstacle without contacting by receiving virtual force from virtual spring and damper installed in virtual surface region. Finally, the mobility of the legged robot by non-contacting impedance control is presented through some experiments. In this paper, a legged mobile robot which has the ability to carry the human is developed. The photograph of prototype model is shown in Fig.1. This robot has three legs. Each leg consists of three linear actuators, so that total degree of freedom is 9. As shown in Fig.1 and Fig.2, to maintain the rigidity, three actuators are assigned to triangular configuration and farther the part of so-called shin has twin screw axle mechanism that are synchronously actuated by single DC-motor. Main specification of proposed legged robot is presented in Table I. 978-1-4244-2495-5/08/$25.00 \u00a9 2008 IEEE. 395 B. Inverse Kinematics A local servo control in terms of linear actuator length is generally convenient to implement whole robot control algorithm. Then it is need to formulate the relationship between each link length and a specific position at free leg and/or at the seat" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000088_j.jsv.2006.01.044-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000088_j.jsv.2006.01.044-Figure3-1.png", + "caption": "Fig. 3. The crack in crankpin-web fillet region.", + "texts": [ + " In this section, we will build a cracked crankshaft model based on the success of previous sections. Under the firing condition, crankshaft is endured an alternative impulsion load. The magnitude of stress in the crankpin-web fillet region will change periodically with a great extent, and then the damage with fatigue crack occurs frequently in this region. Generally, the crack occurs between the main journal and the crankpin, and it is transverse, 451 to the cross-section and at the surface of crank web [16] (see Fig. 3). For the crack in the crankpin-web fillet region extends from the surface into the crank web and does not extend to other part of the crankshaft. It is proposed to represent the cracked crank web by a 2-node slant crack beam element and simplify the other parts of the crankshaft as the method described in Section 3. Comparing with the model shown in Fig. 2, the cracked model consists of the same number of elements, and the number of the nodes will be cut down according to the number of the crack beam elements representing cracked crank web" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000739_j.precisioneng.2007.01.001-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000739_j.precisioneng.2007.01.001-Figure3-1.png", + "caption": "Fig. 3. Spring creep instrument: (1) load mass; (2) torque arm; (3) air b", + "texts": [ + " A version of Edlen\u2019s equation [7] was sed to correct the refractive index for thermal and barometric isturbances in the lab: \u2212 1 = 2.879294 \u00d7 10\u22129(1 + 0.54 \u00d7 10\u22126(C \u2212 300))P 1 + 0.003671 \u00d7 T \u22120.42063 \u00d7 10\u22129 \u00d7 F (3) Partially differentiating the above equation with respect to oth temperature (T) and pressure (P) provides the following orrection factors for the index of refraction of air, \u22120.92 ppm/K nd 0.356 ppm/Torr, respectively. The data acquisition system, uilt in LabVIEW, measures and records these disturbances, and he rotational creep values are corrected in post-processing. Fig. 3 depicts a drawing of the mechanical portion of he instrument which consists of the frame, torque arm, air earing, and rear spring support. The torque arm is attached earing; (4) spring socket block and worm gear assembly; (5) bearing block. t ( s o t w b a g t t 3 p t t t t b i a c s a a t \u00b1 o e c s o t e h m \u2018 b F t a a 4 w l i i o p c c s o h d s l o the spring via a shaft running through the air bearing providing near frictionless rotational motion, while providing ufficient radial stiffness to applied loads)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002566_icm46511.2021.9385615-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002566_icm46511.2021.9385615-Figure3-1.png", + "caption": "Fig. 3. Two-degree-of-freedom parallel robot.", + "texts": [ + " Icmp was given by Icmp = 1 Ktn \u03c4\u0302 , (35) where \u03c4\u0302 \u2208 R n denotes the estimated torque disturbance derived by the disturbance observer as \u03c4\u0302 (t) = 2\u2212 gTs 2 + gTs \u03c4\u0302 (t\u2212 Ts) + gTs 2 + gTs ( \u2212KtnIm(t) + Jn\u03b8\u0308(t) ) + gTs 2 + gTs ( \u2212KtnIm(t\u2212 Ts) + Jn\u03b8\u0308(t\u2212 Ts) ) , (36) where Ts and g are the sampling time and cut-off frequency, respectively. The block diagram of the acceleration control system is shown in Fig. 2(b). To compare the proposed periodic/aperiodic hybrid position/impedance control with hybrid position/impedance control, we conducted two experiments using the robot shown in Fig. 3. Two direct-drive servo motor SGMC-02BDC41 from Yasukawa were used in the robot. We used an Advanced Robot Control System for Linux, and the source code was written in C++. The experimental parameters are shown in Table I. The hybrid position/impedance control was used in Experiment 1, and the proposed periodic/aperiodic hybrid position/impedance control was used in Experiment 2. Each hybrid control system was implemented in the task space with x-axis and y-axis. Finally, we compared the results of Experiment 1 and 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000491_bfb0109980-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000491_bfb0109980-Figure7-1.png", + "caption": "Figure 7. Body coordinate reference frames and degrees of freedom. As depicted in Figure 7, the turret has one degree of freedom, 0,, about the", + "texts": [ + " The goal of the control design is to increase the pointing and tracking performance of the gun by including friction compensation. The control will be tested at the Apache 30mm Chain Gun Test Bed ADAWS Lab, Picatinny Arsenal Figure 6. The test bed gun is driven by a direct drive electric motor that is simply modeled as an input torque. The friction in the motor is dominant, so this is a simple friction control problem, not sandwiched friction problem. 5.1 Dynamic Model of Apache Gun A four-degree of freedom (DOF) model of the apache gun system was developed. A schematic of the multibody-flex model appears in Figure 7. The model consists of a rigid turret with flexible forks that connect to a rigid gun which has a flexible barrel attached to it. A rigid blast suppressor is attached to the muzzle end of the barrel. A two channel bending actuator, developed by TSi to increase pointing accuracy, is mounted to the flexible barrel. The actuator can deliver two pairs of torques to produce muzzle angular deflection in both azimuth and elevation. The bending actuator will not be used for this study. The gun system model was developed using the Mathematica package ProPac [23]. Figure 7 shows the three bodies into which the gun system is broken for modeling. Each body has a local reference frame that is located at the inboard joint. The reference frames and their associated degrees of freedom are also shown in Figure 7. The model was developed for designing and testing slewing controls. In order to keep the model dimension to a minimum; motions not related to slewing or slewing disturbances are not modeled. For example, the elevation of the gun is assumed a fixed value, thus eliminating a potential degree of freedom. In addition, any bending of the forks such as that which might be caused by a firing disturbance is not modeled. azimuth axis. The flexibility of the forks is modeled by a torsional spring with one degree of freedom, 02 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001893_s11426-007-0045-5-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001893_s11426-007-0045-5-Figure1-1.png", + "caption": "Figure 1 The preparation process of {Hb/nano-Au\u2295}n/L-cys modified gold electrode.", + "texts": [ + " A precursor layer of negatively charged L-cys monolayer was adsorbed by immersing the gold electrodes into L-cys aqueous solution (0.02 mol\u00b7L\u22121, pH 5.5) for 2.0 h at room temperature. After being washed with water, the L-cys/gold electrodes were alternately immersed in an aqueous dispersion of nano-Au\u2295 for 6.0 h and an Hb solution (3 mg\u00b7mL\u22121, pH 8.0) for 1.0 h at 4\u2103. This cycle was repeated to obtain the {Hb/nano-Au\u2295}n LBL films with the desirable number of bilayers (n). The preparation process of {Hb/nano-Au\u2295}n/L-cys modified gold electrode is shown in Figure 1. 1.4.1 Electrochemical impedance spectroscopy. The impedance technique was used for detection of impedance change and thickness during the layer-by-layer assembling process. The electrochemical impedance measurements were performed in a solution of 5 mmol\u00b7L\u22121 Fe(CN)6 3\u2212/4\u2212 + 0.1 mol\u00b7L\u22121 KCl + 0.1 mol\u00b7L\u22121 PBS (pH 7.0) at 25\u2103. The frequency range was 0.05\uff0d 105 Hz at 220 mV versus SCE. 1.4.2 Cyclic voltammogram and chronoamperometry response. Cyclic voltanmmetric measurements were performed in a conventional electrochemical cell containing a three- electrode arrangement and the potential swept from \u22120" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002804_j.matpr.2021.04.568-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002804_j.matpr.2021.04.568-Figure8-1.png", + "caption": "Fig. 8. Fatigue analysis of E-Glass/Banana/Epoxy.", + "texts": [], + "surrounding_texts": [ + "CAD models are designed in CATIA which contains special tools for the generation of traditional surfaces which will then be transformed into solid models, using traditional and composite monoleaf spring materials. The measurements of a spring leaf of a TATA SUMO vehicle are used for the design of the mono-leaf spring. Table1 displays the mono leaf spring configuration parameters and the planned mono leaf spring CAD model, as shown in Fig. 1. Finite Element Analysis (FEA) is a computational method for deconstruction into very small elements of a complex structure. In the simulated world, ANSYS offers an affordable way to examine the success of goods or processes. Digital prototyping is called this method of product growth. Users will iterate different scenarios using simulated prototyping techniques to refine the software even before the production is launched. This allows the probability and expense of failed designs to be reduced. In this study, a model was developed that was imported into the ANSYS workstation and FEA analysis. Meshing is the mechanism where the entity is divided into very small pieces called components. Elements. It is sometimes called a piece by piece. The leaf spring model is here meshed with a 10 mm brick mesh part scale. The front end is restricted and the rear end only in Y and Z is restricted; translational movement is permitted in X direction. Conditions of loading include applying a force on the middle of the spring of the leaf upward on the base of the leaf spring. The loading range is between 1000 N and 5000 N. The meshed model and limit and conditions of loading of a leaf spring are seen in Figs. 2 and 3." + ] + }, + { + "image_filename": "designv11_83_0002945_978-3-030-69178-3_15-Figure4.9-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002945_978-3-030-69178-3_15-Figure4.9-1.png", + "caption": "Fig. 4.9 Visualisation on the server side; a defined zones (green boxes), triggered zone (yellow box); b arrows pointing from POI to the closest points", + "texts": [ + " In the decision-making process, based on the received status and presence of the robot in one of the configured zones, one of the presented action scenarios is triggered. In the presented case, zones are defined in proximity to assembly stations, where it could be unsafe to control the robot to move away from the operator. When the robot end-effector is within defined zones and a potential collision is detected, the selected scenario is to stop the robot (collision avoidance scenario 2 in Fig. 4.7). Configuration of zones used is presented in Table 4.3 and the zones themselves are visualised in Fig. 4.9. The collision avoidance algorithm implements the procedure described earlier, preceded by a selection of the shortest collision distance found among collision vectors received from distributed client nodes. The \u201cmoving away\u201d scenario is implemented as a special case of the \u201cmodifying path\u201d scenariowhere the destination target is configured as the most recent point on the original robot path. The differential inverse kinematics module converts corrections of the POI position obtained from the collision avoidance algorithm to corresponding corrections in joint space" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001151_oceans.2007.4449163-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001151_oceans.2007.4449163-Figure3-1.png", + "caption": "Fig. 3. n-link underwater robot model", + "texts": [ + " Since digital computers are utilized for robot controllers in practical situation, the control performance is affected by discritization error of controller and sensor signals. Experimental results show that the control performance of our proposed method is better than the computed torque method. II. MODELING The UVMS model used in this paper is shown in Fig. 2. It has a robot base (vehicle) and a 2-DOF manipulator which can move in a vertical plane. First, we derive kinematic and dynamic equations of n-DOF manipulator model shown in Fig. 3 to obtain the 2-DOF mathematical model. Next, the 2-DOF mathematical model is described from the n-DOF manipulator model. Symbols are defined as follows: n: number of joint X, inertial coordinate frame Xi: link i coordinate frame (i = 0, 1, 2, , n; link 0 means vehicle) 0-933957-35-1 \u00a92007 MTS Ia,, added inertia tensor of link i with respect to Zi xo: position and attitude vector of 70(= IVOT0T]T) xe: position and attitude vector of end-effector(= [pT, OT]TT) Xo linear and angular vector of o0(= [v T, wTIT) linear and angular vector of end-effector(= [V[T, w Y > Z): (1) The layer is parallel to the intermediate Y-axis, but inclined to the X- and Z-axis (Fig. 1a). (2) The layer is parallel to the X-axis, but inclined to the Y- and Z-axis (Fig. 1b). (3) The layer is parallel to the Z-axis, but inclined to the X- and Y-axis (Fig. 1c). (4) The layer is oblique to all of the three principal strain axes. If the competent layer is parallel to the intermediate Y-axis, and inclined at a small angle (<45\u25e6) to the Z-axis (Fig. 1a), asymmetric folds develop under bulk coaxial plane strain, because the layer rotates slower towards the X-axis than the incompetent matrix (Price, 1967; Anthony and Wickham, 1978; Frehner and Schmalholz, 2006; Zulauf et al., 2020a). The same holds for multilayers (Cobbold et al., 1971; Watkinson, 1976). At a higher degree of layer inclination (>45\u25e6, Fig. 1a) boudinage is expected, which is asymmetric at low finite strain (Abe and Urai, 2012; Komoro\u0301czi, 2014), but nearly symmetric at high finite strain (Zulauf et al., 2020a). If the competent layer is parallel to the X-axis, and inclined to the Zand Y-axis, respectively (Fig. 1b), the layer behaves like a corresponding passive plane and rotates with the same velocity like the matrix. During progressive coaxial plane strain, the layer crosses different strain fields (shortening, reduced shortening, elongation) resulting in extensionparallel folds. If the viscosity ratio is high, these structures are portrayed by coeval boudinage and folding, the latter with fold axes parallel to X (Zulauf et al., 2020b). There is a continuous transition from coeval boudinage and folding to pure boudinage if the initial layer inclination changes incrementally from the XZ-to the XY-plane (Fig. 1b). If the layer is parallel to the Z-axis, but inclined to the X- and Y-axis, coaxial plane strain should also lead to extension-parallel folds (i.e. coeval folds and boudins). However, if the initial attitude of the layer approaches the YZ-plane, both structures should switch into pure folds with hinge lines parallel to the Y-axis (Fig. 1c). Folds and boudins, which grow in such type of oblique layers under bulk coaxial plane strain, have yet not been investigated. In the present study, we focus on such a setting. We carried out analogue modelling to investigate the deformation behavior of a single competent layer, which is parallel to the Z-axis but oblique to the X- and Y-axis, respectively. The layer was embedded in a less viscous matrix, and from run to run it was initially rotated around the Z-axis. This rotation should be attended by a rotation of the fold axis at 90\u25e6, from the Y- to the X-axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000332_50003-4-Figure3.31-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000332_50003-4-Figure3.31-1.png", + "caption": "Fig. 3.31. Side force characteristics of the single row brush model up to large levels of spin. (Compare with Fig.3.28 where spin is small and atp = 0.33 sin),).", + "texts": [], + "surrounding_texts": [ + "THEORY OF STEADY-STATE SLIP FORCE AND MOMENT GENERATION 131\nIntegration over the contact length after addition of v~ and multiplication with the stiffness per unit length Cpy gives the side force and after first multiplying with x the aligning torque. We obtain the formulae:\nsgnfP{A 1 3 1 3 } F - c l(a2Xsl - 3 X s l ) - A2(a2Xs2 - 3Xs2) +\na\n+ Cp (Ysl +Xsltana) (Xsl -Xs2) - 2 t a n a (Xs~ (3.86)\n- _ sgnfp IA ( l a 4 e e 1 4) ( 1 4 2 2 1 4)} M - l c - a + a a + Z 2 P a ~ 1 2 Xsl 2Xs l - A 2 - Xs2 2Xs2 +\n{2(y l t a n a 3 3 } + Cp sl + Xs l tana) (Xs2~ - Xs22) - 3 (Xsl - xs2) (3.87)\nThe resulting characteristics have been presented in Figs.3.31 and 3.32. The graphs form an extension of the diagram of Fig.3.28 where the level of camber correspond to 'small' spin. It can be observed that in accordance with Fig.3.25 the force at zero side slip first increases with increasing spin and then decays. As was the case with smaller spin for the case where spin and side slip have the same sign, the slip angle where the peak side force is reached becomes larger. When the signs of both slip components have opposite signs, the level of side slip where the force saturates may become very large. As can be seen from Fig.3.30b the deflection pattern becomes more anti-symmetric when with positive spin the slip angle is negative. This explains the fact that at higher levels of spin the torque attains its maximum at larger slip angles with a sign opposite to that of the spin. The observation concerning the peak side force, of course, also holds for the slip angle where the torque reduces to zero.\nSpin , l o n g i t u d i n a l and side slip, the width effect\nThe width of the contact patch has a considerable effect on the torque and indirectly on the side force because of the consumption of some of the friction by the longitudinal forces involved. Furthermore, for the actual tyre with carcass compliance, the spin torque will generate an additional distortion of the carcass which results in a further change of the effective slip angle (beside the distortion already brought about by the aligning torque that results from lateral forces). Amongst other things, these matters can be taken into account in the tread simulation model to be dealt with in Section 3.3.\nIn the part that follows now, we will show the complexity involved when longitudinal slip is considered beside spin and side slip. To include the effect of the width of the contact patch we consider a model with a left and a right row of tread elements positioned at a distance Y L - -brow and Y R - brow from the wheel", + "132 THEORY OF STEADY-STATE SLIP FORCE AND MOMENT GENERATION\ncentre plane. In fact, we may assume that we deal with two wheels attached to each other on the same shaft at a distance 2bro w f r o m each other. The wheels are subjected to the same side slip and turn slip velocities, Vsy and ~, and show the same camber angle 7. However, the longitudinal slip velocities are different, for the case of camber because of a difference in effective rolling radii. We have for", + "THEORY OF STEADY-STATE SLIP FORCE AND MOMENT GENERATION 133\nthe longitudinal slip velocity of the left or right wheel positioned at a distance YL, R from the centre plane:\nVsxL,R - Vsx - YL,R { ~ - (1 - ey) s sin), } (3.88)\nThis expression is obtained by considering Eq.(2.55) in which conicity is disregarded, steady-state is assumed to occur and the camber reduction factor e~ is introduced. The factors 0 are defined as (like in (3.54, 3.55)):\nX O r x ( y ) - - e r , O ( x , y ) - - e r - sin), (3.89) ?-\ne\nFrom Eqs.(2.55, 2.56) using (3.88) the sliding velocity components are obtained\nOUL,R V x L ' R - VsxL 'R - OX V (3.90)\nOV V y - L y - - - w + x { ~ - ( 1 - ~)K2sin),} (3.91)\nOX r\nAfter introducing the theoretical slip quantities for the two attached wheels\nv V x L ' R tr - sy tr - - . (3 92) ~ - - ~ 9\ntTxL'R = W Y V gt V F t f\nwe find for the gradients of the deflections in the adhesion zone (where Vg = 0) if small spin is considered (sliding only at the rear):\nOblL,R _ OX - r (3.93)\nov 1 - - tr - x + (1 - ey) -~e sin (3.94) OX Y u/\nwhich yields after integration for the deflections in the adhesion zone (x < xt):\nuL, R - (a - x) axL R (3.95)\n1 sin),} (3.96) v - (a - x ) Cry + V2(a 2 - x 2 ) { o +(1 - e ) r e\nThe transition point from adhesion to sliding, at x= x,, can be assessed with the aid of the condition\nc eL, R - !1 q z (3.97)" + ] + }, + { + "image_filename": "designv11_83_0001507_fie.2007.4418155-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001507_fie.2007.4418155-Figure5-1.png", + "caption": "FIGURE 5 PROCESS LINE WITH MACHINE TOOL EMULATED ON A 3D ENVIRONMENT", + "texts": [ + "00 \u00a92007 IEEE October 10 \u2013 13, 2007, Milwaukee, WI 37th ASEE/IEEE Frontiers in Education Conference S3G-20 The real scale model of the Process Line with Machine tool and its corresponding Emulation are shown in Figures 3 and 4, respectively. At the moment, a diverse library of Third-Dimension (3D) Emulations is being developed. 3D Emulations offer, as it could be expected, more characteristics than the ones in Two Dimensions (2D). In a 3D visualization it is possible to get immersed into the process and actually adapt the view to observe specific tasks or circumstances occurring during the process. On a 3D environment the student can zoom, rotate, or scroll the process at will (see Figure 5). Taking advantage of the 3D visualization, students can observe the process from perspectives that most of the times are not available on real systems. Before the development of the Emulations, the students had to solve the exercises and the Automation projects at the laboratory, and only within the laboratory sessions. This caused that pretty often, students could not finish the exercises on time, since the automation tasks normally require a considerable amount of tests in order to get to a correct solution" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000430_ias.2005.1518684-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000430_ias.2005.1518684-Figure5-1.png", + "caption": "Figure 5. Circuit model of SynRM with equivalent stator iron loss resistance..", + "texts": [ + " The iqc is found as an inflection point of Ld \u2013 iq curve drawn for each id=3.0, 5.0 and 7.0 (A). Finally using the obtained Ld1(id) values (in our case Ld1(3.0), Ld1(5.0) and Ld1(7.0)), the three constants, B1, C1 and D1, can be found by a similar least-squares identification. The constants in (4) can be determined by the same procedure. The curves in Fig. 4 are calculated from (3) and (4). It is found that the calculated values agree well with the measured values. C. Measurement of equivalent iron loss resistance Fig. 5 shows a circuit model of SynRMs with an equivalent stator iron loss resistance Rm [5]. In Fig.5, the relation between currents and flux linkages is written as + = c b a m mc mb ma c b a P R i i i i i i \u03bb \u03bb \u03bb 1 (5) where = mc mb ma cbcca bcbab caaba c b a i i i LMM MLM MML \u03bb \u03bb \u03bb ( )reggsa LLlL \u03b82cos20 \u2212+= ( )3/22cos20 \u03c0\u03b8 +\u2212+= reggsb LLlL ( )3/22cos20 \u03c0\u03b8 \u2212\u2212+= reggsc LLlL ( )3/22cos2/ 20 \u03c0\u03b8 \u2212\u2212\u2212= reggab LLM ( )reggbc LLM \u03b82cos2/ 20 \u2212\u2212= ( )3/22cos2/ 20 \u03c0\u03b8 +\u2212\u2212= reggca LLM . Here, P is differential operator (=d/dt), Rm is equivalent iron loss resistance (\u2126) and ls is armature leakage inductance (H)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001861_micai.2008.76-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001861_micai.2008.76-Figure7-1.png", + "caption": "Figure 7. Flexible manipulator robot prototype with pneumatic actuator.", + "texts": [ + " The values of the valves represent a percentage of the aperture, where 0 means totally closed and 1 is totally open. The valve A2 is equal to the A1, because in practice it is the same valve, but the mathematical model needs the three values. The value of A3 depends of A1, we only need to control the air return of the system, to get a better behavior; this valve is important to stop the piston displacement when the arm is close to the set point. )()( )()( )()()( 133 12 101 kk kk kvkppk TAKTA TATA TVKTeKATA (9) The figure 7 shows the flexible manipulator prototype simulation with pneumatic actuator, where the arm is made of PVC material. Remember that on figure 1 we have a scheme that shows the air flow in both directions, and the air return is controlled by two independent proportional valves. Then, these valves are used to avoid that some of the cameras remains without pressure, generating a braking effect to the advance of the piston. Due to the high non-linearity of the system, the hand-tuning method is used to set the member ship intervals used in the fuzzy process for each input and output" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000834_robot.2007.363875-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000834_robot.2007.363875-Figure3-1.png", + "caption": "Fig. 3. (left) Robot\u2019s representation in Cartesian and Polar coordinate systems. (right) Kinematics characteristics of PIONEER 1 (differential drive robot).", + "texts": [ + " (7) Nevertheless, the use of this function could suppose taking a wrong action under some particular conditions. For example, it happens when the robot\u2019s orientation has a high discrepancy with the goal (\u03b1 > 90\u25e6); in this case the robot will move at a high speed away from the goal. On the other hand, the Speed function does not take into account the closeness to the goal, thus when the robot is near the goal this function will promote a wrong action: fast navigation. The Goal function measures the alignment of the robot orientation with respect to the goal, defined with the parameter (\u03b1) (Fig. 3(left)). It computes an orientation error, assuming that the robot moves with a constant velocity w during the interval of time of the control loop \u2206t. This function could be defined as (e.g., [4], [7], [8]): Goal (w) = 1\u2212 |\u03b1\u2212 w \u00b7\u2206t| /\u03c0. (8) Note that it does not include the angular closeness due to translational velocity; this drawback is emphasized in those cases where the robot is near the goal but with a wrong orientation (high \u03b1). The Dist function, presented in (1), represents the distance to the nearest obstacle over a circular trajectory with a curvature given by the velocities (v, w)", + " In addition, further improvements to the original DWA are introduced to avoid the inconveniences found in the terms of the objective function that impose a fast (7) and oriented to the goal (8) navigation. Next, kinematics equations used to model the robot\u2019s motion are introduced. Then, the proposed control law based on Lyapunov stability criteria is presented. Finally, a new objective function is given. Assuming the robot is represented by a point, its kinematics equations, in a Cartesian space, can be expressed as: x\u0307 = v \u00b7 cos (\u03b8) , y\u0307 = v \u00b7 sin (\u03b8) , (9) \u03b8\u0307 = w, where \u03b8 defines the robot\u2019s orientation according to a global coordinate system(Fig. 3(left)). These equations can be expressed in a polar coordinate system associated with the goal: \u03c1\u0307 = \u2212v \u00b7 cos (\u03b1) , \u03b1\u0307 = \u2212w + v \u00b7 sin (\u03b1) /\u03c1, (10) \u03b8\u0307 = \u2212\u03b1\u0307. Although the robot has been represented as a point, this model can be extended to different kinds of robots, for instance synchronous drive robots or differential drive (see Section IV). In these cases, a direct implementation could be to represent the robot rotation center as the reference together with its minimum robot\u2019s bounding circle. This section presents an ideal control law (vi, wi) that allows driving the robot to the goal guaranteeing a global convergence", + " High \u03bb1 and \u03bb2 values will result in a goal oriented behavior, while a high \u03bb3 value will result in a highly reactive behavior that favor the collision avoidance part. Actually, the tuning of these parameters is defined by characteristics such as: the environment\u2019s structure (number and distribution of obstacles), the required robot behavior (level of reaction), and the degree of knowledge of the environment. Experimental results obtained with both simulated and real environments are presented. In both cases a PIONEER 1 robot was used (see Fig. 3(right)). It presents a maximum translation velocity of 600 mm/s, a maximum rotation velocity of about 2.5 rad/s, and a distance between drive wheels of 325 mm. Before going into details about the obtained experimental results the required extension of IDWA to tackle differential drive robots is presented. This section presents an extension of the I-DWA algorithm in order to be able to tackle differential drive locomotion problems. Kinematics equations, together with the corresponding transformations in the velocity space, to handle this kind of robots are also introduced", + " Drive wheels are independently controlled while the passive rear wheel is only used as an additional leaning point to keep the robot\u2019s balance. The rear wheel is automatically oriented according to the robot motion. The robot\u2019s displacement is achieved by means of a separated control of each drive wheel. An instantaneous center of curvature (ICC), defined by the intersection of the drive wheels\u2019 axis with the passive wheel\u2019s axis, is automatically defined according to the robot\u2019s displacement (see Fig. 3(right)). Kinematics equations define the interaction between control commands and the corresponding space state. Thus, in a differential drive locomotion robot, these equations will reflect the robot\u2019s position (x, y, \u03b8) when the velocity of each drive wheels is controlled (vright, vleft). Therefore, from (9), translational and rotational velocities can be expressed by means of the drive wheel\u2019s velocities: v = (vright + vleft) /2, w = (vright \u2212 vleft) /L, (18) where L is the length of the drive wheels\u2019 axis (Fig. 3(right)). A clockwise displacement (w > 0) is performed when vright > vleft, otherwise a counter-clockwise displacement will be executed. From (18) the kinematics equations of a differential drive robot can be expressed as: x\u0307 = ((vright + vleft) /2) \u00b7 cos (\u03b8) , y\u0307 = ((vright + vleft) /2) \u00b7 sin (\u03b8) , (19) \u03b8\u0307 = ((vright \u2212 vleft) /L) . Equation (18) can be depicted by means of a matrix as: [ v w ] = [ 1/2 1/2 1/L \u22121/L ] \u00b7 [ vright vleft ] (20) hence, its inverse representation can be easily expressed as: [ vright vleft ] = [ 1 L/2 1 \u2212L/2 ] \u00b7 [ v w ] (21) The (20) and (21) equations allow the representation of the robot\u2019s velocities (assuming the robot is represented by a point) together with the drive wheels\u2019 velocities" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002838_012021-Figure14-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002838_012021-Figure14-1.png", + "caption": "Figure 14. Strain Distribution in Kevlar 29", + "texts": [ + " The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.3.2. Stress Distribution The Max. and Min. Stress Distribution in Kevlar 29 is 36.704 MPa and 0.26231 MPa respectively shown in Figure 13. 3.3.3. Strain Distribution The Max. and Min. Strain distribution in Kevlar 29 is 0.00054321 and 0.000005542 respectively shown in Figure 14. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.4. Analysing Testing Result of S-Glass 3.4.1. Total Deformation The Max. and Min. Total Deformation in S-Glass is 0.18974 mm and 0 mm respectively shown in Figure 15. The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. 3.4.2. Stress distribution The Max. And Min. Stress distribution in S-Glass is 59.887 MPa and 0.3646 MPa respectively shown in Figure 16" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002020_acemp.2007.4510512-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002020_acemp.2007.4510512-Figure1-1.png", + "caption": "Fig. 1 Analyzed results by various methods.", + "texts": [ + "his paper presents the dynamic modelling of the vector magnetic hysteretic property, which is called \u201cDynamic E&S Model\u201d. This paper deals with the dynamic vector magnetic hysteretic engineering model, \u201cDynamic E&S Model\u201d, which can analyse the effect of eddy current under vector magnetic behaviour. We will report the dynamic magnetic characteristic analysis taken care account of eddy current. 1 Introduction Figure 1 shows the distribution of magnetic flux density of the ring core obtained by commercial soft program and static E&S model [1][2]. This core is the non oriented electrical steel sheet, which is easy to be magnetized in the rolling direction. The case-1 and case-2 are used commercial soft. Case-1 is isotropic analysis obtained by using one magnetizing curve. Case-2 is used two kinds of curve, which are the rolling direction and the transverse direction. Case-3 is by E&S model. The results of case-1 and case-2 are not realistic" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001645_sice.2008.4655086-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001645_sice.2008.4655086-Figure2-1.png", + "caption": "Fig. 2 Sagittal plane model of a quadruped robot: Rush.", + "texts": [ + " The body consists of a platform that carries actuators, transmission devices and computer interface electronics. The total weight of the robot is 4.3 kg. The length and width of the body are 30 cm and 20 cm, respectively. The height of a leg is 20 cm when the robot stands. A DC motor of 7.5 (W) power with 19 reduction ratio is equipped at hip joint. Knee joint is passive. At this moment, Rush is not power autonomous. Detailed values of physical parameters are listed in Table 1. The sagittal plane model of Rush is shown in Fig.2. This model is composed of a rigid body and a pair of spring-loaded two segment legs. The suffix f and h represent the fore legs and hind legs, respectively. The angle of the hip joint with respect to the toe is \u03b3leg (i.e. hip and knee) and a contact sensor at the toe. It has active hip joint and passive knee joint with spring on each leg. spring for energy restoration hip joint When the four legs stance phase appears 2 , the phase transitions in the bound gait is illustrated as Fig.3. This bound gait contains four phases (flight, forelegs stance, four legs stance and hindlegs stance)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002979_iemdc47953.2021.9449609-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002979_iemdc47953.2021.9449609-Figure1-1.png", + "caption": "Fig. 1 Initial designs of (a) modular and (b) non-modular SPM machines.", + "texts": [ + " However, in the existing literature about modular machines, their optimization did not consider the flux gap width as variable. Instead, the flux gaps are often introduced after the non-modular PM machines are optimized. This cannot fully reflect the significant impact of flux gaps on machine electromagnetic performance and also on other machine parameters. To address this shortcoming in the literature, the flux gap width together with other parameters such as the stator yoke height, stator tooth width, split ratio (rotor outer diameter over stator outer diameter), and rotor yoke thickness (as shown in Fig. 1) will be used as variables in the global optimization process in this paper. The average torque is the primary objective and the torque ripple, copper loss, and total machine mass are considered as secondary objectives. At last, to show the benefits of modular machines, the optimized modular machine will be compared with its optimized non-modular counterpart. This comparison is carried out for machines with different slot/pole number combinations in order to achieve more general conclusions. II. INITIAL SPECIFICATION OF MODULAR MACHINE The initial specifications of the modular and non-modular machines used in aerospace application are given in Table I and the cross-sections are shown in Fig. 1. By introducing the flux gaps into the stator iron core, the effective airgap length is significantly increased due to the high reluctance of the flux gaps, which will in turn affect the airgap flux density [2]. According to [5], a flux focusing/defocusing factor, which relates to the slot/pole number combination, is introduced to take the flux focusing or defocusing effect into account. For example, for machines with slot number lower than pole number, e.g. 24-slot/28-pole, the flux gaps will help to increase the open-circuit airgap flux density" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002249_s10015-021-00678-y-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002249_s10015-021-00678-y-Figure1-1.png", + "caption": "Fig. 1 Insect-type microrobot", + "texts": [ + " The microrobot can move its six legs independently, and its mechanical structure allows for several gaits by changing the driving waveform. Thus, we have developed neural networks that have a switching mechanism for the gait pattern, which we have discussed using the simulation results [7]. The NNIC can generate the tripod gait and ripple gait patterns typical of insects. The former is used for fast walking, whereas the latter is used for slow walking. In this study, we develop an NNIC that has a switching mechanism and discuss the differences in the generated gait pattern between the measurement and simulation results. Figure\u00a01 presents the insect-type microrobot. It has a size of 4.6 \u00d7 9.0 \u00d7 6.4\u00a0mm (width, length, and height, respectively). Its leg and body parts were created using the microfabrication technology. The microrobot consisted of leg and body parts, artificial muscle wires, and a circuit board. The circuit board comprised an NNIC (2.5 mm2) and capacitors. The link mechanism converted the expansion and shrinking of the artificial muscle wires in the microrobot\u2019s gait pattern. The microrobot could move its legs independently; thus, various gait patterns were realized by inputting the driving waveform", + " The excitatory\u2013inhibitory switchable synaptic model could switch the connection by VW. This chapter describes the design, simulation results, and measurement results of the NNIC, which could switch between a tripod gait pattern and ripple gait pattern. 1 3 Figure\u00a06 presents the connection diagram of the NNIC, whereas Fig.\u00a07 shows half of the circuit diagram of the NNIC. The NNIC consisted of 10 inhibitory synaptic models, six cell body models, and two excitatory\u2013inhibitory switchable synaptic models. In Fig.\u00a01, the cell body models of C1, C2, C3, C4, C5, and C6 output pulses to Leg 1, Leg 2, Leg 3, Leg 4, Leg 5, and Leg 6, respectively, of the microrobot. The NNIC could change the gait pattern by changing the coupling coefficient voltage VW of the excitatory\u2013inhibitory switchable synaptic model. Figure\u00a08 presents an example of the simulation results of the output waveform of the tripod gait pattern in the case of VW = 3.0\u00a0 V. Moreover, Fig.\u00a0 9 shows an example of the simulation results of the output waveform of the ripple gait pattern in the case of VW = \u2212 3", + " In the future, we will mount the NNIC chip on a microrobot and conduct a gait experiment. Acknowledgements This work was supported by JSPS KAKENHI Grant Number JP18K04060. Also, the part of this research supported by the Research Institute of Science and Technology Nihon University College of Science and Technology Leading Research Promotion Grant. The fabrication of the hexapod-type microrobot was supported by the Research Center for Micro Functional Devices, Nihon University. The VLSI chip of Fig.\u00a01 has been fabricated by Digian Technology, Inc. This work is supported by VLSI Design and Education Center (VDEC), the University of Tokyo in collaboration with Synopsys, Inc., Cadence Design Systems, Inc. and Mentor Graphics, Inc. The VLSI chip in this study has been fabricated in the chip fabrication program of VLSI Design and Education Center (VDEC), the University of Tokyo in collaboration with On-Semiconductor Niigata, and Toppan Printing Corporation. Also, we appreciate the Nihon University Robotics Society" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000817_1.2712953-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000817_1.2712953-Figure1-1.png", + "caption": "FIG. 1. Cross-sectional views of mechanical design of the magnets before and after viscous deformation. The symbols H and M denote the specific aligned angles to the tangential direction and specific anisotropy angles to the tangential direction.", + "texts": [ + " It was clarified that in the arc-shaped magnets, the anisotropy directions were controlled continuously by using an alignment field held in a constant direction together with a mechanical design of the preformed magnets before viscous deformation. Resultantly, characteristics of a motor comprising of a rotor with the arc-shaped magnets were superior compared with those previously reported SPMSM. The cross-sectional design of the preformed magnet with continuously controlled anisotropy directions by using an alignment field in the constant direction was executed by the nonlinear structural analysis, as shown in Fig. 1. Although we estimated that H was almost equal to the anisotropy direction M , as shown in Figs. 1 a and 1 b , the tangential directions of the alignment direction for specific angle H were 0.2845 mm in inner radius and 0.3655 mm in outer radius. The intended dimensions of the magnets using viscous deformation were 18.95 mm in inner radius, 20.45 mm in outer radius, and 14.5 mm in length, respectively. In the above-mentioned magnet, the relation between H and mechanical angle was designed by H =5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002241_j.actaastro.2021.02.001-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002241_j.actaastro.2021.02.001-Figure6-1.png", + "caption": "Fig. 6. (a) \u2018\u2018Signal Flag\u2019\u2019 as described in Section 5.1 and by Kulwicki et al. [1]. (b) \u2018\u2018Continuous Signal Flag\u2019\u2019 as described in Section 5.1.2.", + "texts": [ + " A maneuver referred to as \u2018\u2018Signal Flag\u2019\u2019 (X.1) is described by Kulwicki et al. [1] which can result in rotation around the AP axis. This maneuver was also investigated by Stirling et al. [14], in a study of how quickly the maneuver can be learned by fully and minimally trained subjects in an Earth-based analog environment. Both previous papers offer step by step descriptions of the maneuver. In the original description, the hands are always held 180\u25e6 out of phase with each other throughout the motion, as shown in Fig. 6(a). Using the terminology from 2.2, this maneuver is an example of a \u2018\u2018cartwheel\u2019\u2019. This maneuver successfully initiates incremental rotation around the AP axis. Significant rotation occurs while moving the arms in the extended position. However, some counter-rotation occurs when resetting to the initial position. Artists are not always wishing to engage in incremental rotation however. Two refinements which I will discuss are: (1) selecting an angle for the legs to ensure a more stable rotation and (2) an adjustment to the sequence to allow for continuous rather than discrete rotation", + " The \u2018\u2018continuous signal flag\u2019\u2019 maneuver allows for continuous rotation which comes from changing the phase relationship between the arms. This new version of the maneuver starts with the arms above the head, they are brought down to the sides by moving one arm through extension in , while the other arms moves in flexion along the head\u2013 tail axis. The arms are then brought back up with the opposite arm moving in extension and flexion. See Figs. 9 for a sequence of photos demonstrating the maneuver. This sequence is compared to the \u2018\u2018signal flag\u2019\u2019 in Fig. 6(b). During the maneuver, both arms should be kept confined to as much as possible. Keep the elbow of the \u2018\u2018sliding arm\u2019\u2019 as close to the head\u2013tail axis as possible. This \u2018\u2018continuous signal flag\u2019\u2019 could be thought of as \u2018\u2018passing off\u2019\u2019 the angular momentum from one arm to the other arm. When the arms are raised fully over the head and the thighs are brought to 30\u25e6 forward of the coronal plane, \ud835\udee5\ud835\udc29\ud835\udc50\ud835\udc5c\ud835\udc5a = (1.9 cm)?\u0302? + (14.0 cm)?\u0302?. We find \ud835\udee5\ud835\udc29\ud835\udc50\ud835\udc5c\ud835\udc5a = (2.5 cm)?\u0302?+(7.4 cm)?\u0302? while the arms execute the raising and lower operations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002953_tia.2021.3089662-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002953_tia.2021.3089662-Figure1-1.png", + "caption": "Fig. 1: Cross section of the induction motor under test.", + "texts": [ + " The paper is organized as follows: 1) section II reports the relevant data of the IM under study, the test bench setup and the measurement procedures; 2) section III describes the equations adopted for RFO during the post-process of the measurement; 3) section IV briefly presents the simulation strategy based on RFO FE analysis; 4) section V compares the experimental and simulation results of currents and flux-linkages; 5) section VI shows how to get the estimated torque and rotor resistance throughout the quantities derived in section V. A. IM Under Consideration Fig. 1 shows the IM geometry and Table I reports its geometrics, winding and nominal electrical data. The IM here considered has a skewed rotor with closed slots. The idea is to supply the machine with a given stator current and frequency. The rotor speed is varied controlling the master motor speed. In this way, from the synchronous speed, the rotor is gradually slowed down, to the standstill condition. Inside the motor, the stator currents are differently split into d- and q-axis components, at the different imposed rotor frequencies" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003407_j.optlastec.2021.107425-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003407_j.optlastec.2021.107425-Figure3-1.png", + "caption": "Fig. 3. 3D geometric model and finite element mesh model of WSPC powder (a) 3D geometric model (b) Finite element mesh model.", + "texts": [ + " In the SLS process, because of the effect of laser heat source, the energy absorbed by WSPC powder in X, Y and Z directions is different, resulting in the formed sintering area cannot be described by twodimensional model. Therefore, a three-dimensional finite element geometric model should be established and simplified according to the actual forming process of SLS. The processed model ensure the accuracy, also improve the computational efficiency of the model. In this paper, the model of WSPC powder was simplified considering the calculation accuracy and efficiency. The size of the simplified model was 3.2 mm \u00d7 1.2 mm \u00d7 0.3 mm and its structure was demonstrated in Fig. 3(a). To ensure that the numerical simulation analysis carry out smoothly, the following assumptions should be made for the model: (1) WSPC powder is regarded as isotropic and continuous homogeneous medium, the powder is spherical and does not deform when approached; (2) In the laser sintering process, the thermal convection coefficient and thermal radiation coefficient between the surface of WSPC powder bed and the surrounding environment are both fixed constants, which do not change with temperature; (3) The influence of liquid binder in the sintering pool on the temperature field is ignored; (4) The vaporization of the liquid binder in the sintering pool is not considered", + " In the numerical simulation process of SLS forming, the used finite element model is a cuboid with relatively simple structure. Therefore, the requirements of finite element simulation technology can be met without the use of mesh local refinement technology when the model grid is divided. In summary, eight-node linear hexahedral elements were selected in this paper to partition the mesh of the WSPC powder 3D model. The mesh density of the laser heat source area and other areas was 0.05 mm. The 3D finite element mesh generation model was shown in Fig. 3(b). In the process of SLS numerical simulation, the powder materials melted even vaporized by instantaneous high temperature produced by the laser beam, thermophysical parameters of materials are highly nonlinear, so the specific heat capacity and thermal conductivity of material are particularly important. WSPC powder consists of the walnut shell powder and Co-PES powder (mass ratio of 1:4) [31], which is used for thermal physical property parameters test in order to obtain thermophysical parameters such as thermal conductivity, specific heat capacity and density, as shown in Table 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002135_s00202-020-01163-8-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002135_s00202-020-01163-8-Figure6-1.png", + "caption": "Fig. 6 (a) The magnetic model of the maximum flux density of the proposed TSCMG and (b) mesh model (Maximum length of elements 1\u00a0mm, Maximum number of elements 5000, skin depth 1\u00a0mm, surface triangle length 1\u00a0mm, Maximum number of surface elements 5000) and (c) magnetic gear reluctance", + "texts": [ + " Yet, considering the manufacturing limitation and thermal resistance of the material, it can be said that glass due to its fragile and polyethylene due to its deformed shape under increased temperature cannot be proper candidates for TSCMGs. Consequently, after analyzing and modeling in the FEM, the best candidate in this paper is aluminum. Therefore, the ultimate and optimal material of the TSCGM is shown in Fig.\u00a05. Also, Figs.\u00a06 and 7 depict the magnetic model of the maximum flux density with the aluminum middle rotor and its distributed flux density, respectively. Referring to Fig.\u00a06, the maximum flux density of the TSCMG with a middle aluminum rotor is 2\u00a0T which is suitable. Also, Fig.\u00a07 illustrates that the flux density in the TSCGM with a middle aluminum rotor is very desired and appropriate for distribution. A multi-speed gear with a conventional structure is presented in Fig.\u00a08(b). After conducting thorough studies, it has been clarified that this type of gear with such a structure in which iron is positioned in a dual-layer configuration in the middle rotor provides low efficiency because of the inappropriate design of the middle rotor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002276_s12206-021-0114-2-Figure10-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002276_s12206-021-0114-2-Figure10-1.png", + "caption": "Fig. 10. Thread geometry.", + "texts": [ + " 7, the box thread is magnified to illustrate the detailed geometric parameters of the thread. The thread thickness is 4.72 mm with a space length of 7.58 mm. The pitch is 15.25 mm and the thread length is 7.04 mm. Fig. 8 describes the parameters related to the pin. The outside diameter and thickness of the pin is 139.7 mm and 10.54 mm, respectively. Fig. 9 shows a magnified image of the pin thread and summarizes the thread geometry parameters. The space length, pitch, and thread length are 7.25 mm, 15.25 mm, and 7.97 mm, respectively. Fig. 10 describes the parameters related to the thread. The upper (stabbing flank) and lower corner radius (load flank) are both 0.2 mm. Once the 3-D model for analyzing the premium connection system was prepared, the FE mesh model was defined using five parameters: global size, local size, load flank size, stabbing flank size, and number of flank elements. The quad element was primarily used for 2-D analysis. To reduce the analysis time, the mesh model was created with a coarse mesh by selecting a relatively large mesh size" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003228_j.taml.2021.100281-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003228_j.taml.2021.100281-Figure1-1.png", + "caption": "Fig. 1. a Kinematic chain of the mechanism with two degrees of freedom, b the plan of linear velocities of the chain and c link 5 with applied forces.", + "texts": [ + " Ivanov Theoretical and Applied Mechanics Letters xxx (xxxx) xxx w f a t c t a 5 b b l l a f V f m a b w c i t o C i t ( t M t f l t B l a d 0 The article presents the basics of the theory of mechanisms ith two degrees of freedom with one input. The kinematic and orce analysis of two-mobile adaptive mechanisms is developed nd the principle of motion definability is described which allows o implement an additional constraint with a sufficiently high effiiency. The kinematic chain of a fundamentally new mechanism with wo degrees of freedom and one input ( Fig. 1 a) contains a rack 0, n input carrier H 1 , a movable closed contour with gears 1-2-3-6- -4 and an output carrier H 2 . A closed contour having zero moility contains a satellite 2, a solar wheel block 1-4, a ring wheel lock 3-6, and a satellite 5. The carriers H 1 and H 2 are the initial inks of the mechanism. The initial links can have set angular veocity \u03c9 H1 and \u03c9 H2 and the corresponding moments of forces M H1 nd M H2 . The moment M H1 = const , the moment M H2 can take diferent values within a given range. Linear kinematic and force parameters: B = \u03c9 H1 r H1 , V K = \u03c9 H2 r H2 , F H1 = M H1 / r H1 , F H2 = M H2 / r H2 . Here r H1 , r H2 are the radii of the points B, K of application of orces. The presented two-mobile scheme of the mechanism with a ovable closed contour allows you to build a plan of linear speeds ccording to the set speeds of the two initial links ( Fig. 1 b) and ring the acting forces to the intermediate link 5 ( Fig. 1 c). Presented in Fig. 1 a the scheme of a two-mobile mechanism ith a movable closed contour allows you to create equilibrium onditions. (In a mechanism with two degrees of freedom, it is mpossible to create equilibrium conditions without a closed conour). The equilibrium of the system leads to the implementation f the law of conservation of energy in the mode of steady motion. losed-contour theorem 1 A kinematic chain with two degrees of freedom, containing an nput link, an output link, and a movable closed chain placed beween them, can exist in a steady-state mode of motion. For the scheme ( Fig. 1 a), it is necessary to prove that the work or power) on the input link H 1 is equal to the work (power) on he output link H 2 , that is, H1 \u00b7 \u03c9 H1 = M H2 \u00b7 \u03c9 H2 . (1) We will assume that with the steady motion of the system in he form of a gear planetary mechanism, all the links move uniormly, and there are no inertia forces. To prove the theorem, we transfer all the acting forces to satelite 5 ( Fig. 1 c). A movable closed contour allows you to perform his transfer. According to Fig. 1 a the force F H1 = M H1 / r H1 acting at the point is transmitted to the link 5 from the carrier H 1 to the satelite 2 through the wheel blocks 3\u20136 and 1\u20134 in the form of rections R 65 = 0 . 5 F H1 r 3 / r 6 and R 45 = 0 . 5 F H1 r 1 / r 4 . The input force re- uced to link 5 is equal to the sum of these reactions F \u2032 H1 = . 5 F ( r r + r r ) / r r . Here r is the radius of the wheel i . H1 3 4 1 6 4 6 i K.S. Ivanov Theoretical and Applied Mechanics Letters xxx (xxxx) xxx F f p t t l F i t P a \u03c9 o e M q t s \u03c9 g M w a c d o d o e i t p f t a a s c c D f t a a v v i a f t l p c P T g c a e R r R a R t R ( f i n b a M e t fi d l v e f M R r p \u03c1 \u03c1 t r F m u \u03b7 The position of the point of application B \u2032 of the reduced force \u2032 H1 is determined by the formula KB \u2032 = r 5 ( r 1 \u2212 r 4 ) / r 4 . Reduction of orce F H1 to a link 5 corresponds to the condition of equality the ower of these forces F H1 V H1 = F \u2032 H1 V \u2032 H1 . The force F H2 is applied at he point K. Next, consider the equilibrium of link 5 ( Fig. 1 c). The sum of he moments relative to the instantaneous center P of the link veocities is zero \u2211 M P = 0 . Or \u2032 H1 \u00b7 ( P K + KB \u2032 ) \u2212 F H2 \u00b7 P K = 0 . (2) Here KB \u2032 = e is the eccentricity. From here, the position of the nstantaneous velocity center P (distance P K) can be determined hrough the given forces. K = F \u2032 H1 e/ ( F H2 \u2212 F \u2032 H1 ) . (3) Taking into account the motion of link 5 around a point P with n angular velocity \u03c9 5 , you can use the substitutions in Eq. (2) P K + KB \u2032 = V \u2032 B / \u03c9 5 , P K = V K / \u03c9 5 ", + " In the consider case, it is possible to use the relaionship of the friction moment M f K in the hinge K of the satelite 5 with the angular velocity of the satellite relative to the out- ut carrier \u03c9 5 \u2212H2 = \u03c9 5 \u2212 \u03c9 H2 . In Eq. (4) this constraint will be inluded in the form of power consumed by friction in the hinge f K = M f K \u03c9 5 \u2212H2 . heorem 2 on the definability of motion The definability of motion of a kinematic chain with two derees of freedom can be provided by an additional mobile-limiting onstraint of force with speed (or moment with angular velocity) s a function of power P f K = M f K \u03c9 5 \u2212H2 included in the equilibrium quation. For proof consider the equilibrium of link 5 ( Fig. 1 c). The force 05 is determined from the second condition of the link 5 equilibium \u2013 the sum of the forces is zero \u2211 F = 0 . From this condition 05 = F H2 \u2212 F H1 . (6) The definability of the motion of a two-mobile system can be chieved by replacing the moment M 05 of this unbalanced force 05 relative to the point K (which is stationary at the start) with he friction moment M f K in the satellite hinge K. The replacement (balancing) moment of friction M f K = M 05 = 05 \u00b7 P K. After substituting the value R 05 from Eq", + " Ivanov Theoretical and Applied Mechanics Letters xxx (xxxx) xxx 0 h m o i f u T w a l o c t v fi M s t k [ b o p M a d w r s t 0 c r f s M r s M c p M f M c P d n M f M i s d m n a g g g t At the start (when starting from the place) we can accept P f K = . 4 M H1 \u03c9 H1 . The start is performed without using the clutch. We ave \u03b7 = 0 . 6 . On direct transmission (at \u03c9 H2 = \u03c9 H1 ) P f K = 0 . We have \u03b7 = 1 . The use of a movable-limiting constraint with a balancing moent of friction in the hinge K ( Fig. 1 b) will provide a wide range f control at a low relative speed of the links. The angular veloc- ty line \u03c9 5 can rotate under the action of the resistance force F H2 rom the position providing \u03c9 H2 = 0 (shown by the dotted line) ntil \u03c9 H2 = \u03c9 H1 , when the dotted line coincides with the line \u03c9 H1 . hus, the range of gear ratios of the mechanism u H 1 \u2212H 2 = \u03c9 H1 / \u03c9 H2 ill be within 1 u H 1 \u2212H 2 \u221e . The main task of the adaptive drive is to overcome a given varible moment of resistance corresponding to the output angular veocity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002167_jae-201576-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002167_jae-201576-Figure3-1.png", + "caption": "Fig. 3. Arrangement of conductor coil to constraint eddy current for simulation purpose.", + "texts": [ + " To express this phenomenon in a 2D analysis, the eddy current will have to be constrained. The FEM conductor was set for each magnet in the eddy current analysis, Table\u00a01 Design specification of concentric magnetic gear Size (mm) High speed pole pair (pi) Low speed pole pair (po) Pole piece (np) Gear ratio Overall radius: 90 12 18 2 Inner yoke: 63.5 14 20 2.33 Outer yoke: 80 16 22 2.66 Inner rotor radius: 68.5 6 18 24 3 Outer rotor radius: 75 Stack length: 30 20 26 3.33 Inner air gap: 0.5 22 28 3.66 Outer air gap: 1 24 30 4 as shown in Fig.\u00a03. Opening the terminal on one side of the FEM conductor in the circuit will allow the constraint. Another condition is to set the insulation between the magnets to avoid eddy current being short-circuited to the adjacent magnet. Eddy current loss is expected to increase as the speed increases, but if the gear ratio increases, the PM segmentation can suppress losses due to eddy current\u00a0[34]. The design specification of the simulation and speed setting for all gear ratio is shown in Tables\u00a01 and\u00a02" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001726_isam.2007.4288440-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001726_isam.2007.4288440-Figure3-1.png", + "caption": "Fig. 3. Fishbone model", + "texts": [ + " In this paper, we assume that a belt object is inextensible. Then, the deformed shape of the object corresponds to a developable surface. It means that the object bends in direction d1 and it is not deformed in direction d2. Namely, a line the direction of which coincides with direction d2 is kept straight after deformation. In this paper, the central axis in a longitudinal direction of the object is referred to as the spine line and a line with zero curvature at a point on the object is referred to as a rib line as shown in Fig.3. We assume that bending and torsion of the spine line and direction of the rib line of each point specifies deformation of a belt object. This model is referred to as a fishbone model in this paper. Let \u03b1(u, 0) be rib angle, which is the angle between the spine line and direction d1 as shown in Fig.4-(a). Consequently, the shape of a straight belt object can be represented using five variables \u03c6(u), \u03b8(u), \u03c8(u), \u03b4(u), and \u03b1(u). Note that they depend on only the distance u from one end of the object along the spine line" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002859_s12046-020-01541-9-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002859_s12046-020-01541-9-Figure5-1.png", + "caption": "Figure 5. Cooling curve at three points of the same section thickness.", + "texts": [ + " This variation in the eutectic arrest with section thickness can be explained by the fact that the precipitation of graphite during eutectic reaction leads to the release of energy due to latent heat, which balances the heat losses of the melt. Therefore the higher the extent of graphite precipitation higher will be the heat released, which leads to longer thermal arrest during the eutectic reaction. Consequently, the time of eutectic arrest is an indication of the amount of graphite formation at the eutectic reaction. It can be observed from figure 5 that the cooling curves at three different points on the same section thickness do not show any legitimate variation in the cooling curve pattern. The thermal arrest is similar in each cooling curve given in figure 5. It can also be observed in figure 4 that the degree of undercooling is higher in the thin section, whereas the 15 mm section shows almost no undercooling with a comparatively flatter cooling curve as against those of 5 and 10 mm thickness. It is generally accepted that undercooling and grain size are inversely related to each other; therefore it can be said that the grain size in the thin section will be smaller; consequently, the number of grains in thin section sections will be more compared with thick sections of the casting" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002226_012033-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002226_012033-Figure4-1.png", + "caption": "Figure 4. Stand for research of the spline connection:", + "texts": [ + " In addition, monitoring several couplings simultaneously allows establishing statistical levels of the damping coefficient in each frequency range of the fractional octave spectrum: the base level (B), as well as warning (W) and danger (D) thresholds (Figure 3). To simulate the main dynamic vibrations of the prongs of the spline connection of the tail shaft couplings of the Mi-26 helicopter when the nature of the loading effect changes, a laboratory stand was made to monitor the technical state of the spline model (Figure 4). Thus, a technique for diagnostics of tribological couplings of the tail drive couplings of Mi-26 helicopters has been proposed. It allows identifying in real time the stability of elastic-inertial and dissipative characteristics of frictional interaction, periods of running-in, normal operation and catastrophic wear using the values of fractional octave spectra. Dynamics of Technical Systems (DTS 2020) IOP Conf. Series: Materials Science and Engineering 1029 (2021) 012033 IOP Publishing doi:10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002408_tmag.2021.3064023-Figure9-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002408_tmag.2021.3064023-Figure9-1.png", + "caption": "Fig. 9. No-load flux distribution of ladder rotor machine. (a) Flux linkage of phase A+ is maximum. (b) Flux linkage of phase A+ is minimum.", + "texts": [ + " 5(b), the ladder rotor structure can effectively reduce the low-order torque harmonic and optimize the torque ripple while maintaining the same torque. SRM is considered to be a promising high-speed machine type because of its no slip rings or brushes, simple structure and high reliability, and strong fault tolerance. The machine proposed in this article inherits the above advantages of SRM, but the vibration and noise level will be significantly reduced due to its very smooth current waveform. The no-load flux distribution of ladder rotor machine is shown in Fig. 9. For HSVRM, when the dc component is fed into winding, only the excitation field is established, which induces back electromotive force (EMF) in the winding and the machine will not produce constant torque. As illustrated, when the flux linkage of Phase A+ is maximum, local saturation exists in tooth tips, and the flux paths are much longer. In addition, when the flux linkage of Phase A+ is minimum, the path of the flux line comes back from the adjacent teeth, the flux line is concentrated in one of the stator yokes in which local saturation exists" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000848_coase.2008.4626552-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000848_coase.2008.4626552-Figure1-1.png", + "caption": "Figure 1: Schematic diagram of c-groove traps used to capture silicon elements. The plastic template, which contains the array of c-groove traps, is sandwiched between 2 glass slides to form a narrow fluidic channel to transport the silicon elements through the array.", + "texts": [ + " Various microcomponents can be independently batch microfabricated, released from their substrates into a powder-like collection of microcomponents, and then integrated onto a template via self-assembly using a variety of mechanisms including fluid flow, gravity, and electromagnetic forces.[1-3] In this paper we show how free-standing circular silicon microcomponents can be self-assembled onto plastic substrates, using an array of microfluidic traps. The template used in our study is a polyethylene terephthalate (PET) substrate patterned with electrical interconnects and an array of semi-circular microfluidic traps made of SU8. The shape of the trap is complementary to the circular shape of the elements. As shown in Figure 1, the template is sandwiched between two glass slides forming a narrow channel for the flow of the self-assembly fluid which contains a suspension of microcomponents. The self-assembly medium carries the microcomponents through this channel containing the array of microfluidic traps. When a silicon microcomponent flows into a trap, it is pinned down against the fluid flow, positioning them correctly over their electrical connections. This method of self-assembly gives us a number of benefits over existing paradigms: 1) Gravitational affects are minimized, so that we can scale down the size of the individual components 2) Assembly of thousands of components in parallel, scalable to larger arrays 3) Gives electrical and mechanical connections to the substrate in a single step 4) Self-alignment down to the micron scale without need for cameras and high precision motion stages This results in an assembly process which is high throughput, cost-effective, versatile, and scalable which may allow for entirely new approaches to fabricate new and novel devices in the future" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001281_ac60207a082-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001281_ac60207a082-Figure2-1.png", + "caption": "Figure 2. paratus Diagram of complete ap-", + "texts": [ + ", Amstel 21, Amsterdam (Netherlande). motor, part no. 362479-2), which can be rotated in either direction (Figure 1). Two switches, SI and &, serve to stop the motor when the piston has reached the end of its travel. The motor drives three identical gears (Honeywell gear, part no. 79172). One of these is fitted on the buret drive pulley; another one is coupled to a 10-turn potentiometer, which yields the analog output signal. Mounted on the third gear is a disk provided with 100 circumferential slots (Figure 2). The light emitted by a lamp mounted over the disk is intermittently allowed to pass on to a photoresistor (lamp and photoresistor-unit, Philips PR 901%) which is coupled to an electromagnetic counter via a bridge amplifier circuit (Figure 3). If not needed, one of the outputs can be omitted. A full piston stroke-i.e., 10 revolutions of the gear-consequently corresponds to 1000 counts. Working with piston burets invariably gives rise to hysteresis troubles caused by the inevitable backlash in the driving elements, deformation of the Teflon piston upon reversal of the direction of rotation, and changes in the volume of air bubbles, if present, Hysteresis can be eliminated by with- drawing some titrant from the filled buret and zeroing the counter before starting with the titration" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001466_icnc.2007.597-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001466_icnc.2007.597-Figure2-1.png", + "caption": "Fig. 2 Simplified five-link model", + "texts": [ + " Virtual humanoid robot has 6 DOFs(degree of freedom) on each leg, 5 DOFs on each arm, and 3 DOFs on head, with the result that the virtual humanoid robot has 25 DOFs. The structure of humanoid robot and the DOFs are presented in Fig. 1. To actualize virtual humanoid robot a transmission mechanism is employed, all joints are driven by DC motors, almost all joints have harmonic drive gears and pulleys for gaining a drive torque. Fig.1 Structure Scheme of humanoid robot During walking, the arms of the humanoid robot will be fixed on the chest. Therefore, it can be considered as a five-link biped robot in the saggital plane, as shown in Fig. 2. The motion of the biped robot is considered to be composed of a single support phase and an instantaneous double support phase. 3.1. Object function Firstly, gait relative parameters are defined as follows: length of upper leg, 2l length of lower leg, step length, height of knee rise, height of sciatic rise, b walking period of humanoid robot, 1l D qH bH T bV walking velocity of humanoid robot, , walking period of knee joint, qV walking velocity of knee joint, . bb TDV /= qT qq TD/V = During walking, humanoid robots adopt smooth wave gait" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002862_09544062211016076-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002862_09544062211016076-Figure5-1.png", + "caption": "Figure 5. Topological schematic drawing of the prototype: (a) schematic drawing of the original compliance distribution; (b) schematic drawing of VJM model.", + "texts": [ + " In order to achieve comprehensive optimization of configuration parameters, all the possible deformations of each part should be considered. All the elastic parameters are obtained from virtual experiments in a finite element environment19 and the results are shown in Table 2 in the Appendix. Meanwhile, the stiffness of the base is considered together with the actuators. The compliance distribution would alter after virtual joints are added to the alignment mechanism and relevant schematic drawings is shown in Figure 5(b). Because of the limited experimental condition, it is improper to carry out static tests on the prototype. And based on the engineering demand, the dynamic method would be adopted to verify the obtained VJM model instead, which would take the mode frequencies as a kind of convenient index of the holistic static stiffness characteristics. Considering that all the compliance parameters are obtained from finite element (FE) model, it is necessary to ensure the verification of FE model before verifying the VJM model and thus the virtual dynamic experiments were also carried out" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001644_iros.2008.4650858-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001644_iros.2008.4650858-Figure1-1.png", + "caption": "Fig. 1. Virtual trailer link model for vehicle following. At any time instance, the follower vehicle perceives the pose of the leader with an on board sensor. The pose, T, (xT , yT ) of the virtual trailer link is then computed. The follower vehicle will be commanded to the new position, T. The whole process will be repeated at the next time instant.", + "texts": [ + " However, as presented in our earlier work [17], it was demonstrated that vehicle following can be achieved by implementing a virtual trailer link model. In that model, the leader vehicle is modelled as the towing vehicle and the follower as the trailer. As the model suggested, the leader vehicle (towing) is effectively pulling a follower vehicle (trailer) via a virtual trailer link. Also, it has been proven that the length of the virtual trailer link must equal the length of the follower vehicle itself for an intrinsically safe vehicle following system (figure 1) [17]. If a chain of vehicles is to follow a leader, vehicles further down the chain suffer from a phenomenon known as string stability [2]. This issue can be addressed using this virtual trailer link model [17]. Definition 2: With the virtual trailer link model [17], vehicle following is redefined as: \u2016xF (t + \u03b4t) \u2212 xT (t)\u2016 = 0 \u2200 t > 0 (4) where xT (t) is the pose of the virtual trailer at time t and \u03b4t is the time increment between measurement. With reference to figure 1, the motion model, in discrete time state space, of the follower vehicle can be represented as1: xF,(k+1) = f(XL,k,XT,k,Uk, \u03c9k) (5) where XF,k, XL,k and XT,k are the histories of the state of the follower vehicle, leader and virtual trailer respectively. Uk is a vector of motion control signals input to the follower vehicle, \u03c9k is the motion uncertainty and f(\u00b7 \u00b7 \u00b7) is a nonlinear function representing the motion of the follower. The sensor is used to acquire the noisy observation zk of the leader vehicle taken from the follower" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000332_50003-4-Figure3.29-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000332_50003-4-Figure3.29-1.png", + "caption": "Fig. 3.29. The corresponding Gough plot.", + "texts": [ + "74) omitted) produces the moment - M z : if I tanal - Io-yl _< Or, st ~- l /Or 1 - 310y~y I + 3 ( 0 ; ~ y ) = - 10;~yl ' M z - l a (3.76) t a - - F 3 ya 1-10y yl+' \"3(0y% )2 and else: t~ = 0. The graph of Fig.3.27 clarifies the configuration of the various curves and their mutual relationship. For different values of the camber angle 7 the characteristics for the force and the moment versus the slip angle have been calculated with the above equations and presented in Fig.3.28. The corresponding Gough plot has been depicted in Fig.3.29. The relationship between 7 and the spin ~o follows from Eq.(3.55). The curves established show good qualitative agreement with measured characteristics. Some details in their features may be different with respect to experimental evidence. In the next section where the simulation model is introduced, the effect of various other parameters like the width of the contact patch and the possibly camber dependent average friction coefficient on the peak side force will be discussed. The next item to be addressed is the response to large spin in the presence of side slip" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003278_j.vacuum.2021.110494-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003278_j.vacuum.2021.110494-Figure1-1.png", + "caption": "Fig. 1. Established space and plane coordinate system of the screw rotor.", + "texts": [ + " Furthermore, it can also provide a potential new method to design the rotor with good machinability in such a way that the rotor profile in longitudinal section can be designed first, containing no concavity, and then its profile in transverse section can be deduced and checked if it is self-conjugated curve [22,23], or vice versa. In order to process the screw rotor with high precision, it is necessary to collect enough coordinate points on each surface of the rotor, thus the transformation of the profile curve equations in different section is involved. Before describing the transformation, the establishment of coordinate system of screw rotor is needed. Fig. 1 shows the space rectangular coordinate system established. Its XOY plane is located in the end of the screw rotor with its origin O coinciding with the rotor center, and the rotation axis of the rotor as Z axis. For the deduction purpose, another two plane coordinate systems O1(x, y) and O2(x \u2032 , y\u2032 ) are also established. The former contains X, Y and O, and the later X, Z and O as shown in Fig. 1. Then four planes of the screw rotor are defined. The transverse section of the rotor at Z = 0 is defined as profile plane 1, which is obviously in the coordinate system O1(x,y). The transverse section of the rotor at Z = h is defined as plane 2, and h is the distance between plane 1 and 2. The longitudinal section at Y = 0 is plane 3 which is in the coordinate system O2(x \u2032 ,y\u2032 ). The section of the rotor at Y = a is defined as plane 4, and a is the distance between plane 3 and 4. Besides that, r * Corresponding author" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003385_ur52253.2021.9494662-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003385_ur52253.2021.9494662-Figure2-1.png", + "caption": "Fig. 2. 6DOF manipulator 3D model", + "texts": [ + " The 6-DOF of the cobot do not include the one DOF of the gripper, and therefore, a cobot has a total of seven DOF, including that of the gripper. The designed payload of the 6-DOF manipulator is 1.5 kg, and the accuracy is expected to be a maximum of 2 mm in the no-load condition in terms of the end-effector. The working radius is 600 mm, in which tasks can be performed in a wide space. The cobot can be used to carry a 500 ml water bottle or a cup of coffee, or pick-and-place lightweight objects. The 6-DOF manipulator and its detailed specifications are presented in Fig. 2 and Table II, respectively. A mecanum wheel is generally mounted on a geared motor, which takes up a large space inside the mobile platform owing to a large volume in the lengthwise direction. Accordingly, the space inside the platform becomes narrow and the battery volume cannot be increased, resulting in a reduced operation time. Suspension equipment is also required when using a mecanum wheel so that four wheels 125 Authorized licensed use limited to: University of Glasgow. Downloaded on August 12,2021 at 13:49:11 UTC from IEEE Xplore", + " A mobile manipulator that can move in any direction at a high speed was developed by combining the 6-DOF manipulator and the mecanum mobile platform presented above. Fig. 5 shows the fully assembled mecanum mobile manipulator. The product of exponential formula (POE) was used to compute the kinematics of the manipulator. Unlike the Denavit\u2013Hartenberg (DH) parameters, which define by assigning coordinates to each link, only fixed coordinates {s} and end-effector coordinates {b} are defined. Each joint is assumed to execute screw motion, and the rotating axis is defined as the screw axis. Fig. 2 shows the screw axis of the 6-DOF manipulator. The homogeneous transformation matrix with respect to the end-effector can be computed as shown in (2). M is the homogeneous transformation matrix from {s} to the end-effector when the robot is in the zero position, and Bi is the screw axis of each joint in the zero position with respect to the end-effector, consisting of (\u03c9,v). T = Me[B1]\u03b81e[B2]\u03b82e[B3]\u03b83e[B4]\u03b84e[B5]\u03b85e[B6]\u03b86 (2) where, M = 1 0 0 0 0 1 0 0 0 0 1 l1 + l2 + l3 + l5 + l7 0 0 0 1 B = 0 0 1 0 0 0 0 \u22121 0 \u2212l2 \u2212 l3 \u2212 l5 \u2212 l7 0 0 0 \u22121 0 \u2212l3 \u2212 l5 \u2212 l7 0 0 0 \u22121 0 \u2212l5 \u2212 l7 0 0 0 0 1 l6 0 0 0 0 1 \u2212l7 0 0 The Jacobian is a function of the joint variable \u03b8, which transforms the rate of change of the joint variable \u03b8\u0307 into linear velocity v and angular velocity \u03c9 of the end effector" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002742_13506501211018937-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002742_13506501211018937-Figure8-1.png", + "caption": "Figure 8. The measurement areas.", + "texts": [ + " The requirements for measuring lines for the MBAC are not strict, the endpoints need not to be strictly controlled at both ends of the wear scar, and the accuracy of the BAC measurement in the wear area is improved by fitting multiple measurements. Moreover, the rough peak profile does not need to be matched and stitched, which is a faster and more convenient way to calculate wear. Table 2. Mechanical properties of H13 steel. Elastic modulus (GPa) Poisson\u2019s ratio Density (kg/m3) Hardness (HRC) Yield strength (MPa) Tangent modulus (MPa) Figure 4. Experimental sample: (a) upper sample, (b) lower sample, and (c) friction pair. Calculation of three-dimensional wear of the improved BAC The data for smooth surface areas (I) and (II) in Figure 8 are collected, as shown in Figure 9. The average BAC before and after wear is obtained by fitting these four calculated BACs. The average relocation value H is calculated to be 0.3871 \u03bcm according to the single displacement of each picture. The MBAC after the relocation is shown in Figure 10, and the equivalent wear depth HMBAC is calculated to be 0.3518 \u03bcm. The theoretical wear area in this experiment is 124.88 mm2, and the wear amount is calculated to be 0.0472 mm3. The MBAC takes the difference of the arithmetic mean before and after wear as the reference displacement in the MBAC method" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003345_ur52253.2021.9494692-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003345_ur52253.2021.9494692-Figure3-1.png", + "caption": "Figure 3. Isometric view of the simulator in MATLAB/ SimMechanics.", + "texts": [], + "surrounding_texts": [ + "In this section, the proposed control scheme has been discussed in detail. Nonlinear ESO has briefly explained in the first part, 2nd part describes the ISMC and the 3rd explained the integration of ISMC with ESO." + ] + }, + { + "image_filename": "designv11_83_0002583_s12034-021-02361-1-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002583_s12034-021-02361-1-Figure2-1.png", + "caption": "Figure 2. The fabrication of multilayer self-folding structure. (a) Three-layer assembly of polyimide tape (structural layer) and PVA/SCF composite film (shape change layer) and (b) the dimensions of the sample (dimensions are not to scale).", + "texts": [ + " Using CAD software and 3D printing technology, we could flexibly design and print the mould with the required shape and size. The cut PVA/SCF composite films were placed in the printed mould to obtain various permanent shapes through plasticity. After the electroactive shape memory PVA/SCF composites were prepared, a multilayer assembly technology was used to fabricate it into a multilayer self-folding structure. The self-folding structure was designed as a threelayer composite structure, whose shape and dimensions are shown in figure 2. The middle layer was a PVA/SCF composite film with a thickness of 0.2 mm, which was a shape changing layer. The upper and lower layers, made of 0.055 mm-thick polyimide tape (PI) with an adhesive on one side, formed the non-deformable structural layer. These two layers were adhered to the PVA/SCF composite film using precise cutting tools. Polyimide possessed the advantages of high temperature resistance and high insulation, which could well preserve the heat generated inside the material during ohmic heating" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002794_5.0049922-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002794_5.0049922-Figure2-1.png", + "caption": "FIGURE 2. Contour plot of stresses (a) Normal stress plot in the nineteenth ply, (b) shear stress plot in the eighty-ninth layer.", + "texts": [ + "125 mm have been used for analysis (Table 1). Internal and external diameters are 40 mm and 62 mm, respectively, the length of the driveshaft is 660 mm, so the number of layers is 90 [18]. The meshed model using shell 181 elements (Figure. 1) developed in ANSYS mechanical workbench having 756 elements and 768 nodes with average quality index 0.99 is used for analysis. The torsional strength required for a specific vehicle having sell growth of 9% every year is 3557 Nm [18]. This torque is applied on the glass-epoxy composite driveshaft. Figure 2 (a) shows the variation of normal stress in the nineteenth ply with maximum normal stress is at the outer fiber, which is compressive and equal to 204.02 N/mm2. The maximum shear stress is in the eighty-ninth ply as shown in Figure 2(b), equal to 45. 6 N/mm2. 020007-2 Figure 3, shows the principal stress in the global x-direction is varying linearly across each layer and suddenly changing at the interface. The principal stress is maximum at the middle of the second top fiber, which is 200N/mm2. The failure analysis is done based on local stress and strain in the layers. Figure 4, shows the local normal strain in the material direction 1 and in-plane shear strain. Both the stains are varying linearly with the distance in the individual lamia of the layer from mid-plane and abruptly changing at the interface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000775_detc2007-35917-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000775_detc2007-35917-Figure3-1.png", + "caption": "Figure 3. A beam simply-supported on a rotating rigid disk.", + "texts": [ + " The temperature constraint equations are given by ( , )t t =q 0\u03a6 (65) and the kinematic and driving constraint equations read ( )k =q\u03a6 0 (66) ( , )d t =q 0\u03a6 (67) where T T(1)T ( )T (1)T ( )T,N N = = q q q q q qL L represent the system generalized coordinates, and ( )iq and T( ) ( ) ( ) ( ) ( ) T 0 0 i i i i ix y \u03b8 = q p represent the generalized coordinates of iB , and then Lagrange\u2019s equations of the first kind are given by ( )Tt ttt t = \u2212 q q QM q& 0 \u03a6 \u03a6\u03bb\u03a6 (68) and ( ) ( ) ( ) ( ) T T 2 k d k k k d d d d d t tt = \u2212 \u2212 \u2212 \u2212 q q q q q q q qq M Qq q q q q q && & & & & & \u03a6 \u03a6 \u03a6 \u03bb \u03a6 \u03a6 \u03bb \u03a6 \u03a6 \u03a6 0 0 0 0 (69) where t q\u03a6 , k q\u03a6 and d q\u03a6 represent the Jacobian matrices, and t\u03bb , k\u03bb and d\u03bb represent the Lagrange multipliers. The system generalized mass and force matrices read (1) ( )N = M M M O , (1) ( )N = Q Q Q M (70) (1) ( )N = M M M O , (1) ( )N = Q Q Q M (71) 6 SIMULATION EXAMPLE 6.1 Nonlinear effect: A rotating hub attached with a simply-supported beam, which is shown in figure 3 is simulated. The geometric property and material data of the beam are given in table 1. Uniform heat generation is given by QT = 1\u00d7107 (w/m3). The rotating speed of the rigid disk is given by 5 Copyright \u00a9 2007 by ASME Terms of Use: http://www.asme.org/about-asme/terms-of-use 0 0 2 sin for 0 2( ) for , s s s s sd s s t t t t t tt t t \u2126 \u2126 \u03c0\u2126 \u03c0\u03c9 \u2126 \u2126 + \u2212 \u2264 \u2264= + > (72) where 0.1sst = , \u2126 0 = 4 rad/s, \u2126 s = 0.2 rad/s. The boundary conditions of the two tip ends are 1(0) 0u = , 2(0) 0u = , 1( ) 0u l = , 2( ) 0u l = " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002772_978-3-030-67750-3_14-Figure30.8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002772_978-3-030-67750-3_14-Figure30.8-1.png", + "caption": "Fig. 30.8 Side view of loader bucket", + "texts": [], + "surrounding_texts": [ + "bucket, which is 100 mm linear actuator and 150 mm linear actuator depending on the length of the arm needed to move. All these actuators are controlled by a motor located at the wheelbarrow body tray and near to the handle where the user can easily manage to operate. The connection harness attached under the wheelbarrow bucket is for safety reason. All the items can be referred in Figs. 30.8 and 30.9 that shows the detail view of the loader bucket." + ] + }, + { + "image_filename": "designv11_83_0000737_ijtc2007-44228-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000737_ijtc2007-44228-Figure2-1.png", + "caption": "Figure 2. Schematic of the instrumented bearing and electrode pattern.", + "texts": [ + " A 4 \u00b5m thick layer of aluminium nitride has been sputtered onto the outer raceway on the outer bore. A rectangle of silver paint has been applied to the coated surface as the top electrode. The bearing outer race is the bottom electrode. Wires are soldered 1 Copyright \u00a9 2007 by ASME ms of Use: http://www.asme.org/about-asme/terms-of-use Dow to these electrodes and then to a tag bonded to the bearing for durability. In the photograph (figure 1) the coating is difficult to see but the soldering tag and wired are visible. Figure 2 is a schematic of the instrumented bearing showing the sensor location and detail of the sputtered coating and upper electrode. The sensor has been coated onto the outer surface of the outer raceway. The electrode has dimensions of 0.3 mm by 3 mm. The high frequency ultrasonic wave has the property that it tends not disperse greatly. The size of the contact patch during these tests is an ellipse varying from 0.34 by 4 mm to 0.6 by 7.2 mm. The transducer was connected to an ultrasonic pulserreceiver (Panametrics 5072PR) that was used to excite, receive and amplify the reflected signals which were then passed to a digital scope and PC for storage and analysis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001034_ijvd.2008.021154-Figure17-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001034_ijvd.2008.021154-Figure17-1.png", + "caption": "Figure 17 Surface of a steel disk after test performed on SAE#2 machine (50 cycles): (a) original clutch design; (b) single-sided design and (c) inverted double-sided design", + "texts": [ + " For each of the three pack designs two tests were performed. Figure 16 shows a photograph of rubbing surfaces of metal disks from the original pack design taken after 50 engagement cycles. Five distinct hot spots, manifesting themselves by dark discolorations, were produced. The hot spots were present at all sliding interfaces across the pack, with fairly small differences in severity between the surfaces. They were aligned with the fingers, with slight circumferential shift in the direction of movement of the countersurfaces. Figure 17 shows the rubbing interface of steel disk from the three designs, each after 50 cycles. The two alternative pack designs show much lighter discoloration of sliding surfaces, with different pattern in each case. In the single-sided plates the discolorations were essentially band-shaped, while in the inverted double-sided pack, 14\u201316 light focal discolorations around the circumference occurred. Similar results were obtained in the second series of tests. These observations are consistent with finite element simulations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002969_j.matpr.2021.05.300-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002969_j.matpr.2021.05.300-Figure4-1.png", + "caption": "Fig. 4. Al oxide clutch plate heat flux results.", + "texts": [ + " Calculation of stresses, and greatest weights utilizing EulerLagrange conditions. Developing 3D model of MPCs utilizing Solid edge programming. Exporting the model of MPCs from strong edge to ANSYS 19.2 work bench for FEA. Static and dynamic examination of MPCs utilizing ANSYS work seat at various working condition. Determining the wear, most extreme disfigurement and equal worries for every material utilizing FEA Showing the connection of disfigurement and obstruction property of every material utilizing figures and tables (Fig. 1. Fig. 2. Fig. 3. Fig. 4. Fig. 5. Fig. 6. Fig. 7. Fig. 8. Fig. 9. Fig. 10. Fig. 11. Fig. 12. Table 1). Comparing and conversation of results for choice of better covering materials. Fig. 5. Cast iron clutch plate heat flux results. Fig. 6. Cast iron clutch plate heat flux results. Grip disappointment and harm because of extreme frictional warmth and warmth changes to the grasp counter mate circle frequently happens to a car grips. This circumstance adds to warm weakness to the segment which causes the grasp counter mate plate to break and twist" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000312_978-3-540-44410-7_8-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000312_978-3-540-44410-7_8-Figure2-1.png", + "caption": "Fig. 2. Single arm supported by a vehicle.", + "texts": [ + " z\u0307 = k#(\u00b5\u0307\u2212 kx\u0307) (13) At this point, while referring to Appendix C for some additional comments concerning condition (12) and related solution (13), we can conclude this section by simply noting that, in order to be comprehensive of the overall cases \u00b5\u2217 < \u00b5 < \u00b50 or \u00b5\u2217 < \u00b5 > \u00b50, while also avoiding any possible chattering in the vicinity of the threshold value \u00b50, it is actually always convenient to adopt the following expression for z\u0307: z\u0307 = (\u03b1k)#(\u00b5\u0307\u2212 kx\u0307) (14) being \u03b1(\u00b5) a continuous scalar function of \u00b5, which is unitary for \u00b5 \u2264 \u00b50 and bell-shaped, tending to zero within a finite support for \u00b5 > \u00b50. Obviously enough, with such final adjustment, the smooth transition between the two different cases of task priority turns out to be automatically guaranteed. 3 Control of a Single-Arm Nonholonomic Mobile Manipulator The case of a redundant arm mounted on a 3D moving base as in Fig. 2 is now considered. The vehicle is assumed to nonholonomic, in the sense that it allows a linear velocity vector \u03bd only directed along the principal vehicle axis, and an angular velocity vector \u2126 only lying on a plane passing through a known point of such principal axis, and orthogonal to it. The arm and the vehicle are regarded as two separate \u201cbasic robotic units\u201d, whose motions however needs to be suitably coordinated for the execution of a common task (i.e. making again < e > converge toward < g >) to be realized in a cooperative way", + " With the above considerations in mind, let us now approach the overall control problem of having < e > converging to < g > by again considering the candidate Lyapunov function (3). Its time derivative now takes on the form being x\u0307 and X\u0307 the contributions to the end-effector motion separately produced by the arm and the vehicle, respectively, both projected on world frame <0>. More specifically, for X\u0307 we actually have X\u0307 = S\u03b8\u0307 (19) where \u03b8\u0307 is the three-dimensional vector resulting from the collection of the two non-null components w of \u2126 and the sole non-null component u of \u03bd , provided that both \u2126 and \u03bd are projected on the vehicle fixed frame < b > as indicated in Fig. 2; that is \u03b8\u0307 = [ wT u ]T . (20) Also in (19) the matrix S = HQ (21) where is a (6\u00d76) matrix representing the instantaneous rigid-body velocity transformation from vehicle frame < b > to the end-effector frame < e > (input velocities projected on < b >, output velocities projected on world frame < 0 >), while Q is simply a full-rank (6\u00d7 3) selection matrix, suitably composed by 0 and 1 elements. Notice that the (6\u00d7 3) matrix S is also full-rank. Folding (19) into (18) gives V\u0307 = \u2212eT (x\u0307 + S\u03b8\u0307). (22) At this point, by choosing the Cartesian reference velocity x\u0307 in (22) as in the form (15) and (16) yields V\u0307 = \u2212eT ( \u02d9\u0302x + z\u0307 + S\u03b8\u0307) = \u2212eT [ (x\u0307 + \u03b6\u0307) + z\u0307 + S\u03b8\u0307 ] " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001810_iccas.2008.4694328-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001810_iccas.2008.4694328-Figure1-1.png", + "caption": "Fig. 1 Two-dimensional humanoid models.", + "texts": [ + " This chapter describes the kinematics of a humanoid robot in consideration of variable-speed biped walking and direction turning. Humans can walk, run, and twist by rotating joints whose axes are placed normal in three orthogonal planes; the sagittal, coronal, and transverse planes. These planes are determined from viewing angles to an object human. In the sagittal plane, a human walks forward or backward, while he or she moves laterally by rotating joints in the coronal plane. It is apparent for turning the body that the joints locating in the transverse plane have to be rotated. Fig. 1 illustrates the present humanoid model described in the three planes. There are three types of angles in the model; \u03b8 , \u03c6 , and \u03c8 . The sagittal angles 6,,1, =ii\u03b8 are the joint angles that make a robot proceed forward. As shown in Fig. 1(a), the joint angles 21, \u03b8\u03b8 , and 3\u03b8 are associated with the supporting leg and an upper body, while 54 , \u03b8\u03b8 , and 6\u03b8 are assigned to move a swaying leg. Treating the motor angles is more straightforward and simpler than using the sagittal angles. These motor angles are shown in Fig. 1(a) as rljthknanij i ,,,,, ==\u03b8 , where the subscripts an, kn, and th represent ankle, knee, and thigh, and the superscripts l and r stand for left and right, respectively. Every sagittal joint motor has its own motor angle that has a one-to-one relationship with the six sagittal angles [1]. The coronal angle \u03c6 plays the role of translating the upper body to the left or right for stabilizing the robot at the single support phase. Fig. 1(b) shows that there are four joints lying on the coronal plane in the present model. These joints, however, consistently revolve with 1 DOF to maintain the posture in which the upper body and the foot of a swaying leg always are made vertical to the ground in the coronal view. The transversal joints with revolute angles rlii ,, =\u03c8 are to twist the upper body around each axis and thus change the current walking direction. The left transverse joint ( l\u03c8 ) can undertake a right-hand turning by first revolving in left leg supporting and by second doing in right leg supporting", + "723 lyy += For the modeling of motions in the transverse plane, we take into account the case of a right-hand rotation using the left transverse joint as a pivot axis. This joint must revolve twice for completion of turning in standing; first for body rotation, and second for restoring the posture of standing with parallel foots. The first rotation resulting from actuation of l\u03c8 leads to the circular motion of all the joints except those installed in the supporting leg. Since the coordinate of the left transverse joint is ),,( 222 zyx as shown in Fig. 1, the resultant coordinates of the k-th joint 6,,3),,,( =kzyx t k t k t k are derived by using the following rotation matrix: \u23a5 \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a2 \u23a3 \u23a1 + \u23a5 \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a2 \u23a3 \u23a1 \u2212 \u2212 \u2212 \u23a5 \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a2 \u23a3 \u23a1 \u2212 = \u23a5 \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a2 \u23a3 \u23a1 2 2 2 2 2 2 000 0cossin 0sincos z y x zz yy xx z y x k k k ll ll t k t k t k \u03c8\u03c8 \u03c8\u03c8 (4) where the superscript t implies that the coordinates are rotated on the transverse plane. Owing to the configuration of coronal angles, the pelvis bone 7l is parallel to the ground, and thus the z-coordinates remain the same as shown in Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002014_inmic.2008.4777701-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002014_inmic.2008.4777701-Figure1-1.png", + "caption": "Figure 1. The schematic of grid connected DFIG. Figure 2. Definition of the positive current, voltage and flux directions.", + "texts": [], + "surrounding_texts": [ + "Keywords- Doubly fed induction generat; Output powe; PI controller; PSO algorithm\nII. DFIG EQUATIONS IN ABC-FORM\nThe convention adopted in this paper for positive current, voltage and flux directions is shown in Fig. 2. As the machine is working in generator mode, positive currents are flowing out of it.\nThe sign of the self-flux linkage produced by a current in a circuit is the same as that of the current. The polarity of the voltage induced by a changing flux is so that it results in a current that opposes the change (Lenz's law). Applying the Kirchhoff voltage law to Fig. 2 we have:\nWhere R s ' Vas, ias ' and VIas are in [p.u] and are the stator phase-a winding resistance, voltage, current and flux linkage, respectively; (1JB [rad/sec] is the system base frequency which\nis equal to the synchronous frequency, i.e. (1JB =2nf ; and t is\nthe time in [sec]. In (1), the flux VIas is:\n(1)\n(2)\n'IIas = (Lse/j,s + L1eak,s )ias + Lmut,s (ibs + ics ) +...\nL sr (cos(Or )iar +cos(Or + 21! )ibr +cos(Or - 21! )icr ) 3 3\nI. INTRODUCTION\nWind energy is one of the most important and promising sources of renewable energy all over the world, mainly because it is considered to be nonpolluting and economically viable. At the same time, there has been a rapid development of related wind turbine technology [1, 2]. Nevertheless, this kind of electric power generation usually causes problems in the electrical system it is connected to, because of the lack or scarcity of control on the produced active power.\nDue to its many advantages such as the improved power quality, high energy efficiency and controllability, etc. the variable speed wind turbine using a doubly fed induction generator (DFIG) is becoming a popular concept and thus the modeling of the DFIG based wind turbine becomes an interesting research topic.\nThis paper aims to present a suitable method based on PSO algorithm followed to control the DFIG active power. To this end, the system non-linear dynamical model is derived.\n978-1-4244-2824-3/08/$25.00 \u00a92008 IEEE", + "where Or = Or (t) is the angle between the stator a-axis\n(stationary) and rotor a-axis (rotationary) as shown on Fig.2; Lse/j,s and Lleak,s are the self and leakage inductance of a\nstator winding, respectively; Lmut,s is the mutual inductance\nbetween two stator windings; and Lsr is the peak value of the mutual inductance between stator and rotor windings. For the other phases of the stator and rotor, similar equations can be written.\nIII. ABC-DQ TRANSFORMATION\nFor easier control, three-phase variables are transformed into dq-variables. In matrix notation, we have:\n(3)\nfJ/ qs = L ss iqs + L m i qr\nfJ/ ds = L ss ids + L m i dr\nfJ/ qr = L rr i qr + L m i qs\nfJ/ dr = L rr i dr + L m ids\n(5)\n(6)\nWhere vqdo T = [vq Vd vo], Vabe T =[va Vb Vel and\nTo is the abc 0 dq transformation matrix. In this paper, the\npower invariant transformation is chosen, and the d-axis is leading the q-axis. Fig. 3 shows the dq-frame with respect to the stator 3-axis frame.\nThe corresponding transformation matrix is given in (4).\nIn (5), OJ is the rotational speed of the dq-frame, i.e. 0) = dO where \u00b0= O(t) is the angle between the d-axis and dt\nstator a-axis (Fig. 3); and OJr is the rotational speed of the\nrotor, i.e. O)r = dOr . For the synchronously rotating frame, dt OJ is the synchronous speed, thus in [pu] OJ =OJs =1 . In (6), Lss =LSe/j,s + Lleak,s - Lmut,s and L m =L sr\nsinB sm(e-2:) sm(e+ 2:) T8=~ cosO cos(e- 2:) cos(e+~)\n1 1 1\nJ2 J2 J2\n(4)\nIV. ELECTROMAGNETIC TORQUE EQUATIONS\nThe electromagnetic torque Te is obtained by dividing the\nair gap power by the mechanical speed of the DFIG's rotor [3].\n(7)\nApplying (4) to (1)-(2) and the other stator and rotor phase equations, gives the DFIG dq-model in [pu]:\n(8)\nIf the turbine, gearbox, shafts and generator are lumped together into an equivalent mass H tot [s], the following swing Equation in [pu] models the drive train:\n(9)\nIn (9) D[p.u.s / roo] is the damping coefficient.\nV. CONVERTER MODEL\nIn this paper, the converter connected between the DFIG rotor and the grid consists of two PWM voltage source inverters (VSI) separated by a dc-link as shown on Fig. 4. The grid side converter varies its modulation index ll12 in order to maintain the dc-voltage Vdc constant and operates at unity power factor, i.e. only the stator delivers the required reactive power to maintain the terminal voltage (stator voltage); the" + ] + }, + { + "image_filename": "designv11_83_0000790_s1560354708040072-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000790_s1560354708040072-Figure1-1.png", + "caption": "Fig. 1. A hexagonal pencil whose edges have width a rolling about a temporarily fixed edge on a plane of inclination \u03b8 to the horizontal. The plane containing the axis of the pencil and the instantaneous axis of rotation makes angle \u03b1 to the vertical.", + "texts": [ + " This problem is an extensive elaboration of considerations of impulsive motion given, for example, by Routh [3]. 2. ANALYSIS FOR N = 6 VIA A COEFFICIENT OF (IN)ELASTICITY If the plane has angle of inclination \u03b8 < \u03c0/6, the center of mass of a hexagonal pencil rises during the first part of any 1/6 turn. Hence, the pencil will not roll down the plane from rest unless \u03b8 > \u03c0/6. 2.1. \u03b8 > \u03c0/6 In this case a (sharp-edged) pencil rolls down the plane spontaneously from rest.1) The equation of motion of the rotating pencil is \u03c4 = I\u03b1\u0308 = mga sin \u03b1, (2) where, as shown in Fig. 1, \u03b1 is the angle to the vertical of line from the instantaneous axis of rotation of the pencil to its center of mass, a is the width of each of the six faces of the pencil, m is the mass of the pencil, I is the moment of inertia of the pencil about an edge, and g is the acceleration due to gravity. During each 1/6 turn, \u03b8 \u2212 \u03c0/6 \u03b1 \u03b8 + \u03c0/6. (3) The equation of motion (2) is a so-called Mathieu equation, which does not lend itself to analytic solution. Instead, we use an energy analysis to estimate the asymptotic linear and angular velocity of the rolling pencil", + " (8) The asymptotic time-average angular velocity is related to the asymptotic time-average kinetic energy of the pencil by 1 2 I \u3008\u03c9\u221e\u30092 = \u3008E\u221e\u3009 , (9) where I = kma2 is the moment of inertia of a hexagonal pencil about an edge, k = 17 12 (solid hexagon), k = 11 6 (hexagonal shell). (10) REGULAR AND CHAOTIC DYNAMICS Vol. 13 No. 4 2008 One way to estimate the time-average asymptotic kinetic energy is to replace the average over time by an average over angle \u03b1, \u3008E\u221e\u3009 \u2248 E\u221e + 3 \u03c0 \u03b8+\u03c0/6\u222b \u03b8\u2212\u03c0/6 mga[cos(\u03b8 \u2212 \u03c0/6)\u2212 cos \u03b1] d\u03b1 = mga 2 [ 1 + \u03b5 1 \u2212 \u03b5 sin \u03b8 \u2212 ( 6 \u03c0 \u2212 \u221a 3 ) cos \u03b8 ] , (11) recalling Fig. 1. This analysis suggests that there is a minimum angle \u03b8min for steady rolling given by tan \u03b8min = 1 \u2212 \u03b5 1 + \u03b5 ( 6 \u03c0 \u2212 \u221a 3 ) = 0.178 1 \u2212 \u03b5 1 + \u03b5 . (12) The empirical evidence that steady rolling can exist for very small \u03b8 suggests that the coefficient of (in)elasticity \u03b5 is close to unity. However, we will find a slightly more restrictive limit on \u03b8min in Section 2.2. Combining Eqs. (8), (9) and (11) we estimate the asymptotic linear velocity to be \u3008v\u221e\u3009 = 3 \u03c0 a \u3008\u03c9\u221e\u3009 \u2248 3 \u03c0 \u221a ag k \u221a 1 + \u03b5 1 \u2212 \u03b5 sin \u03b8 \u2212 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000880_kem.381-382.93-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000880_kem.381-382.93-Figure3-1.png", + "caption": "Fig. 3 Multiple measurements on gear pitch Fig. 4 Result of multiple measurements", + "texts": [ + "147, UB Giessen, Giessen, Germany-29/05/15,09:11:01) The multiple measuring technique has been used for symmetric shape artifacts, e.g. cylinder [2]. The technique is applicable to a gear pitch (circumferential pitch) measurement. The techniques eliminates for all geometrical errors of a CMM, systematic probing effects. The biggest uncertainty source is a probing error on the gear pitch measurement. The probing system of the CMM has a characteristic error and the error is systematic. Therefore, the error is eliminated by the multiple measurement technique. Fig.3 shows the schematic diagram of the multiple measurement technique. When we measure the gear of which pitch diameter is 90 mm and the module is 2.5, the number of tooth is 36 and the pitch angle is 10 degrees. First, we measure the gear pitch by a CMM and then we rotate the gear 10 degrees and measure the gear again. We repeat the same procedure till 36 times and then, we average the 36 measurement results. Fig. 4 shows the result of the multiple measurement technique on the gear measurement. The figure shows the systematic errors are eliminated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002208_012043-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002208_012043-Figure2-1.png", + "caption": "Figure 2. Distribution of forces of lateral and longitudinal creep at frictional contact a) Fe-Fe b) Fe-Al", + "texts": [ + " As it was mentioned above, when using the TMC for wear protection of the working surfaces of the flange, one can obtain additional tractive effort due to the frictional connection P of the surfaces of the flange and the side surface of the rail head. This two-point state of interaction of the locomotive wheel with the rail occurs when the locomotive fits into the curves, and when the required tractive effort increases significantly to compensate for the additional resistance of the trailed rolling stock when it moves in curved track sections. It also happens under wind load when moving with a banking engine. The traction force scheme is shown in Figure 2. The possibility of existence of this traction two-point contact is associated with practically equal sliding speeds of the wheel relative to the rail at points 1; 2 and the absence of the output of the traction force curve at point 2 on the horizontal sections of the traction characteristic page Fig. 4. Issues related to the wear and tear of the wheel sets flanges have always been and do remain relevant. Reducing the wear rate of the working surfaces of the wheels flanges to the level and below the level of the wear rate of the surfaces of the rolling circles makes it possible to reduce the number of wheel turning due to the \"thin\" flange and, accordingly, to increase the number of turns on the \"rolling\" of the tread " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001178_20080706-5-kr-1001.00083-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001178_20080706-5-kr-1001.00083-Figure4-1.png", + "caption": "Fig. 4. Determination of periodic motions and decaying oscillations", + "texts": [ + " It is based on a modification of the DF method that involves consideration of instantaneous amplitude, frequency and decay rate instead of respective constant values. Now, with those methodology and results available, let us look at the problem of the existence of periodic motions, asymptotic decay and finite time convergence considering the harmonic balance in the system. Consider application of the concepts of the DF method to analysis of possible periodic motions. Periodic motions can exist in the system if the Nyquist plot of the linear part )( \u03c9jW intersects the negative reciprocal of the DF )(1 aN \u2212\u2212 (Fig. 4). In Fig. 4, two Nyquist plots corresponding to the second- )(1 \u03c9jW and third-order )(2 \u03c9jW linear parts and two negative reciprocal DF corresponding to the relay control )(1 1 aN \u2212\u2212 and to the twisting algorithm )(1 2 aN \u2212\u2212 (Boiko et al., 2004) are depicted. Intersection of )(2 \u03c9jW and either of the DFs provides a periodic solution (points A or B) of finite frequency and amplitude. Plot )(1 \u03c9jW does not have any points of intersection with either )(1 1 aN \u2212\u2212 or )(1 2 aN \u2212\u2212 except the origin. However, the character of the process in the system is different \u2013 depending on whether the control is a conventional ideal relay (plot )(1 1 aN \u2212\u2212 ) or the SOSM control (plot )(1 2 aN \u2212\u2212 )", + " As for the second option, a periodic motion cannot occur at any frequency (including \u221e=\u03a9 ). There is a condition that we shall further refer as a phase deficit. Quantitatively, let us call the phase deficit the minimum phase value that needs to be added (with the negative sign) to the phase characteristic of the linear part to make the phase balance condition hold at some frequency (including the case of \u221e=\u03a9 ). Note: we do not consider now the case of possibly non-monotone frequency characteristics. The phase deficit is depicted in Fig. 4 as d\u03d5 . Therefore, \u03c0\u03d5\u03d5 \u2212=+\u2212\u03a9 )(arg)( aNdl , (28) assuming that 0\u2265d\u03d5 and 0)(arg \u2265aN for SOSM. Now consider controllers that include a nonlinearity with infinite derivative in zero. For this type of nonlinearity, the DF \u221e\u2192)(aN if 0\u2192a and, therefore, 0)(1 \u2192\u2212 \u2212 aN if 0\u2192a . Also, assume that )(1 aN \u2212\u2212 is a straight line on the complex plane (other types of )(1 aN \u2212\u2212 will be considered below). Formulate the following theorem. Theorem 2. For the second-order linear part given by (1) and the controller containing at least one ideal relay function, and having the describing function )(aN of the controller such that the ratio const )(Re )(Im = aN aN (the negative reciprocal DF of the controller is a straight line on the complex plane), the following three modes of oscillations can occur" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000162_bf01972478-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000162_bf01972478-Figure1-1.png", + "caption": "Fig. 1. Arthrograph for metacarpophalangeal joint", + "texts": [ + " The original data were obtained by using an apparatus in which the sinusoidal motion was derived from that of a pendulum, and later by a crank driven by a variable speed motor which permitted investigation of a wider *) Paper presented to the British Society of Rheology Conference on Rheology in Medicine and Pharmacy, London, April 14 15, 1970. range of frequencies (from 0.03 3 cycles/second), velocities (15 radians/see minimum) and accelerations (280 radians/sec maximum). A portable version of this apparatus is shown in fig. 1. To contrast these results with those of a weight bearing joint, an arthrograph operating on similar principles was devised for the knee (fig. 2). The lower leg was attached to an adjustable holder, which could aecommodate any size of leg. It was strapped on so that no movement was permitted at the ankle joint during oscillation. Oscillations could be started from any position of flexion or extension. No tests were done in the final 20 30 ~ of full extension, due to the rotation of the lower leg which occurred at that amplitude" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001015_aas-007-0628-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001015_aas-007-0628-Figure1-1.png", + "caption": "Fig. 1 Triangle formation in formation coordinate system", + "texts": [ + " Note that, approaching to its trajectory, leader will also consider the movement characters of the whole formation. So the speed of group may be immolated, however, formation accuracy can be improved. And the speed of any robot can be considered invariable in a short time interval because it satisfies the piecewise-constant condition. In formation control, the relative poses of robots are the most important parameters. To express the relative poses parameters, a formation coordinate system needs to be constructed firstly (see Fig. 1). Once the linear speed vp(t) and angular speed \u03c9p(t) of leader Rp are determined, the curvature center O of the current instantaneous trajectory can be affirmed. Formation coordinate system is thus created with O as its origin and the direction of polar axis in global coordinates as its direction of polar axis. Note that it is an instantaneous polar coordinate system and may vary with the movement of the group. The instantaneous trajectory of Rp is called the reference trajectory, and the direction of trajectory is the reference direction in this formation coordinate system. Also, passing through its current position, each follower has a virtual trajectory that is parallel to the reference trajectory. This virtual trajectory is called its following trajectory. Formation shape can be described by the polar coordinates. For example, in this case, the coordinate matrix of the triangle formation depicted in Fig. 1 can be defined as M = 2 4 vp/\u03c9p \u03bbp \u2212 sign(\u03c9p)\u03c0/2 \u03c1f1 \u03b8f1 \u03c1f2 \u03b8f2 3 5 (1) where \u03bbp denotes the orientation of leader, \u03c1fi, \u03b8fi(i = 1, 2) are polar coordinates of followers in formation coordinate system. Though the formation shape may be defined by M uniquely, it is not a vivid presentation for the relative poses. To visualize the formation shape, the matrix M can be transformed into two relative-distance matrixes Dm\u00d7m = {dij}i,j=1,\u00b7\u00b7\u00b7 ,m and Lm\u00d7m = {lij}i,j=1,\u00b7\u00b7\u00b7 ,m. Here, D denotes an offset matrix and L is a spacing matrix" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003065_s12206-021-0607-z-Figure16-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003065_s12206-021-0607-z-Figure16-1.png", + "caption": "Fig. 16. Overlapping contact pattern on the tooth pair under \u2206rx = 0.05\u00b0, \u2206ry = 0.05\u00b0 and \u2206c = 0.", + "texts": [ + " The contact pattern between the lower tooth and the upper tooth of the S-shape surface gear was evenly distributed on the tooth flanks, as shown in Fig. 15. The above results indicated that the kinematic errors of the misalignment angles were the same for the positive and negative values. The maximum kinematic error occurred in cases 2 and 3. Results of the gear contact simulation confirmed that the contact pattern between the S-shaped surface tooth pair was evenly distributed on the tooth flanks. Based on Figs. 14(a)-(j) for case 6, Fig. 16 shows the overlapping contact pattern on the tooth pair, where the rotation angle of the pinion ranged from \u221214\u00b0 to 22\u00b0 every 12\u00b0. The contact area gradually increased and moved to the tooth flanks and then gradually decreased. All the cases showed tooth frank contact. The contact simulations and kinematic errors indicated that the kinematic errors of misalignment angles remained the same for both positive and negative signs. In this section, a solid model of the S-shaped surface gear was manufactured" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002948_tfuzz.2021.3089053-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002948_tfuzz.2021.3089053-Figure1-1.png", + "caption": "Fig. 1. (a) Membership functions on variable xl as input, (b) Membership functions on variable \u03b5l(xl) as output.", + "texts": [ + " In this section, the 2-FIS internal component (5) of the system (3) is designed with IF\u2013THEN fuzzy rules of the form [37]: Ri : IF xl is Mi,THEN \u03b5l is U li , l = 1, 2; (6) where x1 and x2 are the input fuzzy variables with the following membership functions: M l 0(xl) = 1 \u03a6l 1 xl + 1, if \u03a6l\u22121 \u2264 xl < 0 \u2212 1 \u03a6l 1 xl + 1, if 0 \u2264 xl \u2264 \u03a6l1 0, elsewhere, (7) M l \u22121(xl) = 1, if xl < \u03a6l\u22121 \u2212 1 \u03a6l 1 xl, if \u03a6l\u22121 \u2264 xl \u2264 0 0, if xl > 0, (8) M l 1(xl) = 0, if xl < 0 1 \u03a6l 1 xl, if 0 \u2264 xl \u2264 \u03a6l1 1, if xl > \u03a6l1, (9) where the symmetry \u03a6l\u22121 = \u2212\u03a6l1 is considered. At the output variable \u03b5l we consider singleton-type membership functions (see Fig. 1). The designed rule base matrix is given in Table I. Adopting a Mamdani-type fuzzy inference system, being the inference engine a product inference and the defuzzifier being the center average, the structure of the fuzzy system can be formulated as follows: \u03b5l(xl) = 1\u2211 i=\u22121 { M l i (xl)\u22111 r=\u22121M l r(xl) } U li = 1\u2211 i=\u22121 \u03a8l i(xl)U l i , (10) where U l\u2212i = \u2212U li refers the fired crisp value of the output regarding of x1(t) and x2(t). Moreover, the function \u03a8l i(xl) must satisfy the following conditions: Input xl Output \u03b5l(x) N N Z Z P P \u2022 Since two rules are activated at the same time [37] for \u03a6li\u22121 \u2264 xl \u2264 \u03a6li+1, then term \u2211p r=\u2212pM l r(xl) becomes one, that is: 1\u2211 r=\u22121 M l r(xl) = Mn(xl) +Mn+1(xl) = xl \u2212 \u03a6li+1 \u03a6li \u2212 \u03a6li+1 + xl \u2212 \u03a6ln \u03a6ln+1 \u2212 \u03a6ln = 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001762_chicc.2008.4605542-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001762_chicc.2008.4605542-Figure1-1.png", + "caption": "Fig. 1 The model for AUV", + "texts": [ + " The control performance of the closedloop system is guaranteed by appropriately choosing the design parameters. The remainder of the paper is organized as follows. In Section 2, we introduce the model of the AUV\u2019s diving system. An adaptive neural network controller is presented for controlling free-pitch-angle diving motion of an AUV in Section 3. Certain simulation studies performed on the system to show the effectiveness of the proposed control scheme in Section 4. Section 5 contains a brief conclusion on this paper. In this paper, we consider an AUV as shown in Fig.1 , which has one propeller, two stern planes and two rudders to control the vehicle. The dynamics of a six degree-of-freedom underwater vehicle can be described as [13, 14] M(v)v\u0307 +CD(v)+ g(\u03b7)+ d = \u03c4, \u03b7\u0307 = J(\u03b7)v (1) where \u03b7 = [x,y,z,\u03c6 ,\u03b8 ,\u03c8 ]T is the position and orientation vector in earth-fixed frame, v = [u,v,w, p,q,r]T the velocity and angular rate vector in body-fixed frame as shown in Fig.1 [13], M(v)v\u0307 \u2208 \u211c6\u00d76 the the inertia matrix (including added mass), CD(v) \u2208 \u211c6\u00d76 the matrix of Coriolis, centripetal and damping term, g(\u03b7)\u2208 \u211c6\u00d76 the gravitational forces and moments vector, d denotes the unstructured uncertainty vector, such as exogenous input term and unmodeled dynamics and \u03c4 is the input torque vector. And J(\u03b7) is the transformation matrix which is defined in . Underwater vehicles are generally designed to have symmetric structure; therefore, we assume that the body-fixed coordinate is located at the center of gravity with neutral buoyancy" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000847_s10409-008-0179-5-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000847_s10409-008-0179-5-Figure3-1.png", + "caption": "Fig. 3 The configuration of the internal force of the triangular lattice cell", + "texts": [ + " In the present analysis, the material of cell walls is assumed to be perfectly elastic-plastic. Because of the quantitative influence of the high order deformations upon the collapse of lattice material, the cell wall is simplified as a beam subjected to axial force, shear force and bending moment. 2.1 Mechanical analysis of the triangular cell Consider a regular periodic lattice with an equilateral triangle unit cell as shown in Fig. 2. Under general in-plane loading the forces within cell struts shown in Fig. 3 are denoted as axial force Ni , bending moment Mi and shearing force Qi , and the suffix \u201ci\u201d is the mark of the cell strut. Letters A, B, C denote the three apexes of the triangular cell and D is the middle point of the level strut. Consider the case of biaxial stresses \u03c3x and \u03c3y . Due to the symmetry of the unit cell, the internal forces satisfy the following equations: N1 = N2, Q1 = \u2212Q2, M1 = M2, M \u2032 1 = M2, M \u2032 3 = M3 = M3D = Q3 = 0. (1) Equivalence of internal forces in the lattice strut requires: M \u2032 1 = M1 + Q1l, (2) where l denotes the length of the cell strut" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003169_s00202-021-01352-z-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003169_s00202-021-01352-z-Figure7-1.png", + "caption": "Fig. 7 Flux density distribution of the STPMIG for a bridged structure, b bridgeless structure", + "texts": [ + " In order to investigate the flux of the stator teeth, the flux density of all the stator teeth is shown in Fig.\u00a06 for both the structures with and without iron bridges. The maximum flux density of the stator teeth for the bridged and bridgeless structures is 1.46\u00a0T and 1.56\u00a0T, respectively, which indicates a 7% increase in the outer flux to the stator. On the other hand, due to the number of generator poles, it is expected that there are no variations in the directions of the flux in two parts of the stator teeth, which is illustrated in Fig.\u00a06. 1 3 Figure\u00a07 shows the flux distribution of the proposed STPMIG for both the bridged and bridgeless structures. The bridgeless generator provides higher flux density than the bridged generator, leading to higher core loss. In order to survey the output voltage and the output power of the proposed generator, the simulation is performed by regarding the following assumptions: \u2022 The outer rotor speed equals the rated speed so that the output induction voltage of the generator equals its rated value. \u2022 The inner rotor operates under the no-load condition to be rotated synchronously with the outer rotor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000185_11802372_33-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000185_11802372_33-Figure1-1.png", + "caption": "Fig. 1. Model of the SMC rover system", + "texts": [ + " This paper investigates a planetary rover system called \u201cSMC rover\u201d, developed by Hirose, etc [3][4]. SMC consists of six wheels to support the main body, on which the main control system of SMC, solar battery cell, communication devices, tool changers for the child rovers and other equipments needed for the planetary exploring carried. Each child rover has a wheel for locomotion and an arm for manipulation. Each of the child rovers can be looked as an autonomous child robot, which is detachable from the main body, as shown in Fig.1. The main body can be regarded as a father robot. When a child robot is in connected mode, it acts as one of the wheels of the father robot. When it is in autonomous mode, it can disconnect automatically from the main body and then rove on the planetary ground to sample or manipulate, which called locomotion mode and manipulation mode. When a child robot blocked by an obstacle in locomotion mode, it will send messages to get help from other child robot or father robot. Two or more child robots can connect automatically to form a new robot by grasping another child robot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002874_s10846-021-01410-5-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002874_s10846-021-01410-5-Figure2-1.png", + "caption": "Fig. 2 a isometric view of the proposed benchmarking platform. b top and side views of the model, with measurements and approximate robot workspace annotated", + "texts": [ + " Though defining a grasp as either a pass or fail has been pragmatic for comparison, training, and assessment purposes, it is not clear whether this single-faceted criterion is related to grasp quality, handling quality or placement quality. Therefore, this work proposes new metrics in conjunction with a testing protocol and a self-contained benchmarking platform. Key aspects of design, calibration and other details related to the proposed benchmarking platform are covered in this section. Comprehensive documentation for construction is availa b l e a t : h t t p s : / / d r i v e . g o o g l e . c om / o p e n ? i d = 1VsEjCl6hrX3FeL9VRF-J9CzVL7JHXO15. The proposed benchmarking platform illustrated in Fig. 2a is portable and self-contained. The layout consists of a conveyor system, vision enclosure and robotic manipulator, framed by 45 \u00d7 45 mm slotted aluminium extrusion. Figure 1 illustrates an annotated diagram of the proposed system. The 460 \u00d7 430 \u00d7 420 mm (L \u00d7W \u00d7 H) vision enclosure employs HD webcams, diffuse LED lighting and a photoelectric through-beam sensor. The inside panels were matt white. 2 LED strips were mounted at the top of the enclosure, facing downward. A plastic diffusor was added at the top of the enclosure, below the lighting", + " A comprehensive specification list is tabulated in Table 3. Power, communication, and control subsystems were housed in a control box mounted to the apparatus. AC power was converted to appropriate DC voltages. 24 V was used to power the stepper motor driver and LEDs. 5 V was used to power the microcontroller. Details regarding the componentry within the control box are provided in Table 4. The complete system occupies a cuboid of approximately 1600 \u00d7 700 \u00d7 470 mm. Dimensions and workspace are illustrated in Fig. 2b. The total cost of a single benchmarking platform proposed in this paper is roughly $2300 USD. The robotic manipulator accommodates gripping locations of up to approximately 28 mm and objects weighting up to 500 g. During development and experimentation, the manipulator performed more than 10,000 grasp actions. Therefore, the conveyor completed over 20,000 translation actions. No breakages were encountered, and no significant wear is visible. No parts were replaced during this period, no screws were tightened, and no cables were frayed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003433_978-3-030-40667-7_6-Figure2.3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003433_978-3-030-40667-7_6-Figure2.3-1.png", + "caption": "Fig. 2.3 The six degrees of freedom: forward/back, up/ down, left/right, yaw, pitch, roll", + "texts": [ + " Ry q q q q q ( ) = - \u00e6 \u00e8 \u00e7 \u00e7 \u00e7 \u00f6 \u00f8 \u00f7 \u00f7 \u00f7 cos sin sin cos 0 0 1 0 0 Rx f f f f f ( ) = - \u00e6 \u00e8 \u00e7 \u00e7 \u00e7 \u00f6 \u00f8 \u00f7 \u00f7 \u00f7 1 0 0 0 0 cos sin sin cos The full rotation matrix of the Platform relative to the Base is then given by the vector product of the three rotational matrices: P x y zR R R RB = ( )\u00b4 ( )\u00b4 ( )f q y pRB = - + +cos cos sin cos cos sin sin sin sin cos sin cos sin y q y f y q f y f y q f y q y f y q f y f y q f q q cos cos cos sin sin sin cos sin sin sin cos sin cos + - + - sin cos cosf q f \u00e6 \u00e8 \u00e7 \u00e7 \u00e7 \u00f6 \u00f8 \u00f7 \u00f7 \u00f7 To define the degrees of freedom, or mobility, of the hexapod, it is schematized as a kinematic chain. It is a mathematical model for an assembly of rigid bodies used to find the number of parameters that define the configuration of the chain. A single object in the space, if it is completely free to move, has six degrees of freedom, three translational, and three rotational, so it means that it can reach any position with any orientation in space, as shown in Fig.\u00a02.3. When it is considered a system of n connected rigid bodies, it has M\u00a0=\u00a06n degrees of freedom, 2 Mathematics of\u00a0the\u00a0Hexapod 16 but it has also constraints due to the links that connect these bodies, which decrease the degrees of freedom of the system. The whole structure must be considered as another body of the system, in order to be independent by its position in the fixed frame. This means that N\u00a0=\u00a0n\u00a0+\u00a01 is the number of rigid bodies in the system. The number of constraints ci that a joint imposes in terms of the joint\u2019s degrees of freedom is ci\u00a0=\u00a06\u00a0\u2212\u00a0 fi, where fi is the degrees of freedom of the joint" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000636_bfb0119382-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000636_bfb0119382-Figure3-1.png", + "caption": "Figure 3. The 2 phases of translating an object with a ratchet motion. The robots are denoted by disks. Va denotes the velocity of the active robot. Vp denoted the velocity of the passive robot. Fp is the force felt through the rope by the passive robot.", + "texts": [ + " When Bonnie is the active robot, the object rotates clockwise; when Clyde is the active robot, the object rotates counterclockwise. In the ratcheting skill, two robots translate an object by coordinating movements of the rope between them. As in the flossing skill, this type of manipulation is also preceded by a binding operation. The manipulation system may also combine ratcheting manipulation with flossing manipulation. We have developed a control algorithm for this skill and implemented it using two RWI B14 robots (see Figure 3). The algorithm for translations iterates the following phases. In the first phase the active robot translates forward with velocity v~. The passive robot follows along feeling Fp and keeping the rope taut. This action causes the robot to translate. When this phase is executed for the first time, it terminates with the passive robot sensing contact with the object. This is called the calibration step. During a calibration step, the passive robot measures the approximate distance to the object. In subsequent repetitions of this phase, the passive robot uses dead reckoning to decide when to stop" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002886_iemdc47953.2021.9449529-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002886_iemdc47953.2021.9449529-Figure7-1.png", + "caption": "Fig. 7. Topology B with winding factors kmws = 0.88, khws = 0.72, khwf = 0.86: a) Geometry with auxiliary rotor winding (light brown); b) circuit diagram with passive diode rectifier on the rotor and no smoothing capacitor; c) main machine and harmonic machine stator MMF at same current density", + "texts": [ + " 5 In two-speed grid-connected induction motors, poleamplitude-modulation (PAM) windings are used to allow switching between different synchronous speeds which at fixed grid frequency fel corresponds to two pole-pair numbers pm and ph. The change between synchronous speeds is achieved by changing the winding configuration with the help of mechanical switches. While this switch configuration does not enable the simultaneous creation of both pm and ph fields, the concept of topology A can be transferred to PAM windings by feeding the coil groups which are normally switched by mechanical switches from different inverter legs (Fig. 7). The advantage of Authorized licensed use limited to: California State University Fresno. Downloaded on July 01,2021 at 02:36:18 UTC from IEEE Xplore. Restrictions apply. PAM windings is that arbitrary pole-pair combinations can be created, albeit with complicated and often asymmetrical phase MMFs for combinations other than pm = 2ph. In this work, the well-known pm = 3, ph = 2 two-layer PAM winding shown in Fig. 7c is used. Since for the rotor a combined winding similar to topology A is not feasible, a fractional slot auxiliary winding must be utilized. This allows for a more robust rotor structure due to the absence of active power electronics on the rotor, but also leads to significantly higher copper losses in the auxiliary winding [4] and hence to a reduced excitation system efficiency of \u03b7exc = 85 %. Rated operation for topology B is shown in Fig. 8 According to the harmonic field theory [5] the relative amplitude of the MMF created by harmonic v of an m-phase winding is given by (8)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000360_4-431-27901-6_2-Figure2.1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000360_4-431-27901-6_2-Figure2.1-1.png", + "caption": "Fig. 2.1. Tower\u2019s test rig. A, bearing bush; B, bearing cap", + "texts": [ + " Based on Tower\u2019s experiments, Osborn Reynolds (1842 \u2013 1912), the physicist, formulated a theory of lubrication in 1886 [3]. Since then, Reynolds\u2019 theory has been the foundation of the theory of hydrodynamic lubrication. Recently developed theories of elastohydrodynamic lubrication, thermohydrodynamic lubrication, turbulent hydrodynamic lubrication, and others are regarded as extensions of Reynolds\u2019 theory. The pioneering works of Tower and Reynolds are reflections of Britain\u2019s advanced technology at that time. In this chapter, Tower\u2019s experiment will be explained first and then Reynolds\u2019 theory will be derived. Figure 2.1, which is a simplification of a drawing from Tower\u2019s famous paper of 1883 [1], shows the main part of Tower\u2019s friction test rig for a bearing used in rolling stock. The bearing is a partial bearing, and bearing bush A covers the upper half of the journal. A load (weight of the vehicle) acts on the journal from above through bearing cap B and bearing bush A. The lower part of the journal is immersed in lubricating oil, and the oil adhering to the journal surface is pulled up by rotation of the journal and is supplied to the bearing clearance" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002859_s12046-020-01541-9-Figure11-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002859_s12046-020-01541-9-Figure11-1.png", + "caption": "Figure 11. Effect of section thickness on (a) yield strength and (b) tensile strength.", + "texts": [ + "3 Effect of section thickness on mechanical properties The mechanical properties of the materials largely depend on the micro-constituents and their volume fractions present in the casting. The variations of microstructure with the thickness of the sections will ultimately impact the mechanical properties like strength and hardness of the materials. It is well known that pearlite is a harder phase when compared with ferrite and graphite, and it imparts strength. The yield strength and tensile strength of different sections of stepbar casting are given in figure 11. It can be observed that the yield strength of the materials is in the range 248\u2013370 MPa. The yield strength of the 5-mm section is 260\u2013295 MPa, whereas the yield strength of 10 and 15 mm is in the range 235\u2013250 and 235\u2013250 MPa, respectively. Similarly the tensile strength of the 5-mm section is 330\u2013370 Mpa, whereas the yield strength of 10 and 15 mm is in the range 270\u2013290 and 220\u2013240 MPa, respectively. The Brinell hardness of stepbar casting is in the range 167\u2013227 BHN as shown in figure 12. It can be observed that the hardness of the 5-mm section is in the range 197\u2013227 BHN, whereas for 10 and 15 mm BHN is 180\u2013197 and 167\u2013180, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001034_ijvd.2008.021154-Figure15-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001034_ijvd.2008.021154-Figure15-1.png", + "caption": "Figure 15 Plate with protrusions imitating the finger piston in experimental tests", + "texts": [ + " During the tests cyclic engagements of the clutch are executed. Operating conditions in the tests were identical with those used in numerical simulations (Table 1). The three types of clutch packs were configured like those used in simulations, with six sliding interfaces in each of them. The apply plate, which plays an important role in the system, was included in each pack. The effect of the finger piston was obtained by inserting a special plate between the machine piston and the apply plate, shown in Figure 15, equipped with protrusions which imitate the fingers. Number of protrusions and their circumferential width were the same as those of the fingers in the clutch. For each of the three pack designs two tests were performed. Figure 16 shows a photograph of rubbing surfaces of metal disks from the original pack design taken after 50 engagement cycles. Five distinct hot spots, manifesting themselves by dark discolorations, were produced. The hot spots were present at all sliding interfaces across the pack, with fairly small differences in severity between the surfaces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002408_tmag.2021.3064023-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002408_tmag.2021.3064023-Figure2-1.png", + "caption": "Fig. 2. Configurations of different types of windings. (a) Lap winding. (b) Toroidal winding. (c) Concentrated winding.", + "texts": [ + " The 3-D diagram and cross-section view and direction of dc current injection of the proposed 6/2 HSVRM are shown in Fig. 1. 0018-9464 \u00a9 2021 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See https://www.ieee.org/publications/rights/index.html for more information. Authorized licensed use limited to: University of Prince Edward Island. Downloaded on June 03,2021 at 12:32:51 UTC from IEEE Xplore. Restrictions apply. However, the design of low-slot-pole HSVRM brings the shortcoming of overlong end length. Fig. 2 presents the configurations of three types of windings. In the machines that use lap winding, as presented in Fig. 2(a), the length of the end winding is very long and it can be even longer than the stack length. Toroidal winding shown in Fig. 2(b) in which the coil is wound around stator yoke is proposed to shorten the length of the useless end, but the utilization of the coil is very low [8]. As for the concentrated winding shown in Fig. 2(c), although the end length is shorter, a large eddy current loss will be caused as this winding type has more harmonic components. To deal with the mentioned problems, in this article, a novel winding structure, named half-quasi-squirrel cage winding, is proposed. As shown in Fig. 3, the main features of the proposed structure are as follows: one conductor is inserted into each stator slot to improve the utilization rate of slot, and connected by the end ring at one end. The other end is directly connected to the modular driving circuit" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000629_9780470612286.ch1-Figure1.2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000629_9780470612286.ch1-Figure1.2-1.png", + "caption": "Figure 1.2. Geometric parameters in the case of a simple open structure", + "texts": [ + " The transformation matrix from the frame Rj-1 to the frame Rj is expressed in terms of the following four geometric parameters: \u2013 \u03b1j: angle between axes zj-1 and zj corresponding to a rotation about xj-1; \u2013 dj: distance between zj-1 and zj along xj-1; \u2013 \u03b8j: angle between axes xj-1 and xj corresponding to a rotation about zj; \u2013 rj: distance between xj-1 and xj along zj. The joint coordinate qj associated to the jth joint is either \u03b8j or rj, depending on whether this joint is revolute or prismatic. It can be expressed by the relation: j j j j jq r= \u03c3 \u03b8 + \u03c3 [1.2] with: \u2013 \u03c3j = 0 if the joint is revolute; \u2013 \u03c3j = 1 if the joint is prismatic; \u2013 j\u03c3 = 1 \u2013 \u03c3j. The transformation matrix defining the frame Rj in the frame Rj-1 is obtained from Figure 1.2 by: j-1Tj = Rot(x, \u03b1j) Trans(x, dj) Rot(z, \u03b8j) Trans(z, rj) j j j j j j j j j j j j j j j j j C -S 0 d C S C C -S - r S S S S C C r C 0 0 0 1 \u03b8 \u03b8\u23a1 \u23a4 \u23a2 \u23a5 \u03b1 \u03b8 \u03b1 \u03b8 \u03b1 \u03b1\u23a2 \u23a5= \u23a2 \u23a5\u03b1 \u03b8 \u03b1 \u03b8 \u03b1 \u03b1\u23a2 \u23a5 \u23a2 \u23a5\u23a3 \u23a6 [1.3] where Rot(u, \u03b1) and Trans(u, d) are (4 \u00d7 4) homogenous matrices representing, respectively, a rotation \u03b1 about the axis u and a translation d along u. \u2013 for the definition of the reference frame R0, the simplest choice consists of taking R0 aligned with the frame R1 when q1 = 0, which indicates that z0 is along z1 and 0 1O O\u2261 when joint 1 is revolute, and z0 is along z1 and x0 is parallel to x1 when joint 1 is prismatic" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000812_icit.2008.4608539-Figure9-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000812_icit.2008.4608539-Figure9-1.png", + "caption": "Fig. 9. Minimal and maximal tool inclination angle", + "texts": [ + " In order to suppress these unfavorable phenomena, desirable range of the inclination angle is considered. Fig. 6. Dimples with deformation ( =30\u00b0) Suppose, the following cutting conditions are denoted as Diameter of cutting tool is R mm Clearance angle : Number of cutting blade is N Shape of the cutting tool: Ball-end mill Depth of cut is d mm Feed rate is c mm/tooth where is selected enough big value to suppress the collision phenomena shown in Fig.5(b). The geometrical relation requires that the inclination angle should be achieved in the region shown in Fig.9. P1 is defined as a point on the cutting edge which is cN/2 apart from the trajectory tool center point P0. Therefore, inclination angle should satisfy the following equation. 1 1 180sin cos 1 2 cN d R R (1) Based on the geometrical relation, the maximum inclination angle can be derived from the following equation. 1 180cos 1 2 d R (2) x y The experimental results in the section 2 are discussed based on a numerical model. Authors have proposed a numerical model to describe generation of dimples during the milling process [6],[7]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001278_s1068798x08110087-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001278_s1068798x08110087-Figure2-1.png", + "caption": "Fig. 2. Calculation of a threaded joint, taking account of the contact pliability.", + "texts": [ + " (3), we obtain the following expressions: the proportion \u03c7F of the external load due to the tensile force F at the screws is (4) the proportion \u03c7M of the external load due to the tipping moment M at the screws is (5) where \u03bbjo = k\u03b5/A, \u03bbP = (h1 + h2)/EA, and \u03bbs are the pliabilities of the joint, the parts, and the screw, respectively, i.e., the changes in their dimensions under the action of unit load; h1 and h2 are the thicknesses of the tensional flanges; I is the moment of inertia of the joint; ni is the number of screws at a distance xi from the neutral axis. The formula for the additional load Fsi (Fig. 2) at screw i takes the form (6) where FsF = \u03c7FF/n is the load on screw i due to the external tensile force; FsMi = \u03c7MMxi/\u03a3(ni ) is the load on screw i due to the external tipping moment. The angle \u03b8 of joint rotation may be written in the form (7) To determine the distance by which the parts move together in repeated loading and the loads on the screws, we employ experimental apparatus consisting of two plates, two rods, and a lever (Fig. 3). The supporting surface of the joint is flat. The contact surfaces of the rectangular plates (length l = 90 mm, width b = 120 mm, thickness h = 23 mm) are milled, with roughness Ra1 = 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002687_j.seta.2021.101240-Figure11-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002687_j.seta.2021.101240-Figure11-1.png", + "caption": "Fig. 11. Comparison of net magnetomotive force in basic PMDC motor and modified PMDC motor.", + "texts": [ + " The comparison shows that the basic PMDC motor cogging torque is very high and modified PMDC motor cogging torque is low. There are two poles are formed in PMDC motors. However, basic PMDC motor cross magnetization of the flux will increase due to the flux line arrangement but modified PMDC motor cross magnetization of the flux will be reduced with help of small air gap value and also reduced the net magneto intention strength. Comparison of net magnetomotive force R. Ullah Khan et al. Sustainable Energy Technologies and Assessments 46 (2021) 101240 in basic PMDC motor and modified PMDC motor is illustrated in Fig. 11. Relation between torque and armature speed is directly correlated to power and inversely related to force. Moreover, armature speed is denoted as \u03b3, power is represented as \u03c1 and the relation of this \u03c1 is expressed in Eq. (15), \u03c1 = Tq\u03b3 (15) Graphical representation of armature speed and torque is illustrated in Fig. 12. Armature speed is increased exponentially thus the torque was reduced linearly. Basic PMDC motor reduced cogging torque range is 0.85(n-m) and the modified PMDC motor cogging torque value 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001262_leon.2008.41.1.39-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001262_leon.2008.41.1.39-Figure5-1.png", + "caption": "Fig. 5. A model of the \u201coverbalancing hammer wheel.\u201d (\u00a9 Allan A. Mills)", + "texts": [ + " Casual inspection of the drawing might imply that the wheel will continuously rotate counterclockwise, but Leonardo is well aware that it will, in fact, soon assume a rectified and is copied here as Fig. 4. The arcs upon the face may represent initial attempts to construct the loci of the centers of gravity of hanging weights as the wheel turns, but they neither are symmetrical nor serve to position the hammers around the rim of the wheel rather than hanging vertically\u2014as shown once they leave the lowest point. Simple peglike stops are the simplest way of ensuring this, as shown in the model illustrated in Fig. 5. The most significant part of Fig. 4 is the way in which Leonardo has compounded the moments in mobile-like [8] chains to prove that, as drawn, the wheel is resting in equilibrium about an imaginary vertical line through its axis. Many examples of this diagrammatic construction are to be found in this codex, which has an emphasis on mechanics. For this model (Fig. 5), I fastened 12 brass weights to phenolic-fiber arms. The assemblies pivoted at equal intervals around a 20-cm-diameter disk. Protruding pegs were positioned near each arm so as to limit its swing. The resulting wheel was supported by a low-friction axle through its center and could be rotated manually in both clockwise and counterclockwise directions. Either way, it soon came to rest in equilibrium. As the weights were equal, it was expected\u2014and confirmed\u2014 that the sum of the distances of the centers of the five weights on the left from a central vertical was equal to the sum of the corresponding distances of the seven weights on the right" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002924_iemdc47953.2021.9449510-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002924_iemdc47953.2021.9449510-Figure2-1.png", + "caption": "Fig. 2. FEA models of 12-Ps-10-Pr DP-STVM for the isolated simulations of permeances and MMFs on the stator and rotor.", + "texts": [ + " According to [17], all the working field harmonics mentioned above have the same frequency in 6-slot-4-pole stator coil, therefore can both contribute to the no-load back EMF. The equivalent total airgap working flux density can be calculated and optimized through adjusting several size parameters, hence higher noload back EMF and torque performance can be expected. II. THE AIRGAP FLUX DENSITIES IN DP-STVMS Due to the irregular distribution of the stator teeth, the components of stator MMF as well as the stator permeance are changed with the FMT position ktp. As shown in Fig. 2 (a), by setting slotless iron rotor, the stator permeance and equivalent MMF varied with ktp can be isolated and simulated. The stator permeance and equivalent MMF under ktp=1, 1.2 and 1.4 are shown in Fig. 3, and the harmonic components of both parameters under ktp = 0.6-1.4 are simulated and presented in Fig. 4. It is obvious that when the distribution becomes uneven (ktp \u2260 1), besides original constant and iPs th permeance harmonics, several new permeance harmonics: (2j-1) Zth are introduced, same as the stator MMF components" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001288_dscc2008-2285-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001288_dscc2008-2285-Figure5-1.png", + "caption": "Fig. 5. Kinematic scheme of standard planetary gear.", + "texts": [ + " The resistance element is described by the generalized Stribeck curve shown in Fig. 4. The resistance element/friction curve is modulated by the clutch normal force that represents the active differential control variable. The zero-speed clutch friction torque (stiction) in Fig. 4 is not uniquely defined. This can be resolved by using different types of practical friction models [14,12], such as the classical friction model based on a steep straight line stiction curve, the Karnopp model, or dynamic friction models which incorporate stiction dynamics. Fig. 5 shows the kinematic scheme of the standard planetary gear (cf. Fig. 1d). Based on [7,8], the \"physical\" bond graph model of the planetary gear is shown in Fig. 6a. The analytically more straightforward, equivalent bond graph models are given in Figs. 6b and 6c (cf. [13]). The bond graph elements used are outlined in Appendix. 0 f f R ..f R 1 2 Fig 3. Bond graph model of clutch. plate speeds, Fig. 3 includes the speed difference 0 junction point 2 Copyright \u00a9 2008 by ASME Downloaded From: http://proceedings" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000238_mhs.2006.320274-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000238_mhs.2006.320274-Figure1-1.png", + "caption": "Fig. 1. Model of passive walker with knees", + "texts": [ + " In this paper, we focus on the stability mechanism of fixed point. It is thought that the essence of principle is common to any model. Therefore, we analyze the stability mechanism of fixed point from a simple passive walker with knees. At first, a global stabilization principle of fixed point is derived from Jacobian matrix of arbitrary state. Secondly, we demonstrate that the stabilization principle includes a generation principle of fixed point. Finally, the validity of the stabilization principle is confirmed by the walking experiment. Figure 1 shows model of passive walker with knees. The model consists of stance and swing legs. Knee of the stance leg is locked straight. The motion is assumed to be constrained to saggital plane. For the purpose of simplicity and clarity of analysis as possible, assumptions are given as follows: M m, M m1, M m2 (1) Stance leg is assumed to be fixed on the ground as no slippage or take off. The equation of leg-swing motion can be written as MK(\u03b8K)\u03b8\u0308K + HK(\u03b8K , \u03b8\u0307K) + GK(\u03b8K , \u03b3) = 0 (2) where MK(\u03b8K) = \u23a1 \u23a3 l2 \u2212(b1l + pll1) cos(\u03b8 \u2212 \u03c61) \u2212b2l cos(\u03b8 \u2212 \u03c62) 0 0 b2 1 + pl21 pb2l1 cos(\u03c61 \u2212 \u03c62) b2l1 cos(\u03c61 \u2212 \u03c62) b2 2 \u23a4 \u23a6 HK(\u03b8K , \u03b8\u0307K) =\u23a1 \u23a3 0 (b1l + pll1) sin(\u03b8 \u2212 \u03c61)\u03b8\u03072 + pb2l1 sin(\u03c61 \u2212 \u03c62)\u03c6\u03072 2 b2l sin(\u03b8 \u2212 \u03c62)\u03b8\u03072 \u2212 b2l1 sin(\u03c61 \u2212 \u03c62)\u03c6\u03072 1 \u23a4 \u23a6 GK(\u03b8K , \u03b3) = \u23a1 \u23a3 \u2212l sin(\u03b8 + \u03b3) (b1 + pl1) sin(\u03c61 + \u03b3) b2 sin(\u03c62 + \u03b3) \u23a4 \u23a6 g \u03b8K(= [\u03b8, \u03c61, \u03c62]T ) is the vector of joint angles", + " xk+1 = xf + Jf\u2206xk + J1\u2206xk \u00b7 \u00b7 \u00b7 +Jn\u22121\u2206xk (9) where Jf is Jacobian matrix of fixed point. J i is Jacobian matrix of xf + i\u2206xk. If all absolute of eigenvalues of Jacobian matrix of arbitrary state, |\u2206xk| > |Jf\u2206xk| and |\u2206xk| > |J i\u2206xk| holds. Consequently, Eq. (7) holds as follows: |xk+1 \u2212 xf | = |Jf\u2206xk + J1\u2206xk \u00b7 \u00b7 \u00b7 +Jn\u22121\u2206xk| \u2264 |Jf\u2206xk| + |J1\u2206xk| \u00b7 \u00b7 \u00b7 +|Jn\u22121\u2206xk| < |\u2206xk| + |\u2206xk| \u00b7 \u00b7 \u00b7 +|\u2206xk| = n|\u2206xk| = n \u2223\u2223\u2223\u2223\u2223xk \u2212 xf n \u2223\u2223\u2223\u2223\u2223 = |xk \u2212 xf | (10) Hence, fixed point is global asymptotically stable. In this section, we derive the Jacobian matrix of arbitrary state of model as shown in Fig. 1. In this study, the state just after heel\u2013strike x+ k is focused. State quantity of x+ k is given as \u03b1k and \u03b8\u0307+ k because \u03c6\u0307+ k is dependent variable of \u03b1, \u03b8\u0307+ k (see section IV-A). Successive states is related as x+ k+1 = f(x+ k ). Linear discrete-time state equation of \u2206x+ k can be derived as \u2206x+ k+1 = \u2202f \u2202x+ \u2223\u2223\u2223 x+=x+ k \u2206x+ k \u2261 Jxk \u2206x+ k (11) where Jxk = \u23a1 \u23a2\u23a2\u23a2\u23a3 \u2202\u03b1k+1 \u2202\u03b1k \u2223\u2223\u2223\u2223 k \u2202\u03b1k+1 \u2202\u03b8\u0307+ k \u2223\u2223\u2223\u2223 k \u2202\u03b8\u0307+ k+1 \u2202\u03b1k \u2223\u2223\u2223\u2223 k \u2202\u03b8\u0307+ k+1 \u2202\u03b8\u0307+ k \u2223\u2223\u2223\u2223 k \u23a4 \u23a5\u23a5\u23a5\u23a6 (12) From energy conservation law and Eq. (4), discrete-time state equation of \u03b8\u0307+2 k can be derived as \u03b8\u0307+2 k+1 = e2 k+1 ( \u03b8\u0307+2 k + 2g l { cos (\u03b1k 2 \u2212 \u03b3 ) \u2212 cos (\u03b1k+1 2 + \u03b3 )}) (13) where ek+1 = cos\u03b1k+1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002654_s42417-021-00299-6-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002654_s42417-021-00299-6-Figure5-1.png", + "caption": "Fig. 5 Layout structure of the system", + "texts": [ + " The analytical formulation of the comprehensive meshing stiffness can be obtained through a 3rd harmonic equation as Eq.\u00a0(11) considering the effect of the contact ratio of the gear pairs: (7)kh = EL 4 ( 1 \u2212 2 ) (8)kf = N\u2211 i=1 1 /( cos2 \ufffd 1 E\u0394l [ L\u2217 ( uf Sf )2 +M\u2217 ( uf Sf ) + P\u2217 ( 1 + Q\u2217 tan2 \ufffd 1 )]) (9) kn = 1 /( 1 kh + 1 kb1 + 1 ks1 + 1 ka1 + 1 kf1 + 1 kb2 + 1 ks2 + 1 ka2 + 1 kf2 ) (10)ksingle(t) = N\u2211 n=1 kn where km is the average meshing stiffness and kj1 , kj2 are the harmonic coefficients. Figure\u00a05 shows the layout structure of the helical gear system. Point A and B are the fixed support of the two input helical gears and point C is the center of the single output helical gear. The number of gear teeth ( ) in the layout angle\u2220ACB is not integer in most cases, which leads to the existence of the phase difference. can be divided into an integral part and a fractional part . The time difference of the stiffness excitation between the left and right driving gears can be calculated as \u0394t = 2 \u2215 ( z1 1 ) , and the phase difference of the meshing stiffness can be described as \u0394 = 2 e\u2215 ( z1 1 ) " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002701_s10556-021-00904-1-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002701_s10556-021-00904-1-Figure2-1.png", + "caption": "Fig. 2. Scheme of deformation of a moisture-containing waste layer in the range of the pressing machine drums: (1) moisturecontaining waste, (2) pressing drums, h0 is the thickness of the infeeding waste layer, h is the thickness of the waste layer after pressing, qi is the distribution curve of the load intensity on the material layer in the pressing zone along the length A1 \u2013 B2 (load intensity diagram), r is the radius of the pressing drums, \u03c9 is the angular velocity of the drums, \u0394y, z1 are the width and height of the selected elementary volume of the product, y1,2 in, y1,2 out are the current abscissas of the load intensity diagram at the input and output, \u03b2 is the angle of material layer nip by the drum.", + "texts": [ + " The efficiency of moisture release when pressing waste on drum pressing machines depends a lot on the size of the initially set gap between the drums; therefore, the task of determining the rational size of the gap depending on the intensity of the load on the drums is relevant. Let a moisture-containing waste layer with a width b equal to the length of the pressing drums and a thickness h0 with an average moisture content W and a bulk density \u03b31 (N/m3) be in the range of the pressing machine drums (Fig. 2). Let us select the elementary volume \u0394V = b\u0394yz from the layer of the material being pressed with the assumption that this volume is deformed only in the Oz plane without shear stresses caused by the shear of the material. The bulk density of the deformed waste layer at the inlet to the pressing drums (provided that the pressure changes within 1.2 \u2264 p \u2264 20.0 MPa) is expressed by an empirical dependence \u03b3 = 6800 10p 44 \u2212W 3 . (2) The abscissas of the load intensity diagram in the zone of the inlet to the pressing drums with drum curvature \u03c1 = 1/r can be expressed as y1in = 1 \u03c1 T 100b\u03b3 i \u239b \u239d\u239c \u239e \u23a0\u239f \u2212 h0 \u23a1 \u23a3\u23a2 \u23a4 \u23a6\u23a5 \u23a7 \u23a8 \u23a9 \u23ab \u23ac \u23ad 0.5 , (3) where T is the initial linear density of the material being pressed located on the feeding conveyor belt, kg/m; \u03b3i is the current value of the bulk density in the material layer cross-sections in the inlet zone along the length yi . The intensity of the load on the layer of the material being pressed at the inlet to the pressing drums (N/m) is determined as: \u2013 for length A1A2 (see Fig. 2): qA1A2 in = T 100b(\u03c1y2 + h0 )(2070 \u2212 \u03b31)kWin\u23a1\u23a3 \u23a4\u23a6 \u2212 \u03b31 2700 \u2212 \u03b31 , (4) where kWin is the coefficient that takes into account the moisture content of the material at the inlet, \u03b31 is the bulk density of material at the inlet for length A1A2, y is the length of the abscissa of the load intensity diagram at the inlet, y = A1O = y1in; \u2013 for length A2O: qA2O in = T (44 \u2212W )[ ] 68 \u22c5104b(\u03c1y2 + h0 )\u23a1\u23a3 \u23a4\u23a6 \u23a7 \u23a8 \u23aa \u23a9\u23aa \u23ab \u23ac \u23aa \u23ad\u23aa 6 . (5) Abscissas of the load intensity diagram in the outlet zone (after the line of the drums centers) in the direction of material movement are y1out = 1 \u03c1 T (100b\u03b3 i )\u2212 h \u239b \u239d\u239c \u239e \u23a0\u239f \u23a7 \u23a8 \u23a9 \u23ab \u23ac \u23ad 0", + " When pressing a moisture-containing material in the zone of impact of the drums, the layer will be compressed and spread over the width of the drums; therefore, to describe the motion of the layer, the well-known equation of motion of the material between two plates can be used: \u2202 \u2202x hk 3 \u2202p \u2202x \u239b \u239d \u239e \u23a0 + \u2202 \u2202y hk 3 \u2202p \u2202y \u239b \u239d\u239c \u239e \u23a0\u239f = 6\u03b7 2vc + \u2202hk \u2202x (v1 + v2 ) \u23a1 \u23a3\u23a2 \u23a4 \u23a6\u23a5 , (9) where hk is the height of the layer in the range of the drums, m, vc is the velocity of layer compression, m/s, v1, v2 are the circumferential velocities of the pressing drums, m/s. The height of the material layer in the range of the drums decreases gradually from the inlet of the material to the range to the outlet at a rate of vc = v sin\u03b2 , where v = (v1 + v2 )/2 , sin \u03b2 = \u0445/r (see Fig. 2). Since the height of the layer deformation zone is significantly lower than the radius of the drums, it is assumed that \u2202hk/x \u2248 x/r . As a result of transforming Eq. (9), taking into account the fact that the height of the layer hk along the axis of the drums Ox does not change for any value of the length of the segment xi, 0 \u2264 xi \u2264 b, where b is the length of the drums along the Ox axis, the Poisson equation is obtained: \u2202 \u2202x hk 3 \u2202p \u2202x \u239b \u239d \u239e \u23a0 + hk 3 \u22022p \u2202y2 = 24\u03b7 \u2202hk \u2202x v . (10) The solution of this equation under the boundary conditions p = 0, hk = h0 and for \u2202 2p/\u2202y2 = 0 has the form p = 48\u03b7v (h01 \u2212 hk1)(hk1 \u2212 h1) lh0hkh , (11) where h01, hk1, h1 are the halves of the layer height at the inlet to the range of the drums, in the range, at the outlet of the range, respectively, l = 2(h0 \u2212 h)/r " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002116_s12541-020-00407-8-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002116_s12541-020-00407-8-Figure1-1.png", + "caption": "Fig. 1 Schematics of the threadless screw", + "texts": [ + " Four different methods are proposed to overcome the shaft deformation due to the reaction force. Furthermore, Types of the mechanism implementation and sources of the lead errors are presented. Two configurations of mechanism are designed, fabricated and tested. Results show the reliability of the idea and good linearity. Moreover, results indicate that thrust force can change the lead considerably. In threadless screw, linear movement is caused by rotation of a roller pushed toward a shaft as shown in Fig.\u00a01, where the roller\u2019s axis and the shaft axis are of skew lines i.e. the angle between the axis of the roller and the axis of the shaft (lead angle) is not zero (\u03b1 \u2260 0). In this case, the roller moves in axial direction. The mechanism of motion is similar to a feed screw. The roller may be an external cylinder, or an internal hole. When \u03b1 = 0, no linear motion is produced and the only result is pure rotation of the roller. By precise adjustment of \u03b1, this mechanism can generate movements with very small lead which is not possible by other similar mechanisms such as ball screws or roller screws", + " On the other hand, this mechanism has the following drawbacks: (a) The transmitted force is limited to friction force between roller and shaft. (b) The lead may change due to slippage. (c) There is a single contact point between roller and shaft, and stress concentration in this point is undesirable. (d) The lead is not a fixed number; therefore, the open loop control is not reliable. Some applications and proposals for this mechanism are as follows: (a) Trolley The trolley may consist of a roller and moves over a rotating shaft such that the roller comes into contact with the shaft as shown in Fig.\u00a01. This trolley could move in both directions with different speeds depending on the angle between axes of the roller and bar. (b) Pipe feeder In this mechanism, a roller touches the pipe and its rotation causes the pipe to rotate and move axially at the same time. This mechanism may be used in the procedure of pipe production. (c) Threadless screw with disk rollers It consists of few rollers (ball bearings) which are forced towards the driver shaft while all the roller\u2019s lead angles are the same", + " The lead of the mechanism ( \u0394x ), is the linear displacement during one revolution of the shaft and is related to the angle between axes of the roller and shaft, , and the diameter of the shaft, d , as follows: According to Eq.\u00a0(1), for = 0 , the lead is zero, and for positive or negative , the movement directions are opposite and related to the magnitude of . (1)\u0394x = d tan In this mechanism, the linear motion is generated through the friction between the roller and the shaft. As the coefficient of friction is of the order of 0.1, the preload between roller and shaft, ( F in Fig.\u00a01), should be about 10 times bigger than the transmittable force. Therefore, a considerable force should be presented at contact point. This applies a bending moment and a crush stress to the shaft. Four different methods are proposed to overcome the bending moment as follows: (a) Direct method using additional neutral rollers as shown in Fig.\u00a05a: This is a simple mechanism, but the zero angle and zero lead for additional rollers cannot be guaranteed. On the other hand, the reaction of roller force, is not the same all over the length of shaft, it is bigger on supporting rollers, and less between two supporting rollers", + " Once the angle is computed, coefficients and can be calculated from the table provide in [29] and then the long diameter and short diameter of the contact area ( 2a and 2b ) and the maximum pressure at contact surface ( P max ) can be calculated using the following equations: (2)2A = 1 r 1 + 1 r \ufffd 1 + 1 r 2 + 1 r \ufffd 2 (3) 2B = \u221a ( 1 r 1 \u2212 1 r \ufffd 1 )2 + ( 1 r 2 \u2212 1 r \ufffd 2 )2 + 2 ( 1 r 1 \u2212 1 r \ufffd 1 )( 1 r 2 \u2212 1 r \ufffd 2 ) cos (2 ) (4)a = \u22c5 3 \u221a \u221a \u221a \u221a 3 4 P A ( 1 \u2212 2 1 E 1 + 1 \u2212 2 2 E 2 ) (5)b = \u22c5 3 \u221a \u221a \u221a \u221a 3 4 P A ( 1 \u2212 2 1 E 1 + 1 \u2212 2 2 E 2 ) 1 3 In the case of the threadless roller positioner presented in this paper as shown in Fig.\u00a01, the driver shaft is cylindrical which leads to r\u2032 2 =\u221e or 1 r \ufffd 2 = 0 . Knowing the maximum contact pressure between roller and shaft, suitable diameters for roller and shaft can be achieved. (6)P max = 3 2 P ab In this research, the material of both shaft and roller are carbon steel with Young modulus of elasticity of E = 206\u00a0Gpa and Poisson ration of v = 0.3 . For a lead of 5\u00a0mm, the following three rollers are considered according to Fig.\u00a07: N0.1: r1 = d/2, r\u2032 1 = 2\u00a0mm. N0.2: r1 = d/2, r\u2032 1 = 5\u00a0mm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002276_s12206-021-0114-2-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002276_s12206-021-0114-2-Figure3-1.png", + "caption": "Fig. 3. Cross-section of API standard line pipe connection.", + "texts": [], + "surrounding_texts": [ + "The analysis process is summarized in Secs. 2 and 3. The 2- D axisymmetric model was used to analyze the premium con- nection system owing to its shorter analysis time and simplistic nature in comparison to a 3-D model. For tensile and compressive forces, the analysis results of the 3-D model and 2-D axisymmetric model were in agreement. The 3-D modeling-based analysis was required to analyze situations where bending occurs [13]. Fig. 5 shows the geometry of an API standard premium connection. Fig. 6 describes the parameters applied to the box. The box length is 234.95 mm, with an outer and inner diameter of 153.67 mm and 143.26 mm, respectively. In Fig. 7, the box thread is magnified to illustrate the detailed geometric parameters of the thread. The thread thickness is 4.72 mm with a space length of 7.58 mm. The pitch is 15.25 mm and the thread length is 7.04 mm. Fig. 8 describes the parameters related to the pin. The outside diameter and thickness of the pin is 139.7 mm and 10.54 mm, respectively. Fig. 9 shows a magnified image of the pin thread and summarizes the thread geometry parameters. The space length, pitch, and thread length are 7.25 mm, 15.25 mm, and 7.97 mm, respectively. Fig. 10 describes the parameters related to the thread. The upper (stabbing flank) and lower corner radius (load flank) are both 0.2 mm." + ] + }, + { + "image_filename": "designv11_83_0001965_6.2007-864-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001965_6.2007-864-Figure8-1.png", + "caption": "Figure 8: Force Situation \u2013 Horizontal Plane Maneuvering Priority with \u03b3 Hold Constraint", + "texts": [ + " The computation of the remaining maneuverability in the vertical plane, MINX ,\u03b3& and MAXX ,\u03b3& remains the same, i.e. is given by equations (34) and (35), independent of the fact whether horizontal maneuvering gets priority over vertical maneuvering AND climb-angle hold according to equation (36), or whether horizontal maneuvering gets maneuvering priority with the boundary condition that the actual climb angle is to be maintained according to equation (37). Starting from equations (34), the constraints to be fulfilled to fly in a \u201chorizontal plane maneuvering priority mode\u201d have been presented. Figure 8 illustrates the situation for the case that horizontal maneuvering is prioritized with the boundary condition of climb angle hold. As a straightforward consequence of the previous chapter, where a controlled prioritization between speed/acceleration and climb/descend angle was introduced by means of cross-feeds between the speed and climb-angle reference models based on energy principles, cross-feeds can now be introduced between the climb-angle and course angle reference models to allow a controlled prioritization between horizontal and vertical plane maneuverability based on force principles" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001609_indcon.2008.4768808-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001609_indcon.2008.4768808-Figure3-1.png", + "caption": "Fig. 3. Workspace of the wrist: Only first three joints are needed for it.", + "texts": [ + " Another fact is that the control system now a-days are completely implemented on computers and the same computer can better be utilized as command console as well. Further, some complex maneuvers may need to change only the orientation of the end-effector while maintaining its position. For example, if we want to pour water from a jug to a cup, we need to continuously change the orientation of the jug while keeping the location of the jug at the same place i.e. at the top of the jug. A GUI can provide facilities to do such tasks interactively. The workspace of an articulated manipulator is shown in Fig.3 with its DH parameters in Table I. It shows the zones where we can place the wrist. Obviously, it is not easy to guess the coordinates that can be quickly fed to the computer and inverse kinematics solutions are happily found. This is because the neighborhood locations of a successful coordinate point cannot be assumed to be reachable! What complicates the problem further is that it may also be necessary to keep the orientation of the end-effector at a particular configuration. For example, if the end-effector is supposed to handle a test-tube containing some liquid, it must always remain vertical" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000812_icit.2008.4608539-Figure12-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000812_icit.2008.4608539-Figure12-1.png", + "caption": "Fig. 12. Intersections of metal and trajectory surface", + "texts": [ + " Discussion on experiment 2 Deformation of the dimples mentioned in section 2 can be explained by focusing on the clearance surface of the milling tool. Fig.11 shows the clearance surface of the tool. Lines L1 through L5 are parallel to the cutting edge and equally spaced from neighboring lines. The distance between Line Ln and the cutting edge is 0.1R\u00d7n. Using the numerical model, the three-dimensional trajectory surfaces of the cutting edge and lines L1 through, L5 are calculated in two cases, =15\u00b0 and =30\u00b0. The intersections of the trajectory surfaces and the workpiece surface are obtained as shown in Fig.12. These simulation data show that cutting with a tool of clearance angle =15\u00b0 causes collisions with the workpiece surface. Furthermore, the area where the collision may occur is close to the location of deformation observed in Fig.5. B. Discussion on experiment 3 Intersection of three-dimensional trajectory surface of cutting edge and the metal surface is calculated with the tool inclination angle =20\u00b0. The intersections are shown in Fig.13. The results mean that the movements of the cutting edge cause the similar deformation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001650_pes.2007.386081-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001650_pes.2007.386081-Figure5-1.png", + "caption": "Fig. 5. Vector diagram of internal fault and external fault in forward direction at the sending end", + "texts": [ + " In a word, on the premise of the blocking condition of inphase area and sacrificing some resistive tolerance, the fault component distance relay based on pre-fault compensated voltage can either ensure security or improve the capability against fault resistance. There is no question that the option of the compensated coefficient affects the performance of the fault component impedance relay, but it is difficult to exactly calculate opk , then taking the sending end as example, qualitative analysis is just made in the following text. To simplify analysis, assume that each impedance angle in the two-source system is completely same. The voltage phasor diagrams under internal fault and external fault for the sending end are shown in Fig.5 respectively, where F is fault point, the end of the compensated voltage locates at point P. C3 corresponds to (4) or (6). For the sending end during internal fault, point P moves towards to the source voltage along with the fault point close to the point Y at the end of the reach setting as shown in Fig.5, or point P anti-clockwise shifts along with the increase of fault resistance. This means that the compensated voltage moves close to the inside of C3 even falls into C3, and then the impedance relay may be failure to trip. Therefore, lessening opk is of advantage to enhance sensitivity and improve the relay capability against fault resistance. On the other hand, for the sending end under external fault with smaller resistance in forward direction, point P moves far away from the source voltage along with the fault point F close to the point Y at the end of the reach setting, this means that the compensated voltage moves close to the outside of C3 even exceeds C3, and then the impedance relay mal-operation may take place, so it is necessary to increase opk to improve security" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000882_s1061920808040109-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000882_s1061920808040109-Figure4-1.png", + "caption": "Fig. 4. A nongeneric linkage and its small perturbation.", + "texts": [ + " When it is, the formalism no longer works, but this is easily remedied, for instance, in the motions described by the two tangent bundles in Fig. 3 above, it suffices to identify the two copies of (say) the point M2 in the two bundles, but only in the cse when the velocity at that point is zero. In this section, we study three topologically different examples of nongeneric quadrangles. 5.1. The first quadrangular linkage ABCD that we consider here is specified by the data \u30081; 3/4, 1/2, 3/4\u3009. The position of this linkage is entirely determined by the position of the midpoint M of the rod BC (Fig. 4(a)). The figure shows two positions of the linkage A,B,C,D and A,B\u2032, C \u2032,D, determined by the points M and M \u2032. The geometric locus of the point M is a smooth curve except at the origin, where it has a self tangency. When M is at the origin, all three of the mobile bars lie on the same RUSSIAN JOURNAL OF MATHEMATICAL PHYSICS Vol. 15 No. 4 2008 line (the x-axis). The curve also possesses four collineation points (M1, M2, M3, and M4), where two of the mobile bars are aligned. Note that the singular point O is also a collineation point, and in a sense, a double one: from it, the system can pass to four different regimes of motion (two of them occur when M \u201cbounces back\u201d from O)", + " As before, the locus of the points M is the projection on R 2 under (2) of the canonical configuration space (curve) C0 \u2282 R 4, but now this curve is a singular algebraic curve with self tangency point at the origin. Our goal in this section is to describe the notion of velocity field in the case of a singular configuration spaces of quadrangles. To simplify the exposition, we will ignore the collineation points by assuming that they are all pass points. From the point of view of real-life mechanical systems, this means we are assuming that the rods have a lot of inertia (being heavy) and/or the hinges have no \u201cbounce\u201d to them. In Fig. 4(b), we show a small perturbation (obtained by decreasing l2) of the linkage under consideration: the two \u201chalves\u201d of the curve have fallen apart, and we obtain a generic linkage for which C consists of two disjoint closed curves. Note that if we increase l2 instead of decreasing it, we will obtain a generic quadrangle with a connected configuration space looking like the one in Fig. 1. 5.2. The second quadrangle that we consider here is specified by the data \u30081; 5/8, 7/8, 5/2\u3009. The position of this linkage is entirely determined by the position of the midpoint M of the bar BC (Fig", + " A smooth vector field will be a (smooth) section of an appropriately defined desingularized tangent bundle of the configuration space of our system (the latter, not being a smooth manifold, has no tangent bundle in the ordinary sense). Roughly speaking, this bundle is obtained by desingularizing the self tangency points in accordance with the regime of motion through the singular point. 6.2. To be definite, we consider the particular case of the nongeneric quadrangle L discussed in Subsection 5.1 and pictured in Fig. 4. Denote by C the locus of the midpoint M of the mobile bar BC; one can see in Fig. 4 that C has a self-tangency point at the origin. The curve C is in one-to-one correspondence with the canonical configuration space C0(L) via the projection (2), and we shall think of C as being the configuration space of our linkage. Let us denote by U a neighborhood of the singular point O \u2208 C, by U+ its intersection with the right coordinate half plane {x 0}, and by U\u2212 its intersection with the left coordinate half plane {x 0}; we further denote by U+ + the upper half of U+ (the one lying in the half plane {y 0}), and use (similar) notation (U\u2212 + , U+ \u2212 , U\u2212 \u2212 ) for the other three half neighborhoods of O in C" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001125_aero.2007.352851-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001125_aero.2007.352851-Figure1-1.png", + "caption": "Figure 1. A 6 DOF dynamic model of a small helicopter", + "texts": [ + " In the second scheme, a helicopter maintains its position in the formation by maintaining specified distances from two neighbouring helicopters. The proposed control schemes only use the state information of the neighbouring helicopters. The sliding mode method is used. It is shown that the relative distances and orientations of the helicopters are stabilized even in the presence of parameter uncertainty and wind disturbance. Numerical simulations are presented to demonstrate the efficiency of these techniques. This sections presents the dynamic model of a helicopter, shown in Fig. 1. Two frames are defined for this dynamic model; the inertial frame {0}, and the helicopter body frame {B} with an origin at the centre of mass. Six degrees of freedom are assumed for each helicopter. The three translational degrees of freedom, the surge, sway, and bounce of the centre of mass, are expressed in the inertial frame {O} and are denoted by (x(@), y(\u00b0), z(\u00b0)). The three rotational degrees of freedom are represented by the yaw-pitch-roll (ZYX) Euler angles (b, 0, r). These Euler angles define the orientation of the body frame with respect to the inertial frame through the following transformation matrix: CO CO ROB = SO C0 L so (- SO CO ( SO SO + CO SO S) + C SO CO) ( CO CO (- CO SO + SO SO S) + S SO CO) COSO COcbC (1) where c = cos and s = sin" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003052_icet51757.2021.9451096-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003052_icet51757.2021.9451096-Figure3-1.png", + "caption": "Fig. 3. Sigma-Pi neural network.", + "texts": [ + " The linear dynamic inversion controller is designed as d ad\u0394 B u Ax u u where \u0394B u is the pseudo control for control allocation; adu is the adaptive output of the neural network, which is designed in the next section; du is the output of linear ProportionIntegration controller as follows: d c c cI Pdt 9u x K x x K x x % where , ,I Ip Iq Irdiag K K K K and , ,P Pp Pq Prdiag K K K K are chosen according to the desired response performance; c c c c[ , , ]Tx p q r and c c c c[ , , ]Tx p q r are desired angular rate vector and derivative of desired angular rate vector respectively, which are generated by a second order command-shaping filter. Define c dt 9x x x , then the error dynamics of the fast dynamics can be expressed as adP I x K x K x u 4 & Define , TT T+ ,. /e x x , where 3 3 3 3 3 3 3 30 0 T T I P e eA I K K B I , ad ad ad ad T p p q q r ru u u+ , 4 4 4. /u \u0394 and (15) can be rewritten as ad e ee A e B u \u0394 2 To compensate model errors \u0394 in (16), a Sigma-Pi neural network is designed, which is described in Fig. 3. The output of the network takes the form ad ad ad ad 256 256 256 1 2 3 1 1 1 [ ] \u02c6 \u02c6 \u02c6 = \u02c6 T p q r T j j j j j j j j j T u u u w x w x w x t X ; + , 0 1 . / < < < u C C C W 3 where 256 3\u02c6 =RW is the interconnection weight matrices; 256 1\u02c6; =X W is the basis function vector built as 1 2 3 4 , , , ,t t kron x u C C C C . To form the basis function vector, the input data is preprocessed and normalized between -1 and 1 with 0.1 0.1( ) (1 ) / (1 )x xx e e; . To tackle the fast time-variant uncertainties, the weightupdate law of the neural network is designed based on derivative-free adaptive control as follows: 1 2 \u02c6 \u02c6 Tt t t > eW W x e PB 5 where \u02c6 tW is an estimate of tW , 1 1 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000089_cca.2004.1387490-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000089_cca.2004.1387490-Figure3-1.png", + "caption": "Fig. 3. NF based supelvisory control ofthe robot-sensor system", + "texts": [ + " Two of the external sensors are integrated in the 2 robot arms and two are mounted on a pan-tilt unit attached at the ceiling (denoted as sensor bead). The sensor head is equipped with a 3D stereo camera and a stereo acoustic sensor (microphone array). The sensors integrated into the robot arms are an optical 3D sensor integrated in the gripper as well as a forcetorque sensor mounted between wis t and gripper. In order to demonstrate the functionality of the NF supervisory control concept a benchmark experiment shown in Fig. 3 has been selected. A planar circular motion of the robot tool center point (TCP) will he considered. The robot motion will he disturbed by two-succeeding fault events. At the time t = 1.5 sec in the servo-controlled joint 4 an actuator bias will occur. Moreover, at t = 5.6 sec a collision will occur when the TCP trajectory will intersect a brick like obstacle. The corresponding signal responses ofthe joint positions I, 4,? and the corresponding joint torques are shown in Fig. 4. Obviously by naked eye it is very difficult clearly to detect (not to mention to classify) the malfunction in joint 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002969_j.matpr.2021.05.300-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002969_j.matpr.2021.05.300-Figure2-1.png", + "caption": "Fig. 2. Al alloy clutch plate thermal boundary condition.", + "texts": [ + " Calculation of stresses, and greatest weights utilizing EulerLagrange conditions. Developing 3D model of MPCs utilizing Solid edge programming. Exporting the model of MPCs from strong edge to ANSYS 19.2 work bench for FEA. Static and dynamic examination of MPCs utilizing ANSYS work seat at various working condition. Determining the wear, most extreme disfigurement and equal worries for every material utilizing FEA Showing the connection of disfigurement and obstruction property of every material utilizing figures and tables (Fig. 1. Fig. 2. Fig. 3. Fig. 4. Fig. 5. Fig. 6. Fig. 7. Fig. 8. Fig. 9. Fig. 10. Fig. 11. Fig. 12. Table 1). Comparing and conversation of results for choice of better covering materials. Fig. 5. Cast iron clutch plate heat flux results. Fig. 6. Cast iron clutch plate heat flux results. Grip disappointment and harm because of extreme frictional warmth and warmth changes to the grasp counter mate circle frequently happens to a car grips. This circumstance adds to warm weakness to the segment which causes the grasp counter mate plate to break and twist" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003226_dac.4926-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003226_dac.4926-Figure7-1.png", + "caption": "FIGURE 7 Energy level, Neighbors count vs \u03b1", + "texts": [], + "surrounding_texts": [ + "If xi equivalents to form metrics and ei designates moderate forecast failure which is optimization goal, the MLR may be calculated as yi \u00bc b0\u00feb1xi1\u00feb2xi2\u00fe\u2026\u00febnxin\u00fe ei: \u00f013\u00de Various quality computations can be utilized to decide the achievement and validity of the recommended alphabased mobility model for a fleet of UAVs or UASs. This is carried out by error inspection based on statistics. To assess and differentiate the attainment and correctness of the recommended framework with the standard planned data, the ultimate usual analytical standard estimates. RMSE and R2 have been utilized by applying regression in Matlab 16b. Statistical data analysis for T1FLS and T2FLS is shown (Table 2). In accordance with inputs and outputs data, we have drawn up an alpha-based mobility model for WSN in fuzzy setting. Regarding the empirical data, the model outcome is perceived value. So as to manage the followship weight attainment of the predictive model, we have determined a number of statistical metrics. The forecast competence of TABLE 1 Statistical data analysis of output using ANFIS and MLR Model Approach RMSE MAE MAPE R2 MLR 0.05613 0.0792 0.988 0.9624 ANFIS 0.05387 0.0753 0.977 0.9735 FIGURE 6 IF-THEN rule for the proposed WSN model suggested model is assessed by applying the testing data in the skilled data and differentiating the outputs and calculated values. From the Figures 7\u201310, it is observed that the actual results and the results from T2F technique are very close to each other. The change of \u03b1 is given in Figure 10 which is expected." + ] + }, + { + "image_filename": "designv11_83_0000009_fuzz.2003.1206625-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000009_fuzz.2003.1206625-Figure2-1.png", + "caption": "Fig. 2. The inverted pendulum system", + "texts": [ + " The minimum approximation 2 - <--e\"'QZ>>>< >>>>: \u00f04\u00de For any rotational axis vec with rotational angle u, the rotational matrix R could be written as: R( u)= \u00bd1 cos( u) V2 + sin( u)V \u00f05\u00de where V= 0 vecz vecy vecz 0 vecx vecy vecx 0 2 4 3 5 and vec= \u00bdvecx vecy vecz T. Rotational matrix R1 or R2 corresponding to vec1 or vec2 could be calculated by equation (5). Finally, the pure nodal deformation of the element is obtained (as shown in Figure 1(d)): qi = u0i u000i (i= a, b, c, :::g, h) \u00f06\u00de in which ui 000 represents deformation coordinates of node i. Establishing of dynamic damping model of gear dynamics The traditional damping model defined in equation (1), which is proportional to the magnitude of velocity with opposite directions, is not applicable for dynamics analysis because: (1) damping coefficient is usually less than 0.1 and low damping coefficient could reduce the rate of convergence greatly which multiply the computing cost to unacceptable levels; (2) using the larger damping coefficient to accelerate the the convergent process would result in distortion of computational solution due to over-large damping forces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000195_b978-0-443-07113-3.50014-7-Figure10.1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000195_b978-0-443-07113-3.50014-7-Figure10.1-1.png", + "caption": "Figure 10.1 The effect of a flare on the heel designed to convert a deforming valgus (F f \u00d7 d) moment to a correcting varus moment (F wf \u00d7 c).", + "texts": [ + " A simple example of this is the way in which a shoe can be modified to help to correct a pronated foot. With an unmodified shoe, the line of action of the ground reaction force will be lateral to the line of action of the weight (force) at heel strike, thus producing a valgus moment, which forces the subtalar joint into further pronation. This can be avoided by moving the point of initial contact medially so that the ground reaction force now produces a correcting varus moment. This can be achieved by flaring the heel of the shoe medially, as shown in Figure 10.1. Such a flare will not, however, have any effect upon the final position of the calcaneus relative to the ground, and this is required if pronation is to be controlled. The simplest way to adjust the position of the calcaneus would be to use a heel wedge, either placed in the heel of the shoe or, more effectively, built into an in-shoe orthosis (Fig. 10.2). Wedging, also known as posting, is one of the principles used in functional foot orthoses; the exact angle at which the rearfoot is canted forms part of the assessment of the patient" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003207_eit51626.2021.9491871-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003207_eit51626.2021.9491871-Figure8-1.png", + "caption": "Fig. 8. Experimental setup for CWRU bearing dataset [20].", + "texts": [ + " 7 displays the proposed LSTM frame. As shown in Fig. 7, 1-D feature vector with a size of 4096 is employed as the input for the LSTM. Notice that the LSTM hidden layer has 500 units and the rest includes fully connected layer, soft-max layer, and classification layer to generate the classified output. The vibration data files from the Case Western Reserve University Bearing Dataset (CWRU) bearing fault dataset were adopted, which were recorded and made publicly available on the CWRU bearing data center website [20], [21]. Fig. 8 describes the platform. In our simulations for validation, we chose the vibration data files from the 3 horse power (HP) engine running at 1772 RPM with the motor shaft bearings having faults in depth from 0.000 (none), 0.007, 0.014, 0.021, and 0.028 inches and various fault locations (inner race, rolling element, and outer race). Accelerometer was attached at the 12\u2019clock position at the drive end (DE) of the motor housing to measure the vibrations. The acquired signals were sampled at 12 kHz. The data files contain 1 normal class and 11 single point fault classes after merging outer race positions relative to load zone shown in Table III" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000629_9780470612286.ch1-Figure1.3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000629_9780470612286.ch1-Figure1.3-1.png", + "caption": "Figure 1.3. Link frames for the St\u00e4ubli RX-90 robot", + "texts": [ + " This choice makes rn (or \u03b8n) zero when \u03c3n = 1 (or = 0 respectively); \u2013 for a prismatic joint, the axis zj is parallel to the axis of the joint; it can be placed in such a way that dj or dj+1 is zero; \u2013 when zj is parallel to zj+1, the axis xj is placed in such a way that rj or rj+1 is zero; \u2013 in practice, the vector of joint variables q is given by: q = Kc qc + q0 where q0 represents an offset, qc are encoder variables and Kc is a constant matrix. EXAMPLE 1.1.\u2013 description of the structure of the St\u00e4ubli RX-90 robot (Figure 1.3). The robot shoulder is of an RRR anthropomorphic type and the wrist consists of three intersecting revolute axes, equivalent to a spherical joint. From a methodological point of view, firstly the axes zj are placed on the joint axes and the axes xj are placed according to the rules previously set. Next, the geometric parameters of the robot are determined. The link frames are shown in Figure 1.3 and the geometric parameters are given in Table 1.1. 1.2.2. Direct geometric model The direct geometric model (DGM) represents the relations calculating the operational coordinates, giving the location of the end-effector, in terms of the joint coordinates. In the case of a simple open chain, it can be represented by the transformation matrix 0Tn: 0Tn = 0T1(q1) 1T2(q2) \u2026 n-1Tn(qn) [1.4] The direct geometric model of the robot may also be represented by the relation: X = f(q) [1.5] q being the vector of joint coordinates such that: q = [q1 q2 \u2026 qn]T [1", + "7] There are several possibilities to define the vector X. For example, with the help of the elements of matrix 0Tn: X = [Px Py Pz sx sy sz nx ny nz ax ay az]T [1.8] or otherwise, knowing that s = nxa X = [Px Py Pz nx ny nz ax ay az]T [1.9] For the orientation, other representations are currently used such as Euler angles, Roll-Pitch-Yaw angles or Quaternions. We can easily derive direction cosines s, n, a from any one of these representations and vice versa [KHA 02]. EXAMPLE 1.2. \u2013 direct geometric model for the St\u00e4ubli RX-90 robot (Figure 1.3). According to Table 1.1, the relation [1.3] can be used to write the basic transformation matrices j-1Tj. The product of these matrices gives 0T6 that has as components: sx = C1(C23(C4C5C6 \u2013 S4S6) \u2013 S23S5C6) \u2013 S1(S4C5C6 + C4S6) sy = S1(C23(C4C5C6 \u2013 S4S6) \u2013 S23S5C6) + C1(S4C5C6 + C4S6) sz = S23(C4C5C6 \u2013 S4S6) + C23S5C6 nx = C1(\u2013 C23 (C4C5S6 + S4C6) + S23S5S6) + S1(S4C5S6 \u2013 C4C6) ny = S1(\u2013 C23 (C4C5S6 + S4C6) + S23S5S6) \u2013 C1(S4C5S6 \u2013 C4C6) nz = \u2013 S23(C4C5S6 + S4C6) \u2013 C23S5S6 ax = \u2013 C1(C23C4S5 + S23C5) + S1S4S5 ay = \u2013 S1(C23C4S5 + S23C5) \u2013 C1S4S5 az = \u2013 S23C4S5 + C23C5 Px = \u2013 C1(S23 RL4 \u2013 C2D3) Py = \u2013 S1(S23 RL4 \u2013 C2D3) Pz = C23 RL4 + S2D3 with C23=cos (\u03b82 + \u03b83) and S23 = sin (\u03b82 + \u03b83)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001928_6.2008-6485-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001928_6.2008-6485-Figure1-1.png", + "caption": "Figure 1 Axis system with forces acting on the helicopter", + "texts": [ + " In this approach, one of the basic requirements is: the settling time (with bounded steady state error) of a slower mode must be less than or equal to the integration step size of the faster mode. In this paper, the modes are integrated at sampling rates of 10 Hz, 100 Hz, and 1 kHz for the slow, middle, and fast modes, respectively. The model has a total of 14 states and is modified according to Gavrilets26. All terms are defined and given in the Nomenclature section. A. Slow Mode The states of the slow timescale model slowx are the position [ , , ]Tx y z (in inertial coordinates) and velocity [ , , ]Tu v w (in body coordinates) (as shown in Figure 1), whereas , TT slow mrT\u23a1 \u23a4= \u23a3 \u23a6\u0398u are the control variables which include the Euler angles [ ], , T\u03c6 \u03b8 \u03c8=\u0398 and the main rotor thrust mrT . The kinematic relation between the body frame velocity ( u , v , and w ) and the velocity in the inertial coordinate ( x , y , z ) is given by ( ) x u y v z w \u23a1 \u23a4 \u23a1 \u23a4 \u23a2 \u23a5 \u23a2 \u23a5=\u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5\u23a3 \u23a6 \u23a3 \u23a6 R \u0398 (1) where ( )R \u0398 is the direct cosine matrix from the body coordinate to the inertial coordinate with a sequence of rotation sequence of 1-2-3. The translation dynamics is 1 1 / sin / / sin cos / / cos cos / ped fus mr tr fus tr vf ped mr fus mr vr wq X m g a T mu v wp ur Y Y Y m g Y b T m w uq vp Z m g T m \u03b4 \u03b8 \u03c6 \u03b8 \u03b4 \u03c6 \u03b8 \u2212 + \u2212\u23a1 \u23a4 \u23a1 \u23a4\u2212\u23a1 \u23a4 \u23a2 \u23a5 \u23a2 \u23a5\u23a2 \u23a5 \u23a1 \u23a4= \u2212 + + + + + +\u23a2 \u23a5 \u23a2 \u23a5\u23a3 \u23a6\u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5\u23a2 \u23a5 \u2212 + +\u23a3 \u23a6 \u2212\u23a2 \u23a5 \u23a3 \u23a6\u23a3 \u23a6 (2) The forces acting on the fuselage are American Institute of Aeronautics and Astronautics 5 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002999_jifs-219090-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002999_jifs-219090-Figure1-1.png", + "caption": "Fig. 1. Delta mechanism 3D model diagram.", + "texts": [ + " The motor for driving can 115 be installed on the static platform to drive the con- 116 necting rod; the middle connecting rod is composed 117 of three identical branches, each of which connects 118 the active link with the follower link through the 119 ball hinge; the three branches connect the moving 120 platform together and can install the necessary end 121 actuators on the moving platform. The 3D software 122 is used to draw the model of each part of the compo- 123 nent, and then assemble it to get the model shown in 124 Fig. 1. 125 The overall structure of the Delta parallel robot 126 consists of a static platform (upper platform), a 127 mobile platform (lower platform), three active rods 128 (driving rods) and three driven branches (parallelo- 129 gram closed loop). The static platform of the base 130 transmits motion through three identical kinematic 131 chains, which are respectively transmitted through 132 the three sides of the rotating pair connecting the 133 motion platform. The robot not only has the common 134 advantages of parallel robots, but also has good kine- 135 matics and dynamics characteristics due to its simple 136 and compact structure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001782_isic.2007.4450936-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001782_isic.2007.4450936-Figure4-1.png", + "caption": "Fig. 4. A three-link robot manipulator.", + "texts": [ + " Theorem 1: Consider the nth-order multivariable nonlinear systems represented by (1). The intelligent robust control system is defined as in (13), in which the adaptive laws of RCMAC are designed as in (33)-(36) and the robust controller is designed as in (37). Then, the robust tracking performance in (31) can be achieved for the prescribed attenuation level ....,2,1,, miri ),(\u02c6\u02c6 tT w esW (33) ),(\u02c6\u02c6 tT cc esWc (34) ),(\u02c6\u02c6 tT vv esWv (35) ),(\u02c6\u02c6 tT rr esWr (36) ),()()(2 212 tRC esRRu I (37) A three-link robot manipulator is depicted as Fig. 4. The dynamic equation is given as follows [12]: qgqqqCqqM d)(),()( (38) where 3 321 )](),(),([ Ttqtqtqq is the angular position vector, 3, qq are the joint velocity and acceleration vector, respectively, 33)(qM is the inertia matrix, 3 is the input torque vector, and 33),( qqC is the Coriolis/Centripetal matrix, 3)(qg is the gravity vector, 3 d is the external disturbance. The acceleration of gravity 2/8.9 smg . im is the link mass; ia is the link length; the short hand notations are defined as )( jiij qqsins , )( jiij qqcosc ; and id is defined as in Table 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002135_s00202-020-01163-8-Figure10-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002135_s00202-020-01163-8-Figure10-1.png", + "caption": "Fig. 10 The magnetic model of the maximum flux density in the TSCMG given in Fig.\u00a08(b)", + "texts": [ + " As a result, the torque ripples of the inner, middle and outer rotors of the gearbox are better in Fig.\u00a08(a) by 33, 11 and 50%, respectively. In addition, the flux density in the gear given in Fig.\u00a08(a) is distributed more optimal and appropriately. As observed in Fig.\u00a02, using a middle iron rotor will increase the flux density because a thicker core is required, which results in smaller modulators and decreased efficiency of the gear. This law is true for the gear given in Fig.\u00a08(b) as well because the flux density of the middle rotor in the gear is increased. As Fig.\u00a010 shows, the flux density in the gear given in Fig.\u00a08(b) is high. To reduce the flux density, the thickness of the middle rotor needs to be increased, resulting in reduced torque in the magnetic gear. Thus, using the conducted studies, the structure proposed in this paper provides distributed flux density, higher efficiency of the output torque and better performance. Therefore, we can infer that the gear structure Fig. 5 The structure of the proposed TSCMG 1 3 given in Fig.\u00a08(b) cannot be assumed as a proper structure for multi-speed gears" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002748_icepe-p51568.2021.9423481-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002748_icepe-p51568.2021.9423481-Figure2-1.png", + "caption": "Fig. 2: Model of magnetic levitation system.", + "texts": [ + " The simulation results are discussed in section V to demonstrate the effectiveness of the proposed graphical methodology. Finally, concluding remarks are drawn in section VI. Figure 1 shows the ED-4810 magnetic levitation system which is an experimental benchmark allowing theoretical confirmation of the principle of magnetic levitation by suspense steel ball in the air with magnetic force. This system is safe for universities laboratories and it is suitable for testing the effectiveness of various types of controllers due to the high nonlinearity in its model [16]. Figure 2 shows the model of magnetic levitation system, where: m: Weight of steel ball. y(t): Center position of the ball. i(t): Electric current flowing on an electromagnetic coil. c: Magnetic force constant. L: Inductance of wire. R: Resistance of wire. e(t): Input voltage. equation (1) explain mathematically the relationship between the voltage and the current in the coil. e(t) = Ri(t) + L di(t) dt (1) The total summation of the forces around the vertical shown in figure 2 is: ftotal = fgravity + fem (2) Where fgravity is the gravity force and fem represents the electromagnetic force caused by the upper coil. The following first order differential equations represent the mathematical model of the physical system shown in figure 2. mY\u0308 = mg \u2212 c i2 y (3) x\u0307a = xb (4) x\u0307b = g \u2212 cx2 c(t) mxa (5) x\u0307c = \u2212R L xc(t) + 1 L u(t) (6) where: xa = y(t), xb = y\u0307 ,xc = i(t) ,u(t) = e(t). For convenience, the above equations can be represented in state space format as shown in equation (7) [16]. \u23a1 \u23a3x\u0307a x\u0307b x\u0307c \u23a4 \u23a6 = \u23a1 \u23a2\u23a2\u23a2\u23a3 0 1 0 g xa 0 \u2212cxc mxa 0 0 \u2212R L \u23a4 \u23a5\u23a5\u23a5\u23a6 \u23a1 \u23a3xa xb xc \u23a4 \u23a6+ \u23a1 \u23a2\u23a3 0 0 1 L \u23a4 \u23a5\u23a6u(t) (7) Obviously, the resultant state space model is nonlinear. Hence, these equations should be linearized in order to apply the proposed method" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000195_b978-0-443-07113-3.50014-7-Figure10.6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000195_b978-0-443-07113-3.50014-7-Figure10.6-1.png", + "caption": "Figure 10.6 Diagram of two people pushing a bed with forces Fs 1 and F 2.", + "texts": [ + " In order to define the weight force fully, its point of application on the object has to be defined. This is called the centre of gravity and is, in effect, the same as centre of mass. It is the point at which the mass of the object may be considered as being concentrated. As mentioned above, the addition of vectors is different from the addition of scalars. With scalars, the addition is simply the sum of the two numbers, but with vectors we need to use vector addition. This requires the use of vector diagrams. Figure 10.6 shows two people pushing a bed, each with a different force (F1 and F2) and in a different direction. To find the total force acting on the bed and its direction of motion, we need to redraw the diagram as shown in Figure 10.7. Here the two forces are represented by arrows. The length of each indicates the magnitude of the force, and the angle indicates the direction, or line, of action. The lines are drawn parallel to their actual positions: F2 follows from the end of F1. If there were more forces involved, they could be added to the end of the last arrow in any sequence" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002022_mmvip.2007.4430742-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002022_mmvip.2007.4430742-Figure1-1.png", + "caption": "Fig. 1. Working principle of fibre optical displacement sensor.", + "texts": [ + " The reflective intensity modulated (RIM-FOS) is a kind of non- function type fibre optical sensor, fibre only has the transferred light function. The fibres are divided into two parts, the input fibre and the output fibre. The modulated theory of this kind of sensor is that the input fibre make light source only carry to the measured object surface, then is reflected from measured surface to another output fibre, its light intensities change with the distance of measured surface and fibre, the light which is reflected back to the output fibre is received by the detector, then it is converted to electric signals to export. Figure 1 shows the work principle of reflective intensity modulated fibre optical sensor, the measured object which is a plane surface reflector is in the location which is d in the space between it and the fibre optical end face, moving it along axial of input and output fibre, so the virtual image of the input fibre is formed in the distance d after the plane surface reflector. Therefore, to ascertain the responding of modulator is equivalent to calculating coupling of the virtual fibre and the input fibre" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003132_978-981-16-0119-4_20-Figure19.3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003132_978-981-16-0119-4_20-Figure19.3-1.png", + "caption": "Fig. 19.3 Forces on the \u2018bottom bracket\u2019", + "texts": [ + "\u2019 The final purpose of a \u2018product design\u2019 is practical usage. Manufacturing a product requires sharing a drawing that the manufacturer\u2019s team can use. For this study to be effectively used, it is assumed that a student has an understanding of introductory courses in \u2018Applied Mechanics\u2019 and \u2018Strength of Materials.\u2019 From a \u2018system level\u2019 standpoint, tractive effort is an important parameter needed for designing any vehicle (see Fig. 19.2). This effort is an input to estimate the pedaling force (see Fig. 19.3). As the handlebar controls the steering and supports the inertia of the rider, estimating the load on the handlebar is important (see Fig. 19.4). The saddle supports some rider inertia (see Fig. 19.5). To simplify the study, loads along the lateral direction of the bicycle are not included in this case study. For finding the tractive effort, a process explained by Krishnakumar4 has been adapted. Using mechanics, one can estimate the forces required on the pedal, once a value for the tractive effort on the rear wheel is known" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002657_tmag.2021.3073155-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002657_tmag.2021.3073155-Figure2-1.png", + "caption": "Fig. 2. Topologies of the 6/17 FP-MFSPM machine.", + "texts": [ + " Finally, several conclusions are drawn in Section IV. The 6/17 C-MFSPM machine is presented in Fig. 1, where phase winding connection is that the two coils spatial radially distributed are serially connected into one phase winding, such as A1 and A2 for phase winding A. It should be noted that the armature windings adopt a concentrated structure, and coil conductors layout of phase A is A1+, A1-, A2-, A2+. By changing the winding configuration from concentrated winding to full-pitch winding, a 6/17 FP-MFSPM machine is formed, as is shown in Fig. 2, where coil conductors layout of phase A turns into A1'+, A2'+, A1'-, A2'-. The 6/17 C-MFSPM and FPMFSPM machines are identical except for the winding configurations. W Authorized licensed use limited to: Robert Gordon University. Downloaded on May 29,2021 at 23:52:19 UTC from IEEE Xplore. Restrictions apply. 0018-9464 (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" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002197_j.matpr.2020.12.125-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002197_j.matpr.2020.12.125-Figure5-1.png", + "caption": "Fig. 5. Gear pair assemblyfor (a) 20/20 Pressure angle; (b) 20/40 Pressure angle.", + "texts": [ + " 3 highlights the gear pair contact assembly. Fig. 4 highlights the zoomed view of the mesh transition from coarser to finer elements at the involute contact region which extend till gear tooth fillet. Reduced integrated elements formulation was used to reduce the total time taken for solving the complete analysis. Also *CONTROL_HOURGLASS was included to reduce the possibility of occurrence of hourglass modes. Care was taken to make sure that hourglass energy was less than 5% of the total energy of the system. Fig. 5 highlights the assembly for 20/20 and 20/40 pressure angle teeth gear pair (20 represents coast side & 40 represents the drive side pressure angle). In a static-implicit method, the contact analyses between any mating parts are calculated by representing it based on contact element approach. However, in the LS-DYNA [12] software there is no need of separate contact element, as long as there are definitions of contact surfaces with addition of specific type of contact between them. Additional contact parameters are also specified, so that the program solves easily utilizing these contact interfaces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002456_s11041-021-00623-7-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002456_s11041-021-00623-7-Figure2-1.png", + "caption": "Fig. 2. Scheme of the process of laser beam welding: 1 ) laser beam; 2 ) clamp, 3 ) welding direction; 4 ) sample support.", + "texts": [ + " Before the welding, the surfaces of the sheet were cleaned from the oxide films by polishing with a SiC abrasive paper with grit 1000 and then treated with ultrasound in acetone. The welding was conducted with the help of a YLS-5000 fiber laser with maximum radiation power 5 kW. To protect the molten pool and the solidified fusion zone at high temperature from atmosphere, three kinds of protection with pure argon were applied, i.e., front shielding gas, back shielding gas, and protective cover gas with flow rates 25, 15, and 15 liters min, respectively. Figure 2 presents the scheme of laser welding of alloy Ti60. The welding parameters were as a follows: laser power 2 kW, welding speed 1.2 m min, defocusing 2 mm. After the welding, the joints were subjected to a heat treatment (PWHT) involving 2-h exposure at 940\u00b0C, cooling with the furnace to room temperature, 1.5-h exposure at 600\u00b0C, and cooling with the furnace to room temperature. The scheme of the PWHT is presented in Fig. 3. We cut transverse metallographic specimens and specimens for tensile tests for each mode of welding and welding + PWHT" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000676_978-1-4020-8829-2_9-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000676_978-1-4020-8829-2_9-Figure7-1.png", + "caption": "Fig. 7. Wrong assembly of Delta robot model", + "texts": [ + " Accordingly, relation (28) becomes: [ q\u0307a q\u0307f ] = [ Bqau Bqf u ] x\u0307 , (30) from which we can deduce in comparison with (24) that: Jf = B\u22121 qau . (31) Therefore, in (27), the condition number of Jf is equal to the condition number of Bqau. This optimization problem has also some restrictions to satisfy. Bound constraints have been specified to get a reasonable solution. Minimum and maximum values are reported in Table 2. Concerning the assembly problem, we should avoid to close the mechanism toward an \u201cinside\u201d position of the legs as shown in Fig. 7. Therefore, some inequality constraints c have been formulated in terms of sines and cosines of the dependent variables. All these considerations lead to the extended objective function g (see (8a)). The optimization has been first performed with only two of the five parameters to visualize the process. The leg lengths are optimized but the radii rp and rb, and the distance zc are fixed: rp = 2 cm, rb = 5 cm, and zc = \u221210 cm. Two different starting points Start1 and Start2 are proposed to highlight the existence of two different local minima", + " Thus, we make use of the characteristic length LC to divide the \u201cposition\u201d columns of Bvu, defining the homogeneous Jacobian J\u0303f involved in the optimization as: J\u0303f = [ Bqaxp LC Bqaxo ]\u22121 . (34) The objective function f of this 6-dimension problem is finally: f (zc, LI, LS,RB,RP,H, \u03b1, \u03b2, \u03c8, LC) = 1 32 32\u2211 i=1 1 \u03ba (( J\u0303f ) i ) . (35) Also for practical reasons, design parameters of this robot are limited by bounds (see minimum and maximum values in Table 3). The same restrictions are imposed on the assembly problem to avoid \u201cinside\u201d configurations as in Fig. 7. This leads to the final extended function g (see (8a)). The results obtained with the SQP method are presented in Table 3, and initial and optimal design can be compared in Fig. 11. The graphical results are represented in Fig. 12. It takes 94 iterations and 110 evaluations of the objective function to optimize the 10 parameters. In Fig. 5.2, we can observe the evolution of the objective function in three main steps. These steps correspond to \u201csequences\u201d of the SQP method for which active constraints are defined" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000707_2008-01-0780-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000707_2008-01-0780-Figure1-1.png", + "caption": "Figure 1. Schematic of a main battle tank", + "texts": [], + "surrounding_texts": [ + "The ride vibration environment of a typical high-speed military tracked vehicle traversing rough off-road terrain is a significant factor due to the magnitude of ride vibrations arising from dynamic terrain-vehicle interactions. In this paper, ride dynamics of a \u201c2+N\u201d degrees of freedom (DOF) tracked vehicle mathematical model has been evaluated for rough off-road terrain modeled as a sinusoidal profile of different pitch and height at constant running speeds. The equations of motion are derived using Newton\u2019s laws of motion. A comparison of ride dynamic analysis of the vehicle fitted with conventional passive suspension system and hydrogas suspension system is made." + ] + }, + { + "image_filename": "designv11_83_0001769_iccas.2007.4406887-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001769_iccas.2007.4406887-Figure8-1.png", + "caption": "Fig. 8 PSD arrangement and recognition of slanted wall.", + "texts": [ + " We plan the arc-like path from the origin to the sub-goal, as shown in Fig. 7. We also plot the wall along with the path responsible for the collision en route to the sub-goal. The amplitude of the reflected waves reduces as the angle between the wall and the transducer increases. Therefore, the robot is unable to recognize the slant wall in front of it. We use a position sensitive detector (PSD) [4], [5] in order to compensate for the shortcomings of the proposed system. We install four PSDs in the front section of the robot (as shown in Fig. 8) in order to recognize the situation in which the robot moves obliquely toward the wall. When the PSD system recognizes such a situation, it corrects the sub-goal calculated by the sonar sensor. Further, when the PSD system determines that it is difficult to generate an arc-like path, the robot stops and performs a 360 degree scan. The PSD system works effectively when a robot moves obliquely toward the wall or it exists at the end of the hall. The integration-type ultrasonic wave sensor system usually performs a 180 degree forward scan and searches for a traversable area" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002892_s11665-021-05931-w-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002892_s11665-021-05931-w-Figure2-1.png", + "caption": "Fig. 2 Meshed geometries for finite element analysis (FEA) of heat flow within features of different thicknesses", + "texts": [ + " After the PDAS values were obtained, the corresponding cooling rates were calculated as (Ref 30): dp \u00bc 80\u00f0CR\u00de 0:33 \u00f0Eq 2\u00de where dp is the PDAS (lm) and CR is the corresponding cooling rate ( K/s). Powder x-ray diffraction (XRD) patterns were collected with a diffractometer equipped with a 2.2 kW Cu-Ka x-ray source (Smartlab, Rigaku, Tokyo, Japan) configured in BraggBrentano geometry. The patterns were collected in the 2h range of 20-120 with 0.02 step and at a scan speed of 4 /s. To simulate heat transfer during LPBF, a 3D FEA model was constructed in COMSOL 5.3a Multiphysics software. Three geometries were meshed, corresponding to the 0.5, 2.5 and 5 mm legs (Fig. 2). The finer meshed regions were assigned to the features scanned by the laser beam. The coarser meshed regions represented the powder bed surrounding the deposited build and were assumed to have a thermal conductivity of 0.01 times of the bulk SS 316L alloy (Ref 31). Rectangular elements with a maximum size of 0.03 mm were used for the central, finer meshed region, and free tetrahedral elements with a maximum size of 0.28 mm were used for the coarser meshed regions. To simulate the laser deposition process, the thermal conductivity of the elements switched from that of the loose powder to that of solid bulk SS 316L after the laser spot moved past the elements" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002804_j.matpr.2021.04.568-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002804_j.matpr.2021.04.568-Figure1-1.png", + "caption": "Fig. 1. 3D Model of M", + "texts": [ + " The 3D modelling was carried out and analyzed with ANSYS on both a steel and a composites blade spring. CAD models are designed in CATIA which contains special tools for the generation of traditional surfaces which will then be transformed into solid models, using traditional and composite monoleaf spring materials. The measurements of a spring leaf of a TATA SUMO vehicle are used for the design of the mono-leaf spring. Table1 displays the mono leaf spring configuration parameters and the planned mono leaf spring CAD model, as shown in Fig. 1. Finite Element Analysis (FEA) is a computational method for deconstruction into very small elements of a complex structure. In the simulated world, ANSYS offers an affordable way to examine the success of goods or processes. Digital prototyping is called this method of product growth. Users will iterate different scenarios using simulated prototyping techniques to refine the software even before the production is launched. This allows the probability and expense of failed designs to be reduced" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000088_j.jsv.2006.01.044-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000088_j.jsv.2006.01.044-Figure7-1.png", + "caption": "Fig. 7. Rotating coordinate system.", + "texts": [ + " / Journal of Sound and Vibration 295 (2006) 890\u2013905 899 method is robust for vibration analysis of structure, such that the application proves that the current crack beam element and the approach for cracked crankshaft model are valid for the vibration analysis of cracked crankshaft. In order to find the dynamic response of the crankshaft-bearing system, the rotating crankshaft must be analyzed by combining with the non-rotating (fixed) main bearing. A common coordinate system is required for the combined model. For this reason, the analysis is performed with respect to a right-handed rotating coordinate system \u00f0X ; Y ; Z\u00de that is attached to the crankshaft (see Fig. 7). The X coordinate is along the crankshaft axis towards the flywheel end and the Y coordinate point always to crankpin. The dynamic analysis described in this report assumes that the mass matrix [M] of crankshaft system is time invariant. Let the rotating speed of the crankshaft be o, the angle between the initial position and the coordinate origin \u00f0x; y; z\u00de be 0 (i.e., when t \u00bc 0, system \u00f0X ; Y ; Z\u00de is superposed with system \u00f0x; y; z\u00de). Therefore, under the rotating coordinate system the supporting stiffness [Ka] is \u00bdKa \u00bc KYY KYZ KZY KZZ \" # \u00bc \u00bdT0 T kyy 0 0 kzz \" # \u00bdT0 , (11) where \u00bdT0 \u00bc cos\u00f0ot\u00de sin\u00f0ot\u00de sin\u00f0ot\u00de cos\u00f0ot\u00de \" # " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001195_09544062jmes906-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001195_09544062jmes906-Figure3-1.png", + "caption": "Fig. 3 FE model: (a) coil spring, (b) finite contact spring elements, and (c) configuration of the contact between a spring and a spring pad; the 0.75 turns of the coil spring is contacted with the pad", + "texts": [ + " Keq is the bulk stiffness of the spring pad assembly, K1 is the equivalent bulk stiffness of an upper spring pad, K2 is the equivalent bulk stiffness of a lower spring pad, and Ks is the static spring stiffness corresponding to the noncontact length of the spring. The static stiffness of the coil spring can be determined such that Ks \u00bc Gd4 8D3 Neff \u00f03\u00de whereG is the shearmodulus, d the diameter of a coil, D the central diameter of the helical spring, and Neff the number of spring turns for the non-contact length of the spring (which is also called the effective turns of coil). To develop the contact kinematics between the coil spring and two spring pads, the finite contact element model is constructed as depicted in Fig. 3. The point contact stiffness is assumed to be uniform over the contact area. Therefore, the contact stiffness on both sides of the spring is kc1 \u00bc K1/NC1 and kc2 \u00bc K2/NC2, where NC1 and NC2 are the numbers of the upper point and lower point contacts, respectively. The virtual work due to the upper contact stiffness and the lower contact stiffness can be, respectively, Proc. IMechE Vol. 222 Part C: J. Mechanical Engineering Science JMES906 # IMechE 2008 at PRINCETON UNIV LIBRARY on July 10, 2015pic", + " Mechanical Engineering Science JMES906 # IMechE 2008 at PRINCETON UNIV LIBRARY on July 10, 2015pic.sagepub.comDownloaded from following form Fspring \u00bc kc1 XNC1 k\u00bc1 XN i\u00bc1 fi\u00f0x U k \u00de qi\u00f0t\u00de \u00fe Cc1 XNC1 k\u00bc1 XN i\u00bc1 fi\u00f0x U k \u00de _qi\u00f0t\u00de \u00f014\u00de The magnitude of Fspring can be obtained by substituting equation (12) into equation (14). The modal analysis of the component coil spring is conducted by ANSYS. Brick 95 in ANSYS is used to mesh the coil spring solid geometry. Nearly 14 000 FEs are employed and the mesh is conducted by sweeping the geometry in the axial direction of the spring coil as shown in Fig. 3(a). By this analysis, the modal vectors corresponding to the natural vibration modes of the component coil spring are obtained and used for the construction of the contact stiffness matrices and the contact damping matrices as described in equation (9). The nine vibration modes in the ascending order of the natural frequencies are illustrated in Fig. 4. It is seen that the lower vibration modes are dominated by the longitudinal modes and the bending modes. The corresponding natural frequencies of the coil spring are summarized in Table 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002979_iemdc47953.2021.9449609-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002979_iemdc47953.2021.9449609-Figure4-1.png", + "caption": "Fig. 4 Cross-sections of optimized 24-slot/28-pole (a) modular and (b) nonmodular SPM machines. In (a), the flux gap width is 10.2mm.", + "texts": [ + " Optimization of the 24-slot/28-pole Modular Machine In order to optimize the 6 variables shown in Table II, 1397 cases are generated to find the global optimization. As the primary objective, the average torque is used as y-axis to derive the correlation curve with other three secondary objectives and the optimal case is calculated by the fitness function in (1) and (2), as shown in Fig. 3, where the optimal cases are highlighted. The comparison of variables between the initial design and the optimal design is shown in Table III, while the cross-section of the optimized machine is shown in Fig. 4 (a). TABLE III COMPARISON BETWEEN ORIGINAL AND OPTIMAL CASES OF 24- SLOT/28-POLE MODULAR MACHINES Original Optimal Difference (%) Split ratio 0.80 0.84 5.0 Stator tooth width (mm) 10.0 13.2 32.0 Stator yoke thickness (mm) 10.0 5.6 -44.0 Rotor yoke thickness (mm) 10.0 8.6 -14.0 Flux gap width (mm) 6.2 10.2 64.5 Magnet thickness (mm) 7 6.6 5.1 Average torque (Nm) 285.3 331.2 16.1 Torque ripple (%) 31.4 4.0 -87.3 Copper loss (W) 761 956 25.5 Total machine mass (kg) 37.3 31.2 -16.4 Efficiency (%) 95", + " Optimization of 24-slot/28-pole Non-Modular Machine In order to demonstrate the influence of flux gaps in the optimization process of SPM machines, the conventional nonmodular SPM machine is also optimized by GA with the same objectives and constraints as the modular machines. There are 1320 cases generated by GA and the correlation curves are shown in Fig. 5, while the specifications of the optimized nonmodular machine are listed in Table IV. The cross-section of the optimized 24-slot/28-pole non-modular machine is shown in Fig. 4 (b). It is worth noting that the optimal design chosen is far away from the Pareto front of the torque and torque ripple correlations because the weight (0.1) of the minimized torque ripple is smaller than the weight for other objectives. After the GA optimization, the average torque is increased by 19.7%, and the total machine mass and torque ripple are reduced by 15.8% and 56.3%, respectively. In addition, the efficiency of the optimized non-modular machine is increased by 0.9%. D. Evaluation of Optimized 24-slot/28-pole Modular SPM Machine To prove the importance of considering the flux gap width in the global optimization, the results obtained using method 1 and method 2 are compared in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001007_s1068798x08120113-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001007_s1068798x08120113-Figure3-1.png", + "caption": "Fig. 3. Determining the indenter\u2013surface contact parameters: (a) rigidity coefficient C0; (b) ridge height hr of the deformed material.", + "texts": [ + " 2, we write the following differential equations: for section 1 (1) for section 2 (2) for section 3 (3) for section 4 (4) m1 d2y dt2 ------- B1 dy dt ----- C0iny+ + P Ffr\u2013= when y 0, dy/dt 0;> > m1 d2y dt2 ------- B1 dy dt ----- C0outy+ + P Ffr+= when ymax y hr,\u2013 dy/dt 0;<> > m1 d2y dt2 ------- B1 dy dt -----+ P Ffr+= when y hr,\u2013 dy/dt 0;< < m1 d2y dt2 ------- B1 dy dt -----+ P Ffr\u2013= when y 0, dy/dt 0,>< RUSSIAN ENGINEERING RESEARCH Vol. 28 No. 12 2008 NONLINEAR DYNAMICS OF THE ELASTIC SMOOTHING OF SURFACES 1203 where C0in and C0out are the rigidity coefficients of indenter\u2013surface contact in sections 1 and 2, respectively; Ffr is the frictional force on the indenter. To determine C0, we adopt the calculation scheme in Fig. 3. The dependence of C0 on the indenter depth in the surface is nonlinear, for the following reasons: 1) introducing the indenter in the part increases the indenter\u2013part contact area, with proportional increase in C0; 2) indenter motion leads to the formation of a ridge on the surface, which increases C0; 3) the ridge height depends on the indenter\u2019s introduction depth y and the direction of its vibrational velocity (sgn ). Given that the rigidity of indenter\u2013point contact is proportional to the contact area of the indenter with the surface, we write C0 = K1ScoHB, where Sco is the indenter\u2013surface contact area; K1 is a correction coefficient", + " Energy input to the dynamic system increases the oscillation amplitude or leads to stability loss. The subsequent processes depend on the balance between energy input and scattering. Energy is supplied to the self-oscillatory processes by the indenter\u2013surface interaction. Energy accumulates in the dynamic system when the indenter is withdrawn from the surface and is due to the nonlinear ridge formation. The energy input to the self-oscillatory process may be characterized as follows: was indenter 1 (Fig. 3a) is introduced in the surface, the metal is displaced and hardened, with the formation of a ridge (hr) ahead of the indenter. When the indenter is introduced in the surface, three stages may be identified. 1. Crumpling of the initial roughness (line OA in Fig. 3b). The indenter only interacts with surface projections. Contact points with high stress arise, and the peaks are contorted, filling the troughs. 2. Indenter introduction in the surface layer of the metal (line AB). The indenter begins to displace and compact the metal in the contact region. 3. Formation of a ridge of hardened metal (point B). Ahead of the indenter, an elastic wave of displaced metal is formed; its height is proportional to the depth of indenter introduction. On account of elastic-wave formation ahead of the indenter, the section where it leaves the surface differs significantly from the section where it is introduced", + " With increase in the contact rigidity when dy/dt < 0 and hence in the force on the indenter, the oscillation amplitude and velocity of the indenter increases in each cycle of extraction from the surface. Consequently, energy accumulates in the indenter\u2019s oscillation. The energy stored per oscillatory cycle depends on the ratio of the smoothing speed and the indenter\u2019s vibrational 1204 RUSSIAN ENGINEERING RESEARCH Vol. 28 No. 12 2008 KUZNETSOV, GORGOTS speed. Reducing the smoothing speed changes the indenter\u2019s interaction with the surface (Fig. 3a). With decrease in smoothing speed, breakaway of the indenter\u2019s working surface 2 from the smoothed surface 1, with the formation of a gap L, is observed at a negative oscillatory speed dy/dt of the indenter. Given that the depth y + hr of indenter introduction is significantly less than the indenter radius R, we may assume that This formula permits the determination of the critical smoothing speed at which the rigidity of the indenter\u2013surface contact is sharply reduced for kinematic reasons and the conditions required for self-oscillation are disrupted Thus, in energy terms, self-oscillation is only possible on smoothing at above-critical speed (v > vcr) and consists of the energy-accumulation phase in section 2 and the energy-dissipation phases in sections 1, 3, and 4 (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002859_s12046-020-01541-9-Figure12-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002859_s12046-020-01541-9-Figure12-1.png", + "caption": "Figure 12. Effect of section thickness on hardness.", + "texts": [ + " The yield strength and tensile strength of different sections of stepbar casting are given in figure 11. It can be observed that the yield strength of the materials is in the range 248\u2013370 MPa. The yield strength of the 5-mm section is 260\u2013295 MPa, whereas the yield strength of 10 and 15 mm is in the range 235\u2013250 and 235\u2013250 MPa, respectively. Similarly the tensile strength of the 5-mm section is 330\u2013370 Mpa, whereas the yield strength of 10 and 15 mm is in the range 270\u2013290 and 220\u2013240 MPa, respectively. The Brinell hardness of stepbar casting is in the range 167\u2013227 BHN as shown in figure 12. It can be observed that the hardness of the 5-mm section is in the range 197\u2013227 BHN, whereas for 10 and 15 mm BHN is 180\u2013197 and 167\u2013180, respectively. This variation in hardness with the section thickness is also due to the variation in cooling rate. Small sections contain a higher volume fraction of pearlite, which leads to higher hardness, whereas thick sections have lesser volume fractions of pearlite and hence lower hardness. Brinell hardness is also found to reduce as the section thickness increases" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000471_wst.2004.0679-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000471_wst.2004.0679-Figure1-1.png", + "caption": "Figure 1 Cross section of CHEMFET structure116", + "texts": [ + " Operation principle The ion selective transistors are solid-state miniature sensors based on silicon technology. From the chemical point of view, these sensors can be an attractive alternative on the market to the classical potentiometric sensors because of their relatively good analytical performance, low price, and small size. These sensors, often referred as ChemicallyModified Field Effect Transistors (CHEMFETs), are devoted to the detection of particular species in surrounding electrolyte. The base for all CHEMFETs constitutes the Ion Sensitive Field Effect Transistor (ISFET). As shown in Figure 1, in ISFETs the classic gate of an ordinary FET is replaced by a more complex structure consisting of a reference electrode, an analysed solution and a gate dielectric. The hydrogen ion concentration in the solution influences the gate potential, which in turn modifies the transistor threshold voltage. In this way, the ion concentration exercises electrostatic control on the drain-source current. Such structure is capable of sensing the concentration of the hydrogen ions and is used as a pH sensor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002935_978-3-030-77102-7_10-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002935_978-3-030-77102-7_10-Figure7-1.png", + "caption": "Fig. 7. The agents in the environment (Color figure online)", + "texts": [ + " t1 P2\u03bb \u2192 A(rd) B P1 P3 t2 \u03bb \u2192 B(d) P4 t3 \u03bb \u2192 A( ) t5 t4 P5 \u03bb \u2192 B( ) P6 \u03bb \u2192 A(u) H P8 P7 t7 t6 \u03bb \u2192 H(c1, c2, c4, c5, c6, c8) \u03bb \u2192 H( ) Fig. 5. TATPN1 generating Honeycomb patterns The derivation is shown in Fig. 6. Fig. 6. Honeycomb patterns We can generate different sizes of this picture as the number of times t6 and t7 fire. (ii) We construct a TATPN which generates Swastik pattern. The swastika is a geometrical figure and an ancient religious icon in the cultures of Eurasia. Let . Using these tiles, EER and PER on transitions of TATPN2 as in Fig. 7, we get the swastik pattern. In TATPN2, when the transitions t1t2t3 fires, we get the pattern in Fig. 8(a) and again when the transition t4 fires we get Fig. 8(c) and again when the transition t5 fires, we get Fig. 8(d) and when the transition t6 fires, we get the swastik pattern in Fig. 8(e) and when the transition t7 fires, we get the swastik pattern in Fig. 8(f) and when t8 fires, it removes the token Q from P8 and it uses the rule \u03bb \u2192 Q(u, r, d, l) and deposits the pattern Fig. 9 in P9. When t9 fires we get the swastik pattern as in Fig", + " Transition with catenation rule after firing the position of the partial array after transition t1 fires. Since the catenation rule is associated with the transition, catenation takes place in P3. In A\u2666 = a a a a \u2666 a a a a , the number of columns of A is 3, n \u2212 1 is 2, firing t1 adds the row x x y as the last row. Hence A1\u2666 = a a a a \u2666 a a a a x x y Example 3. Let \u03a3 = {a}, F = P1, where S\u2666 = a a a a \u2666 a a a a , Q1 = (\u2666)m, Q2 = (\u2666)n Q3 = (a)m, Q4 = (a)n S is the initial partial array placed in P1. The PATPNS is shown in Fig. 7. Derivations in PATPNS is given in the following tabular column. Input place Transition Output place S A | Q1 a a a \u2666 a \u2666 a \u2666 a a a \u2666 a a a \u2666 a \u2666 a \u2666 a a a \u2666 Q1 | A \u2666 a a a \u2666 \u2666 a \u2666 a \u2666 \u2666 a a a \u2666 \u2666 a a a \u2666 \u2666 a \u2666 a \u2666 \u2666 a a a \u2666 A \u2212 Q2 \u2666 a a a \u2666 \u2666 a \u2666 a \u2666 \u2666 a a a \u2666 \u2666 \u2666 \u2666 \u2666 \u2666 Partial Array Token Petri Net and P System 143 Input place Transition Output place \u2666 a a a \u2666 \u2666 a \u2666 a \u2666 \u2666 a a a \u2666 \u2666 \u2666 \u2666 \u2666 \u2666 Q2 \u2212 A \u2666 \u2666 \u2666 \u2666 \u2666 \u2666 a a a \u2666 \u2666 a \u2666 a \u2666 \u2666 a a a \u2666 \u2666 \u2666 \u2666 \u2666 \u2666 \u2666 \u2666 \u2666 \u2666 \u2666 \u2666 a a a \u2666 \u2666 a \u2666 a \u2666 \u2666 a a a \u2666 \u2666 \u2666 \u2666 \u2666 \u2666 A | Q3 \u2666 \u2666 \u2666 \u2666 \u2666 a \u2666 a a a \u2666 a \u2666 a \u2666 a \u2666 a \u2666 a a a \u2666 a \u2666 \u2666 \u2666 \u2666 \u2666 a \u2666 \u2666 \u2666 \u2666 \u2666 a \u2666 a a a \u2666 a \u2666 a \u2666 a \u2666 a \u2666 a a a \u2666 a \u2666 \u2666 \u2666 \u2666 \u2666 a Q3 | A a \u2666 \u2666 \u2666 \u2666 \u2666 a a \u2666 a a a \u2666 a a \u2666 a \u2666 a \u2666 a a \u2666 a a a \u2666 a a \u2666 \u2666 \u2666 \u2666 \u2666 a a \u2666 \u2666 \u2666 \u2666 \u2666 a a \u2666 a a a \u2666 a a \u2666 a \u2666 a \u2666 a a \u2666 a a a \u2666 a a \u2666 \u2666 \u2666 \u2666 \u2666 a A \u2212 Q4 a \u2666 \u2666 \u2666 \u2666 \u2666 a a \u2666 a a a \u2666 a a \u2666 a \u2666 a \u2666 a a \u2666 a a a \u2666 a a \u2666 \u2666 \u2666 \u2666 \u2666 a a a a a a a a a \u2666 \u2666 \u2666 \u2666 \u2666 a a \u2666 a a a \u2666 a a \u2666 a \u2666 a \u2666 a a \u2666 a a a \u2666 a a \u2666 \u2666 \u2666 \u2666 \u2666 a a a a a a a a Q4 \u2212 A a a a a a a a a \u2666 \u2666 \u2666 \u2666 \u2666 a a \u2666 a a a \u2666 a a \u2666 a \u2666 a \u2666 a a \u2666 a a a \u2666 a a \u2666 \u2666 \u2666 \u2666 \u2666 a a a a a a a a The firing of sequence (t1t2t3t4t5t6t7t8)k, k \u2265 0 puts a square partial arrays of size 4k + 3 in P1, where the boundaries of the squares are alternatively \u2666\u2019s and a\u2019s", + " This modification is still in the scope of the P colonies, and we were not forced to introduce some outer mechanism setting the hierarchy of the pack. The wolves were able to set the hierarchy themselves, and the simulation proceeded as expected. The following figures show the simulator, the settings, the process of the simulation, and the results. Fig. 5. The instance of the environment In the Fig. 5, there can be seen an instance of the environment. Source environment file is displayed on the right. If we move the mouse cursor to some field, we obtain the coordinates of it. In the Fig. 7, the agents are initialized within the environment. The vectors of the blackboard have all their values initialized to zero. The agents are represented by the blue text color. By moving the mouse cursor to the position of the agent, we obtain detailed information about it in the form: Ai = ({obj1, obj2}, [posX, posY ]), where i is the index of the agent, obj1 and obj2 are the objects inside this agent, and posX and posY are the coordinates of its position. The Step button runs a predefined number of the steps of the simulation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000837_imece2008-67118-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000837_imece2008-67118-Figure1-1.png", + "caption": "Fig. 1: Frontal view of the simplified biped model", + "texts": [ + "org/about-asme/terms-of-use Do although we are tuning all the three PID gains, we have tried to simplify the calculations and by that lower the needed computational power. In our method we have used a trial and error method to tune the membership functions and created the rules by human expert knowledge. In this study, we have proposed a simplified model of human in a special situation trying to maintain it's balance in frontal plane. In this case feet are supposed be fixed to the ground similar to a 1 DOF revolute joint. Also hands are extended to help maintaining balance easier. Therefore, a model of fully actuated planar 3 linked segment, shown in fig. 1, consisting of limb, trunk and extended arms with fixed base is used. Let l1, l2, and l3 be the length of links1, 2, and 3, respectively. Let the angels between link1 and the positive Y axis, link2 and the extension of link1, link3 and perpendicular line to link2 be \u03b81, \u03b82, and \u03b83, respectively. The Lagrange's method is used to model dynamics of the robot. Therefore, the Lagrangian function is written as; UTL \u2212= (1) Where T is the kinetic energy and U is the potential energy. The kinetic energy of the model is obtained as follows; 2 33 2 33 2 22 2 22 2 11 2 11 2 1 2 1 2 1 2 1 2 1 2 1 \u03c9\u03c9 \u03c9 CCC CCC IVmI VmIVmT +++ ++= (2) where m1, m2, m3, I1C, I2C, and I3C are mass and moment inertia of link 1, 2, and 3, respectively and we have: wnloaded From: http://proceedings", + "1: Implemented rules for \"K'p\" ( )e t NB NM NS ZO PS PM PB NB S S S S S S S NM B B S S S B B NS B B B S B B B ZO B B B B B B B PS B B B S B B B PM B B S S S B B ( )e t PB S S S S S S S Tab.2: Implemented rules for \"K'd\" wnloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/28/2 ( )e t NB NM NS ZO PS PM PB NB S S S S S S S NM MS MS S S S MS MS NS M MS MS S MS MS M ZO B M MS MS MS M B PS M MS MS S MS MS M PM MS MS S S S MS MS ( )e t PB S S S S S S S TAB.3: Implemented rules for \"\u03b1\" The effectiveness and capability of the proposed method has been tested by a series of simulations for the robot in fig. 1 with l1=1 m, l2=0.8 m, and l3=2 m. Let m1=35 kg, m2=35 kg, m3=10 kg, I1c=2.9167 kg.m2, I2c=1.8667 kg.m2, I3c=3.3333 kg.m2. The mass of all links of the robot are considered as distributed mass and their COMs are in the middle of each link. The ranges of Kp, Kd and \u03b1 are determined, as shown in Tab.4, based on the method described in [18] and the experimental results. figs.7-9 show the results of simulation for an initial value of angular positions and their first derivatives. Joints number Parameter min max Kp 0 100 Kd 0 100 1 \u03b1 1000 1010 Kp 0 750 Kd 0 75 2 \u03b1 1500 1510 Kp 0 50 Kd 0 25 3 \u03b1 20 30 Tab" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000817_1.2712953-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000817_1.2712953-Figure6-1.png", + "caption": "FIG. 6. Magnetization vector distribution of the sample rotor. The arrow is the direction of a magnetic vector, M. M denotes every 0.5\u00b0 in the mechanical angle and symbol M denotes the specific anisotropy angle to the tangential direction.", + "texts": [ + " The compacting pressure dependency of BH max of the above-mentioned magnets is shown in Fig. 5. At the compacting pressure of approximately 50 MPa, the BH max value did not deteriorate before and after the heating for cross-linking reaction of binder. It was found that optimized magnetic properties are 155\u2013158 kJ/m3 in BH max, 0.95\u20130.96 T in remanence, 0.90\u20130.91 MA/m in coercivity, and 6.1 mg/m3 in density, respectively. We, further, evaluated the controllability on the continuously anisotropy directions of a prepared eight-pole magnet rotor. Figure 6 shows the magnetization vector distribution of the rotor. Furthermore, Figs. 7 a and 7 b show the distribution of M and H in the boundary region of two magnets a , [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: 129.49.23.145 On: Thu, 18 Dec 2014 14:23:39 together with M of other conventional magnet rotors b for comparison. The two magnets were connected at =45\u00b0, and H jumps by 10\u00b0 at =45\u00b0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000360_4-431-27901-6_2-Figure2.4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000360_4-431-27901-6_2-Figure2.4-1.png", + "caption": "Fig. 2.4a-c. Flow velocity distribution. a shear flow, b pressure flow, c summation of shear flow and pressure flow", + "texts": [ + " The boundary conditions for the velocities are, from the assumption that there is no slip between the fluid and the solid surface (assumption 7), as follows: u = U1, w = W1 at y = 0 u = U2, w = W2 at y = h } (2.9) Then the fluid velocities will be as follows: u = \u2212 1 2\u00b5 \u2202p \u2202x y(h \u2212 y) + [( 1 \u2212 y h ) U1 + y h U2 ] (2.10) w = \u2212 1 2\u00b5 \u2202p \u2202z y(h \u2212 y) + [( 1 \u2212 y h ) W1 + y h W2 ] (2.11) where, in the calculations, it is assumed that the pressure p is constant in the y direction (assumption 5). In Eq. 2.10 for the flow velocity u, the latter half of the right-hand side (in brackets) shows the fluid velocity due to the movement of the solid surface in the x direction. It changes linearly as shown in Fig. 2.4a (it is assumed that U2 = 0). This is called shear flow or Couette flow. The former half of the right-hand side shows the flow velocity due to the pressure gradient. It is proportional to the pressure and changes parabolically across the film thickness as shown in Fig. 2.4b. This is called pressure flow or Poiseuille flow. The flow velocity in a general case is the sum of the two. Figure 2.4c shows such an example in which a flow in the reverse direction to the shearing direction occurs due to the pressure gradient at the left end and the shear flow is accelerated by the negative pressure gradient at the right end. At the point of maximum pressure, we have the relation dp/dx = 0, and therefore the flow at that point consists of the shear flow only. The same can be said of Eq. 2.11 for the flow velocity w. d. Continuity Equation The continuity equation for a small volume element in an incompressible fluid (assumption 3) can be written as follows: \u2202u \u2202x + \u2202 \u2202y + \u2202w \u2202z = 0 (2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002881_s42235-021-0043-x-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002881_s42235-021-0043-x-Figure3-1.png", + "caption": "Fig. 3 Design of passive wing pitch.", + "texts": [ + " With the right wing being taken as an example to establish the model of oscillating motion, which can also be extended to the left wing in a straightforward manner, the wing displacement, X(t) is calculated as: 2 2 2 ( ) = cos(2\u03c0 ) + 1 sin (2\u03c0 ). A X t A ft ft L (1) The wing velocity, v(t), is then calculated as: 2 2 2 2 2 2\u03c0 cos(2\u03c0 )sin (2\u03c0 ) ( ) 2\u03c0 sin(2\u03c0 ) , 1 sin (2\u03c0 ) A f ft ft v t Af ft A L ft L (2) where A is the radius of the crank rotation, f is the frequency of the crank rotation, and L is the coupler length. Fig. 3 shows the design of a passive wing pitch that controls the angle of attack passively during the movement. The parts connecting the wings and sliders made of resin material ensure that the wings only pitch within a certain angle and maintain a certain angle of attack through the limit structure. When the slider starts moving from a static state, the wings deflect under the resistance and inertia of the wind, causing the angle of attack to vary. When the limit structures of the connector and sliders are in contact, the wing angle of attack will no longer vary and the wings will move horizontally under the set angle of attack to generate lift" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002135_s00202-020-01163-8-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002135_s00202-020-01163-8-Figure8-1.png", + "caption": "Fig. 8 (a) The structure of the proposed TSCMG and (b) conventional structure of a multi-speed magnetic gear", + "texts": [ + " Therefore, the ultimate and optimal material of the TSCGM is shown in Fig.\u00a05. Also, Figs.\u00a06 and 7 depict the magnetic model of the maximum flux density with the aluminum middle rotor and its distributed flux density, respectively. Referring to Fig.\u00a06, the maximum flux density of the TSCMG with a middle aluminum rotor is 2\u00a0T which is suitable. Also, Fig.\u00a07 illustrates that the flux density in the TSCGM with a middle aluminum rotor is very desired and appropriate for distribution. A multi-speed gear with a conventional structure is presented in Fig.\u00a08(b). After conducting thorough studies, it has been clarified that this type of gear with such a structure in which iron is positioned in a dual-layer configuration in the middle rotor provides low efficiency because of the inappropriate design of the middle rotor. The present study proposes a TSCMG with a different structure for the middle rotor, which is more optimal and superior than the one detailed in Fig.\u00a08(b). As is illustrated in Fig.\u00a08, bot gears have been designed using the same volume and material. The only difference between them is the material and structure of the middle rotor because in the gear given in Fig.\u00a08(b) the middle rotor must be made from iron so that the flux density is closed in the inner and outer rotors. According to Fig.\u00a08, the thickness of the core of both middle rotors is 2\u00a0mm. The thickness of irons in both gears is also 2\u00a0mm. Table\u00a02 tabulates the optimal parameters of the magnetic gear given in Fig.\u00a08(b). To better clarify the torque and operation of the two gears discussed above, they were simulated using the FEM. Based on Fig.\u00a09, the maximum torques produced by the gear given in Fig.\u00a08(a) in the inner, middle and outer rotors are 13%, 24% and 10% larger than those of the gear given in Fig.\u00a08(b). The torque ripples in the inner, middle and outer rotors are 1 3 1 3 3, 0.9 and 2\u00a0Nm in Fig.\u00a08(a), respectively. Also, the torque ripples of the inner, middle and outer rotors are 4, 1 and 3\u00a0Nm in Fig.\u00a08(b). As a result, the torque ripples of the inner, middle and outer rotors of the gearbox are better in Fig.\u00a08(a) by 33, 11 and 50%, respectively. In addition, the flux density in the gear given in Fig.\u00a08(a) is distributed more optimal and appropriately. As observed in Fig.\u00a02, using a middle iron rotor will increase the flux density because a thicker core is required, which results in smaller modulators and decreased efficiency of the gear. This law is true for the gear given in Fig.\u00a08(b) as well because the flux density of the middle rotor in the gear is increased. As Fig.\u00a010 shows, the flux density in the gear given in Fig.\u00a08(b) is high. To reduce the flux density, the thickness of the middle rotor needs to be increased, resulting in reduced torque in the magnetic gear. Thus, using the conducted studies, the structure proposed in this paper provides distributed flux density, higher efficiency of the output torque and better performance. Therefore, we can infer that the gear structure Fig. 5 The structure of the proposed TSCMG 1 3 given in Fig.\u00a08(b) cannot be assumed as a proper structure for multi-speed gears. According to the analysis conducted in the previous section, a prototype of the proposed TSCMG was implemented to verify the simulation results. The gear consists of inner, middle and outer rotors, where the middle rotor is designed heuristically using a belt shaft so that speed and torque can be extracted from this rotor. Figure\u00a011 shows the output of the middle rotor belt. The important point to notice is the parameters of the middle rotor, which play a key role in the torque of the inner and outer rotors", + " Based on these characteristics, the TSCMG was simulated. Then, the TSCM was evaluated using the FEM under different operating conditions. Magnetic flux density and maximum flux were measured on rotors. The obtained results show that the flux density in the TSCMG is 2\u00a0T. Additionally, the FEM results proved that the TSCMG provides larger torque compared to conventional multi-speed gears. The main reason for the increased torque and suitable flux density in the Fig. 9 The maximum torque produced by the TSCMG given in Fig.\u00a08 (a), (b) 1 3 TSCMG is, in fact, the appropriate design of the middle rotor in a single-layer form. The proposed magnetic gear was built for experimental validation and then was tested. The results obtained from the tests highlight the fact that the torque in the implemented TSCMG has a negligible difference with simulation results. The error is roughly 4\u20137%. The results also approves that for the applications like permanent wind turbine, the TSCMG is a powerful equipment to produce an electricity from the wind and at the same time produce a wind using third (middle rotor)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002862_09544062211016076-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002862_09544062211016076-Figure4-1.png", + "caption": "Figure 4. Schematic drawing of the screw analysis of a branch chain.", + "texts": [ + " There are five motion DOF of passive joints and one motion DOF of actuator on each branch chain. Obviously, the parallel mechanism is a complete constraint mechanism and thus the internal force could be left out of account and external wrench distributed to each branch chain could be obtained using projectors. Considering that motion DOF of passive joints is larger than 3, the method based on the kinematic screw analysis would be adopted here to obtain matrix U2i i \u00bc 1; . . . ; 6\u00f0 \u00de. Take the i-th branch chain as an example (as shown in Figure 4), the form of equivalent kinematic pair screw system of passive joints can be expressed as $i1 \u00bc Cg i2 Sg i3 Cg i2 Cg i3 Sg i2 ; hi1Cg i2 Cg i3 hi1Cg i2 Sg i3 0 $i2 \u00bc Cg i3 Sg i3 0; 0 0 0 $i3 \u00bc 0 0 1; 0 0 0 $i4 \u00bc 0 1 0; hi2 0 0 $i5 \u00bc Cg i4 0 Sg i4 ; 0 hi2 \u00fe hi3Cg i4 0 8>>>>< >>>>: (48) where C \u00f0 \u00de and S \u00f0 \u00de respectively represent sine and cosine, and the definition of the parameters could be found in Figure 4. Observing the kinematic screw system, it is obvious that one of the constraint wrenches of the i-th branch chain is $ci \u00bc ~nTi ; ~ri ~ni T \u00bc 0 0 1; 0 0 0 ; i \u00bc 1; . . . ; 6\u00f0 \u00de (49) where ~ni and ~ri respectively represent the direction of the i-th rod and the position vector of a point on the central axe of the i-th rod which could be arbitrarily chosen. The left part of the formula (49) depends only on the topology of the chain and thus holds whatever the definition of coordinate system are. In addition, considering that five twists of the kinematic screw system are always linear independent, the rank of the constraint wrench system is 1 and thus matrix U2i is expressed as U2i \u00bc 1 ~ni ~ri ~ni ~ni ~ri ~ni \u00bc 1 jj~fi jj ~fi ; i \u00bc 1; " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002465_mems51782.2021.9375221-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002465_mems51782.2021.9375221-Figure1-1.png", + "caption": "Figure 1: (a) The structure of TB sensor. (b) The sensor-integrated smart manipulator. (c) The structure of TT sensor. (d) The integrated intelligent system.", + "texts": [ + " Moreover, One-dimensional (1D) convolutional neural network (CNN), capable of extracting useful information from sensor\u2019s original time-domain signal automatically for further processing and classification, has been proven as an efficient ML tool to deal with the triboelectric spectrum with high accuracy [18]. The subtle features hidden in the triboelectric waveform, including contact sequence, collision vibration, etc., other than the individual physical information, effectively enhance the perception and recognition capability of the integrated intelligent system. Herein, we integrate a soft-robotic manipulator with a TB sensor (Figure 1a) for finger bending monitoring and a TT sensor (Finger 1c) for contact position and area detecting. The TT sensor made of silicone rubber shows good flexibility, which is able to fit and follow the pneumatic fingers well under different deformation conditions. The TB sensor measures the bending angle by counting output peaks, capable of avoiding environmental influences (humidity, temperature, etc.). By combining the sensory information both from the TB and TT sensor through the IoT module (Figure 1d), ML enhanced data analytic can be leveraged to help the manipulator realize more complex perception functions, such as the recognition of grasping objects, illustrating the potential of the integrated system in unmanned shop, automatic assembly line and assistive robot applications. Working Mechanism of Sensors for soft manipulator The working mechanisms of TB and TT sensors are shown in Figure 1a and Figure 1c respectively. The gear of the TB sensor is covered with a layer of Nickle fabric (Ni-fabric) as positive triboelectric material. When the pneumatic actuator is inflated, the strip which connected to the gear will be stretched, making gear rotate and gear teeth contact with the negative polytetrafluoroethylene (PTFE) film intermittently. Due to the triboelectrification on contact surface and the electric potential changes in Cu electrode, triboelectric output peaks will be generated as depicted in Figure 2a, where the peak numbers represent the bending angle of the pneumatic finger (30 degrees/peak) and the resolution can be further improved with increased gear size and teeth numbers. Besides, the disc spring mounted on the shaft is able to provide the pull-back force for the pneumatic finger to recover when air goes out. Next, the TT sensor is designed to collect the tactile related sensory information when the pneumatic finger contacts with external objects and the detailed structure is shown in Figure 1c. There are three short electrodes (labeled as E1-E3) and one long electrode (labeled as EL) patterned on the TPU substrate, with a layer of silicone covered on the top serving as the negative triboelectric material. When silicone layer contact with other objects, it tends to attract negative charges held on its surface, thus generating triboelectric outputs in attached 4 electrodes due to electrostatic induction. The different locations of the short electrodes (E1-E3) on patch surface will result in different output amplitudes in these channels during contacting as shown in Figure 2b" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001567_micronano2008-70185-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001567_micronano2008-70185-Figure1-1.png", + "caption": "Fig. 1 Schematic of two types of a S-DGS", + "texts": [ + " [9, 10] in the form o open open open z o oh h F F F K h h h= \u2032\u2202 \u2212 = \u2248 \u2212 \u2202 \u2032 \u2212 (7) where ho is the film separation after a small film perturbation, F\u2032open and Fopen are the opening force corresponding to ho and ho, respectively. The leakage rate of sealed medium through the interface is 2 3 2 2 2i o g i0 g i 1 ( 2 ) 24 ln( / ) p hQ P d P T R R \u03c0 \u03b8 \u03c0 \u00b5 = \u2212 \u2212 \u211c \u222b (8) The maximum fluid film stiffness to leakage ratio is used here to be the geometric optimization principle for a LST-MS according to Refs.[9, 10] zK Q \u0393 = (9) RESULTS AND DISCUSSION Comparisons Here we analyze the solution for both of S-DGSs with or without an inner annular groove shown in Fig. 1(a, b). First the Copyright \u00a9 2008 by ASME Terms of Use: http://www.asme.org/about-asme/terms-of-use Dow comparisons of sealing performance are made between a typical S-DGS and a S-DGS with inner annular groove at the same order of groove depth as spiral groove. The seal parameters are shown in Table 1. As shown in Table 2, compared with a typical S-DGS, a S-DGS with an annular groove exhibits a nearly unchanged value of opening force, but exhibits an increased value of leakage and film stiffness" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002274_j.precisioneng.2021.01.004-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002274_j.precisioneng.2021.01.004-Figure6-1.png", + "caption": "Fig. 6. Flow of hopping in stance phase.", + "texts": [ + " The amplitude A and the phase \u03c6 of the complex variable \u03be correspond to the energy and the phase of the spring-mass oscillator. Y. Abe and S. Katsura Precision Engineering 69 (2021) 36\u201347 De et al. [15] proposed an oscillation method based on the phase of the spring-mass oscillator. The method makes the actuation phase to be equal to the vibration phase of the spring-mass oscillator. The phase of the oscillator is induced as (28). Force reference Fref z is designed in order to synchronize with the phase of the oscillator as Fref z =Fref A cos\u03c6, (29) where Fref A is a positive constant value. Fig. 6 shows the procedure in the stance phase of the hopping motion. A new phase variable \u03b8 is introduced for clear understanding of the phase. The variable \u03b8 is defined as \u03b8=\u03c6 \u2212 \u03c0. (30) The phase variable \u03b8 moves from 0 to \u03c0 in the stance phase while \u03c6 moves from \u03c0 to 2\u03c0. \u03b8 is more clear in terms of that the phase starts from 0 and ends at \u03c0. There are three main events in the stance phase as Fig. 6: touch down, bottom, and lift off. Touch down is the moment when the robot in the air touches the ground. At this moment, \u03b8 is near zero. From (29) and (30), the force reference becomes the negative maximum value at the touch down. The bottom is when the spring is compressed maximum and the mass velocity becomes zero. At this time, \u03b8 becomes \u03c02. From (29) and (30), the force reference is zero in this case. Liftoff is the moment when the robot flies from the ground. Liftoff is set when the velocity of the robot body becomes the maximum" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003163_j.cja.2021.05.023-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003163_j.cja.2021.05.023-Figure1-1.png", + "caption": "Fig. 1 Structure of flexible manipulator.", + "texts": [ + " The main results are presented in Section 3. In Section 3.1, a High-Gain Observer (HGO) is developed. In Section 3.2, a novel Spike Suppression Function (SSF) is developed, and then a dynamic surface controller based on HGO and SSF is proposed. In Section 4, a simulation verifies the effectiveness of the proposed HGO, SSF and controller. In Section 5, the conclusion of the whole paper is presented. 2. Problem formulation 2.1. Dynamic model of flexible joint manipulator The flexible joint manipulator is shown schematically in Fig. 1. The dynamic model of the Flexible Manipulator Control System (FMCS) can be described as follows34: I\u00fe DI\u00f0 \u00de\u20acq\u00fe K\u00fe DK\u00f0 \u00de q h\u00f0 \u00de \u00fe MgL\u00fe DMgL\u00f0 \u00desin q\u00f0 \u00de \u00fe d1 \u00bc J\u00fe DJ\u00f0 \u00de\u20ach K\u00fe DK\u00f0 \u00de q h\u00f0 \u00de \u00fe d2 \u00bc s \u00f01\u00de where q and _q are the rotation angle and angular velocity of link, respectively; h and _h are the rotation angle and angular t d P r of r flexible joint manipulator, Chin J Aeronaut (2021), https://doi.org/10.1016/j. 160 161 162 163 164 165 166 167 168 169 171 172 173 175 176 177 178 179 181 182 183 185 186 187 188 189 190 191 192 193 194 195 196 197 198 199 200 201 202 203 204 205 206 207 208 210 211 213 214 215 216 217 218 220 221 223 224 225 226 228 229 230 231 232 233 235 236 237 238 239 240 241 242 243 244 245 246 247 248 249 250 251 252 253 254 255 256 257 258 Spiking-free HGO-based DSC for flexible joint manipulator 3 c velocity of motor, respectively; I and J are the moment of inertia of link and motor, respectively; Mand L represent the mass and the distance between the joint and the centroid of the link, respectively; g is the gravity acceleration; K represents the elastic stiffness of the flexible joint; s is the motor torque; DI, DK, DJ and DMgL represent the unmodeled dynamics of the FMCS; d1 and d2 denote the disturbance from the environment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001266_s12206-007-1104-8-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001266_s12206-007-1104-8-Figure2-1.png", + "caption": "Fig. 2. Schematic diagram and coordinates of a spinning disk.", + "texts": [ + " 1(b), in which springs intervene between disks. Since the head assembly is of stiff foundation, the coupling among disks, to the authors\u2019 opinion, is relatively weak, the same remarks as made by Orgun and Tonque [13, 14]. The present system is similar to the study conducted by Orgun and Tonque [13] but by a thoroughly different approach. In addition, unlike that of Orgun and Tonque [13], the present study focuses on the vibration characteristics due to coupling rather than the localization due to thickness disorder. Fig. 2 shows a spinning disk, in which (X,Y,Z) is an inertia frame; a, the outer radius; b the inner radius; , the spinning speed; , the angular coordinate with respect to the inertia frame. The equation of motion of a spinning disk in terms of its transverse displacement u (r, t ) with respect to a stationary observer was derived [15] 2 2 2 2 2 2 1( ) [ ( ) ( )]r D u r r r r h u ur r r r r 2 2 2 2 2 2[ 2 ] ( , , ),u u uh p r t t t (1) where D=Eh3/12 (1- 2), the bending rigidity; , den- sity; h, thickness; , Poisson\u2019s ratio; p pressure type loading; r and , initial stresses due to spin, 2 2 2 2 2 2 2 2 2 2 2 (3 )( ) ( ) 8 (1 )[ (3 ) (1 )] ( 1) 8[ (1 ) (1 )] r r a r b a b a b a r (2) 2 2 2 2 2 2 2 2 2 2 2 ( ) [(3 ) (1 3 ) ] 8 (1 )[ (3 ) (1 )] ( 1) 8[ (1 ) (1 )] r a r b a b a b a r (3) The disk is assumed clamped inside and free outside, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002530_cerma.2007.4367740-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002530_cerma.2007.4367740-Figure3-1.png", + "caption": "Figure 3. Distance calculation between the tractor and a trailer", + "texts": [ + " The core of the constraint method is the use of a damping constraint, given by the next expression: d\u0307 \u2265 \u2212\u03be d \u2212 ds di \u2212 ds if d \u2264 di (11) with: \u2022 d the minimal distance between two objects, \u2022 ds the safety distance, \u2022 di the influence distance and \u2022 \u03be the positive damping coefficient In our previous work [14], we have shown that the con- straint imposed by an obstacle over the tractor is given by: (n \u00b7 m)vi + ( ~OP \u00d7 n) \u00b7 z\u0302 \u03c9i \u2264 \u03be d \u2212 ds di \u2212 ds (12) where vi and \u03c9i are expressed in function of tractor velocities, m is the unit vector along the robot X axis and n is the unit vector along the segment ~PrPt of minimal distance between the modules. For a tractor and one trailer as they are shown in figure 3 we can see that the linear velocity for the tractor and its trailer are the same in point l and we can say that: A = At + Ar B = Bt + Br where: At = mt \u00b7 nt Ar = mr \u00b7 nr Bt = ( ~OPt \u00d7 nt) \u00b7 z\u0302 Br = ( ~OPr \u00d7 nr) \u00b7 z\u0302 with mt and mr the unit vectors of axes Xi\u22121 et Xi respectively, z\u0302 represents the unit vector perpendicular to plan and nt represents the unit vector across the segment PtPr, nr is its negative vector. We can rewrite the anti-collision restriction for a robot with a trailer as: (mt \u00b7 nt + mt \u00b7 nt)v + [( ~OPt \u00d7 nt) + ( ~OPr \u00d7 nr)] \u00b7 z\u0302\u03c9 \u2264 \u03be d\u2212ds di\u2212ds (13) We can conclude that an anti-collision constraint is represented by a linear constraint on the tractor\u2019s velocities" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001726_isam.2007.4288440-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001726_isam.2007.4288440-Figure6-1.png", + "caption": "Fig. 6. Example 1", + "texts": [ + " The following set of basis functions are used in the computation of these examples: e1 = 1, e2 = u, e2i+1 = sin \u03c0iu U , e2i+2 = cos \u03c0iu U , (i = 1, 2, 3, 4). (39) Necessary constraints for developability described by eqs.(17), (22), and (23) are divided into 16 conditions at point P(iU/15) (i = 0, \u00b7 \u00b7 \u00b7 , 15) respectively in the following examples. All computations were performed on a 750MHz Alpha 21264 CPU with 512MB memory operated by Tru64UNIX. Programs were compiled by a Compaq C Compiler V6.1 with optimization option -O4. Fig.6 shows the first example of straight object deformation. The length of the object U is equal to 1, its width V is equal to 0.1, and its flexural rigidity along the spine line Rf is constantly equal to 1. In this example, both ends of the spine line are on the same line but directions of the spine line at these points are different. Fig.7 shows computational results. Fig.7-(a), -(b), and -(c) illustrate the top, front, and side view of the object, respectively. As shown in this figure, the object satisfies the given geometric constraints by twisting partially" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003436_j.apacoust.2021.108345-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003436_j.apacoust.2021.108345-Figure1-1.png", + "caption": "Fig. 1. Gear pair dynamic model.", + "texts": [ + " The sound prediction model permits the selection of the gear combination according to the detectability regulations. The implementation of the sound prediction model is divided into two consecutive parts clearly differentiated. On the one hand, a dynamic model that studies the motion of the driven wheel to determine its rattle movement has been developed. On the other hand, the acoustic generation model, which successively produces the noise of the gear teeth that are impacting during the rattle movement has been presented. The scheme of the free gear system is shown in Fig. 1. The driven wheel is mounted on a ball bearing. The motion of the driving wheel u1 is imposed so the system has only one rotational freedom degree u2 in the driven wheel. The equation of motion for the driven wheel is shown in Eq. (1): I2 \u20acu2 \u00bc FgRb2 FtRq2 Ff Rf \u00f01\u00de where I2 is the inertia of the driven wheel,\u20acu2 is the angular acceleration of the wheel. Fg is the gear meshing force while its lever arm is the base radius Rb2. Ff is the frictional force in the bearing which is the only charge in the loose gear pair and Rf is the application radius of the force", + " In the present study, the following shows how the noise produced by the impact of a pair of dry teeth is obtained. To simulate the clash of the teeth, the gears are simplified as a system of two impacting cylinders [25] which vary in time depending on the gear mesh position. Each radius of these cylinders is given using Pythagoras theorem in Eq. (2). Rci \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi Rbi 2 \u00fe Rqi 2 q \u00f02\u00de where Rbi is the base radius of the pinion or the wheel, and Rqi is the involute curvature radius as it is shown in Fig. 1. To reduce computational cost and simplify the programming, the sound produced by a pair of impacting cylinders can be equated as an equivalent cylinder impacting against a semi-infinite elastic half-space [25]. The equivalent radius is given by Eq. (3) where Rc1 and Rc2 are the radii of each cylinder in any instant. Rceq \u00bc Rc1Rc2 Rc1 \u00fe Rc2 \u00f03\u00de The mass of the equivalent cylinder depends on the flank width l and the material density q as is shown in Eq. (4). meq \u00bc pRceq2lq \u00f04\u00de k1 and k2 parameters are used to simplify later notation as it is described in [26], where the clash of two cylinders is presented" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002560_tec.2021.3069096-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002560_tec.2021.3069096-Figure1-1.png", + "caption": "Fig. 1 (a) A cross sectional view of the original 12-slot and 14-pole PMSM (b) Added 12-slot and 10-pole RB. saliency can be well reflected onto the stator winding.", + "texts": [ + " fundamental performance. Although the rotor modification widens the feasible region of self-sensing operation, the saturation effects have not been completely eliminated. It is still observed that the saliency ratio of ringed-pole machine decreases at heavy loads and cross-saturation phase shift is still very large for field-intensified IPM (FIIPM) [16]. Modification of a rotor structure has been proposed in [17] to enhance the sensorless capability of PMSMs against the machine saturation effect. Fig. 1(b) illustrates the additional anisotropy provided by a 10-pole RB attached to one-end of a 14-pole rotor (Fig. 1 (a)). The 12-slot stator winding provides magnetic excitation while the saliency modulation of the RB (SMRB) generates reluctance variations in the process of the saliency modulation. The major contribution of this paper compared to [17] is reflected in the following aspects: 1) This work is developed on top of the concept raised in [17]. Considerable efforts are made to deeply verify the feasibility of sensorless operations and further enhance the performance of closed-loop sensorless control of PMSM" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001994_b978-0-240-80969-4.50014-6-Figure10-10-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001994_b978-0-240-80969-4.50014-6-Figure10-10-1.png", + "caption": "Figure 10-10. Three terminal potentiometer.", + "texts": [ + " RTDs can self heat, causing inaccurate readings, therefore the current through the unit should be kept to 1 mA or less. Self heating can also be controlled by using a 10% duty cycle rather than constant bias or by using an extremely low bias which can reduce the SNR. The connection leads may cause errors if they are long due to the wire resistance. Potentiometers and Rheostats. The r e s i s t ance o f potentiometers (pots), and rheostats is varied by mechanically varying the size of the resistor. They are normally three terminal devices, two ends and one wiper, Fig. 10-10. By varying the position of the wiper, the resistance between either end and the wiper changes. Potentiometers may be wirewound or nonwirewound. The nonwirewound resistors usually have either a carbon or a conductive plastic coating. Potentiometers or pots may be 300\u00b0 single turn or multiple turn, the most common being 1080\u00b0 three turn and 3600\u00b0 ten turn. Wirewound pots offer TCRs of 50 ppm/\u00b0C and tolerances of 5%. Resistive values are typically 10 \u2013100 k , with power ratings from 1 W to 200 W. Carbon pots have TCRs of 400 ppm/\u00b0C to 800 ppm/\u00b0C and tolerances of 20%" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001408_detc2008-50033-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001408_detc2008-50033-Figure1-1.png", + "caption": "Figure 1 A Free Body Diagram Of The Sprung Mass Of A TwoAxled Vehicle In Steady State Cornering [6]", + "texts": [ + " These data are essentially the same as the data required for the customary method of steady state handling analysis outlined in [2]. 2) Weight shift submodel. This submodel computes the lateral weight shift using vehicle parameters listed above for a given 2 Copyright \u00a9 2008 by ASME e: http://www.asme.org/about-asme/terms-of-use Down value of lateral acceleration. This submodel, also discussed in [2] allows for the calculation of the roll angle, roll rate, wheel normal forces and per axle lateral forces corresponding to the given lateral acceleration. The computation is based on the model shown in Fig.1. The three primary contributions to lateral weight shift [2, 6] are listed here briefly: 2.1. Weight shift proportional to roll center height and track width as given by: \u0394w T T a (2) 2.2 Weight shift proportional to product of roll rate and roll stiffness as given by: loaded From: https://proceedings.asmedigitalcollection.asme.org on 02/07/2019 Terms of Use \u0394w K K K \u2013 H H T a (3) where H is the height of the CG of sprung mass above the roll axis, as shown in Fig.1, is computed from: H h h l L (4) 2.3. Weight shift proportional to CG height of unsprung mass as given by: \u0394w T a (5) The total lateral weight shift for the front tires is given by: \u0394w \u0394w \u0394w \u0394w (6) Proceeding similarly, the total lateral weight shift for the rear tires is given by: \u0394w T T a K K K \u2013 H H T a T a (7) Equations (2-7) are obtained from the equations of motion for the sprung mass and the unsprung masses, as shown in Figs 1 and 2, respectively. The inside and outside normal loads are computed from: L \u2013 \u0394w (8) L \u0394w (9) L \u0394w (10) L \u0394w (11) The roll gain or roll rate (in deg/g) of the sprung mass is [2]: K H K K \u2013 H \u03c0 (12) The roll angle (in deg) for a given lateral acceleration is then, given by \u03c6 K a (13) The front and rear axle lateral forces are given by lat w a (14) lat w a (15) 3) Tire submodel" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003062_14644207211026696-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003062_14644207211026696-Figure4-1.png", + "caption": "Figure 4. Test apparatus used in experiments.", + "texts": [ + " Information on the gears tested is given in Table 1. The tested gears made of AISI 4140 were heated to 860 C by 1 h in the furnace, and quenched in oil. Then, gears were tempered at 400 C by 2 h to find out the hardness level of 38 HRC. The test apparatus was developed to perform the STBFTs of gears more functionally and to test gears with different geometries with a single test apparatus. The elements of the test apparatus, which is the revised version of Patent Pending Design,26 are shown in Figure 4. Important elements and parts of the test apparatus designed for experimental study are numbered, and these numbers are explained below: 1. Test apparatus lower table; 2. Linear motion plate adjustment bolt; 3. Meshing gear tooth; 4. Linear motion plate; 5. Force transmission pin; 6. Linear slide; 7. Test apparatus vertical table; 8. Gear fixing apparatus slide cover; 9. Gear fixing apparatus adjustment bolt; 10. Gear fixing apparatus slide; 11. Gear fixing apparatus; 12. Test gear; 13. Gear fixing bolt; 14" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002945_978-3-030-69178-3_15-Figure6.1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002945_978-3-030-69178-3_15-Figure6.1-1.png", + "caption": "Fig. 6.1 A cell layout in CAD and modelled in the SSConf software", + "texts": [ + " The order is sent to the Master hardware device controlling all the certified safety devices. TheMaster hardware controls the change and synchronisation of safety configurations in the devices. TheMaster hardware functions as the interface between certified safety devices and the DSC control. The Master hardware is typically a safety-rated Programmable Logic Controller (PLC) that signals the active configuration to all safety devices. The configuration of the system begins with modelling the layout of the robot cell using the Safety System Configurator. The model shown in Fig. 6.1 is a simplified replication of the real cell built by placing the robot, safety devices, tables, and other static appliances. The layout needs to include the areas where the robot will move with the objects that will be present in the scan area of the safety sensors (e.g., table legs). In the configuration software, the robot is represented in a simplified way, with the safety parameters required for separation distance calculation. As shown in Fig. 6.2, the software has a library of pre-configured robot models and an easy way to include more by modifying a template" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003572_s10015-021-00694-y-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003572_s10015-021-00694-y-Figure5-1.png", + "caption": "Fig. 5 Crawler mechanism for each pair", + "texts": [ + "\u00a0(2) and (3), respectively: where Fni is contact force at crawler i. The six-crawlers are devided to 3 pairs. All pairs are driven by a motor (24\u00a0V 40\u00a0W 57.8\u00a0mNm 9690\u00a0rpm), the power is transfered through the worm gear and chain mechanism (gear ratio 213.28:1) to drive all crawlers. The crawlers have a driving speed about 7.06\u00a0cm/s and a driving torque about 12.3Nm to overcome the obstacles inside the sewer. The crawlers are made from double-side timing belt that are installed on the four-bar linkages, as shown in Fig.\u00a05. The dash lines stand for the chains that connects each crawler to the motor. (1)Dr = Do + 2l sin . (2)Fn2 = Fn3 = mg, Fn1 = 0 (3)Fn2 = Fn3 = mg This section consists of a 3-DOF manipulator with cutting tool and a camera. The tools are used for removing the blockages and drilling the holes. The camera is setup at the manipulator, above the tool. The viewpoint covers the tool, the targets, and the area in front of robot 45\u00a0cm. The viewpoint can be changed by turning the manipulator. The workspace of the manipulator covers all targets area using only 2 rotation joints (J1, J2) and 1 linear joint (J3)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002838_012021-Figure25-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002838_012021-Figure25-1.png", + "caption": "Figure 25. Stress Distribution in Carbon Fiber", + "texts": [ + " Analysing Testing Result of Carbon Fiber 3.7.1. Total Deformation The Max. And Min. Total Deformation in Carbon Fiber is 0.68523 mm and 0 mm respectively shown in Figure 24. The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.7.2. Stress Distribution The Max. And Min. Stress Distribution in Carbon Fiber is 97.378 MPa and 0.28227 MPa respectively shown in Figure 25. 3.7.3. Strain Distribution The Max. And Min. Strain Distribution in Carbon Fiber is 0.0024296 and 0.000021423 respectively shown in Figure 26. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.8. Analysing Testing Result of HSLA Steel 3.8.1. Total Deformation The Max. And Min. Total Deformation in HSLA Steel is 0.19371 mm and 0 mm respectively shown in Figure 27. The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002438_012190-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002438_012190-Figure4-1.png", + "caption": "Figure 4. Case 2 of the domain, region and sensor.", + "texts": [ + " But Iraqi Academics Syndicate International Conference for Pure and Applied Sciences Journal of Physics: Conference Series 1818 (2021) 012190 IOP Publishing doi:10.1088/1742-6596/1818/1/012190 \ud835\udcb4(\ud835\udcc9) = \u2329\ud835\udc531, \ud835\udf11\ud835\udc560\ud835\udc570\u232a = ( 4 (\ud835\udefd1\u2212\ud835\udefc1) (\ud835\udefd2\u2212\ud835\udefc2) ) 1 2\u2044 \u222b \u222b \ud835\udc531( \ud835\udefc1+\ud835\udf011 \ud835\udefc1\u2212\ud835\udf011 \ud835\udefc2+\ud835\udf012 \ud835\udefc2\u2212\ud835\udf012 \ud835\udf011, \ud835\udf012) \ud835\udc46\ud835\udc56\ud835\udc5b [ \ud835\udc570 \ud835\udf0b \ud835\udf011 (\ud835\udefd1\u2212 \ud835\udefc1) ] \ud835\udc46\ud835\udc56\ud835\udc5b [ \ud835\udc570\ud835\udf0b \ud835\udf012 (\ud835\udefd2\u2212 \ud835\udefc2) ] \ud835\udc51\ud835\udf011 \ud835\udc51\ud835\udf012 If \ud835\udc53 satisfy symmetry property for \ud835\udccd01 = \ud835\udf01, this implies \ud835\udcb4( \ud835\udc61) = \u2329\ud835\udc531, \ud835\udf11\ud835\udc560\ud835\udc570\u232a = 0. Therefore, the dynamic system (31) is not \ud835\udc34\ud835\udc45\ud835\udc3a\ud835\udc45\ud835\udc42-observer for the systems (26)-(30).\u25a1 5.2. Case 2 Deliberate case 2 in circular supports which is shown in Fig.4., and described by equations (26)-(32). So the measuring output is stated by \ud835\udcb4(\ud835\udcc9) = \u222b\ud835\udc37\ud835\udc651(\ud835\udc5f, \ud835\udf03, \ud835\udc61)\ud835\udc53(\ud835\udc5f, \ud835\udf03)\ud835\udc51\ud835\udf03, (32) where \ud835\udc37 = (\ud835\udc5f, \ud835\udf03) \u2282 \u2127, is the position of the using sensor \u201d Figure 4\u201d. So that (\ud835\udc37, \ud835\udc53) may be enough to ensure \ud835\udc34\ud835\udc45\ud835\udc3a\ud835\udc45\ud835\udc42-observer [8], and by employing the same way in case 1 implies the following lim \ud835\udc61\u2192\u221e \u2016(\ud835\udc64(\ud835\udc5f, \ud835\udf03, \ud835\udcc9) + \ud835\udc3b\ud835\udc34\ud835\udc45\ud835\udc3a\ud835\udccd2(\ud835\udc5f, \ud835\udf03, \ud835\udcc9)) \u2212 \ud835\udccd(\ud835\udc5f, \ud835\udf03, \ud835\udcc9)\u2016(\ud835\udc3b1(\ud835\udf14))\ud835\udc5b = 0, where { \ud835\udf15\ud835\udccc \ud835\udf15\ud835\udc61 (\ud835\udc5f, \ud835\udf03, \ud835\udcc9) = \ud835\udf152\ud835\udccc \ud835\udf15\ud835\udf012 (\ud835\udc5f, \ud835\udf03, \ud835\udcc9) + ((1 + \ud835\udc3b\ud835\udc34\ud835\udc45\ud835\udc3a) \ud835\udccc (\ud835\udc5f, \ud835\udf03, \ud835\udcc9) +(1 \u2212 \ud835\udc3b\ud835\udc34\ud835\udc45\ud835\udc3a) \ud835\udf15\ud835\udc652 \ud835\udf15\ud835\udf012 (\ud835\udc5f, \ud835\udf03, \ud835\udcc9) + (\ud835\udc3b\ud835\udc34\ud835\udc45\ud835\udc3a 2 \u2212 1) (\ud835\udc5f, \ud835\udf03, \ud835\udcc9)) \ud835\udcac \ud835\udccc (\ud835\udc5f, \ud835\udf03, 0) = \ud835\udccc0 (\ud835\udc5f, \ud835\udf03) \u2127 \ud835\udccc (\ud835\udc5f, \ud835\udf03, \ud835\udcc9) = 0 \u0398 (33) Then, we have the following result: Proposition 5.2. Supposing \ud835\udc37 = \ud835\udc37(\ud835\udc50, \ud835\udc5f) \u2282 \u2127, \ud835\udc50 = (\ud835\udc501, \ud835\udc502). So the process (33) is not \ud835\udc34\ud835\udc45\ud835\udc3a\ud835\udc45\ud835\udc42- observer for (26)-(32), if \u2200\ud835\udc560, \ud835\udc560(\ud835\udc5001 \u2212 \ud835\udefc1) (\ud835\udefd1 \u2212 \ud835\udefc1) and \ud835\udc560(\ud835\udc5002 \u2212 \ud835\udefc2) (\ud835\udefd2 \u2212 \ud835\udefc2) \u2044\u2044 is rational number in order \ud835\udc53 satisfy symmetry property around the \ud835\udc6501 = \ud835\udc5001" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001260_s021797920804750x-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001260_s021797920804750x-Figure6-1.png", + "caption": "Fig. 6. Test specimens (a) conventional, (b) novel. Fig. 7. Fabricated pilot model ((a) plaster, (b)metal powder).", + "texts": [ + " Figure 5 shows the UV curing system and right figure shows a piezo print-head unit with a heating chamber of resin. For a comparison of the two 3DP processes, the specimens were fabricated, and the fabricated specifications are as shown in Table 1. In the novel 3DP process, the printed resin was cured by using the UV lamp per each layer. According to the results, additional curing time is no longer required after completing all the layer printings. Furthermore, there is no need for post-processing. The non-cured powder has only to be eliminated by air spray. The fabricated specimen is shown in Figure 6(a). For the testing of the conventional 3DP process, The jetting fluid was Zb56(water based binder), which is also provided by Z-Corporation. The specimen was brush-painted with Z-bondlOl for the post-process and is shown in Figure 6(b). For a comparison of the fabricating time, the conventional process took approximately 70 minutes, which comprised the printing, drying, and post-processing times, as shown in Table 1, while the proposed process took 35 minutes, consisting of the printing time alone, without the drying and post-processing time required for the conventional process. Furthermore, the materials available for the conventional process were restricted to plaster-mixed and starch-mixed powders. However, in the proposed novel 3DP process, polymer and metal powder could also be used due to the strong coherence provided by UV curing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000368_detc2004-57064-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000368_detc2004-57064-Figure6-1.png", + "caption": "Figure 6. A Parallel Manipulator Without Relative Motion Between the Platforms and Without Self-Motion.", + "texts": [ + " More precisely, if the sum of all degrees of freedom of the kinematic pairs of any of the k serial connecting chains is greater that the dimension of the corresponding infinitesimal mechanical liaisons,V m f j , it follows that the parallel platform has passive degrees of freedom. The total number of these passive degrees of freedom is computed by Fp \u2211 fi k \u2211 j 1 dim Vm f j (39) 5 Examples In this final section, the results developed in this contribution will be applied to a series of interesting examples of parallel manipulators. 5.1 Fixed Platform Without Self-Motion. Figure 6 shows a parallel platform, in which there is no relative movement between the moving and fixed plataforms, this class of platforms was analyzed in Section 2.1. Each serial connecting leg is formed by three revolute pairs whose axes are parallel. Therefore, the closure algebra associated with each leg are given by Am k j ple\u0302 j ; 15 ownloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?u namely, they are the subalgebras associated with planar subgroups. Moreover, the set e\u03021 e\u03022 e\u03023 is linearly independent" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002331_012081-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002331_012081-Figure1-1.png", + "caption": "Figure 1. The Leader-Follower kinematic model.", + "texts": [ + " For the i-th UAV in a formation, it is assumed that the following two decoupled autopilots are in place: heading-hold autopilot, and velocity-hold autopilot. Considering that the UAV\u2019s altitude remains unchanged in flight, the model of the autopilot can be simplified as follows [22]. 1 ( ) c i i i V V V V (1) 1 ( ) c i i i (2) Where c iV and c i are the speed and heading change commands, V and are constant coefficients. The UAV velocity iV and rate change of heading i should be under the constraints of the flight envelope: min max max0 , i iV V V . Figure 1 shows the kinematic relations between the leader and the follower in formation flight where a planar situation is considered. The heading and position of the leader in the inertial reference frame, provided by its navigation system, are transmitted to the follower through wireless communication. So a body fixed frame centered at iF , the instantaneous position of the i-th follower, is defined. The axis ix is aligned with the instantaneous velocity vector and the axis iy is along iF starboard wing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002772_978-3-030-67750-3_14-Figure30.6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002772_978-3-030-67750-3_14-Figure30.6-1.png", + "caption": "Fig. 30.6 a Final conceptual design of wheelbarrow loader bucket. b Exploded view of the design", + "texts": [ + " Based on all the design steps taken before, three possible design sketches have been made to slightly give a picture to develop a possible design for this product. The simple 2D sketches equipped with the specifications and parts from the data collected in the previous design method process had been analyzed later by using a weighted objective analysis method. Therefore, the final design concept is available after finishing up all the designs steps. The final conceptual design was made by using SolidWorks, Fig. 30.6a displays the final conceptual design of the wheelbarrow loader bucket, while Fig. 30.6b is the exploded view of the design. The side view of wheelbarrow bucket loader is shown in Fig. 30.7. The design is based on the existing wheelbarrow but has a modification by using the six-point linkage based on the existing loader bucket design. The pneumatic actuator attached at the loader linkages is to make ease the loading and unloading process of the bucket. Themainmaterial used for the loader bucket is using Aluminiumwhile the wheelbarrow body is maintains to use the existing material, which is steel" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002772_978-3-030-67750-3_14-Figure11.6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002772_978-3-030-67750-3_14-Figure11.6-1.png", + "caption": "Fig. 11.6 Proposed design of material trolley", + "texts": [ + " So, the new material\u2019s trolley has been designed and fabricated to ensure comfort for workers. The current design of the material\u2019s trolley was designed based on four multi-direction wheels. That is the reason that the current design is heavy once it is loaded by the semi-finished good part. Furthermore, it will affect the body of workers and at the same time reduce the productivity of the production process. The newly designedmaterial\u2019s trolley applied the \u201cPipe& Joint\u201d concept to ensure more comfort for the transferring process as shown in Fig. 11.6. Furthermore, the cost of this newly designedmaterial\u2019s trolley is cheaper and it is easilymaintained because it just used a single Allen-key size for the maintenance process (shown in Fig. 11.7). Moreover, this new design of this material\u2019s trolley is lighter. This pipe and joint system is considered as a smart system, dynamic, and flexible modular assembly system consisting of plastic-coated steel pipes and metal joints. The flexibility of this system to tailor with several requirements such as racks, trolleys, workstations, gravity flow racks, and even light-duty machine structures was considered during the selection of this system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002399_s40430-021-02881-1-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002399_s40430-021-02881-1-Figure1-1.png", + "caption": "Fig. 1 Schematic diagram of the SbW system", + "texts": [ + " Both the adaptive state observer and event-triggered sliding mode control are presented in Sect.\u00a03. The simulations and experiments are provided in Sect.\u00a04. Section\u00a05 is the concluding remarks. Notations: \u211dn represents the n-dimensional Euclidean space and \u2124+ represents positive integers. | \u22c5 | represents the absolute value, and \u2016 \u22c5 \u2016 denotes the standard Euclidean vector norm. For the matrice A \u2208 \u211d n\u00d7n , max(A) and min(A) are the maximum and minimum eigenvalues of the matrix A, respectively. \u230a\u22c5\u2309 = \ufffd \u22c5 \ufffd sign(\u22c5), \u2200 \u2208 \u211d. The schematic diagram of SbW systems is shown in Fig.\u00a01, and the SbW system is mainly composed of the steering motor and its driver, reducer, gearbox, etc. The SbW system has two distinct characteristics: (1) the mechanical linkage between the steering wheel and front-wheels is no longer required; (2) an additional steering motor is applied to regulate the steering angle of front-wheels. Unlike conventional vehicles equipped with SbW systems, the desired steering angle of an autonomous vehicle\u2019s front-wheels is provided by its upper modules, including environment perception, decision-making, and path planning rather than the steering wheel" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002020_acemp.2007.4510512-Figure20-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002020_acemp.2007.4510512-Figure20-1.png", + "caption": "Fig. 20 Analytical model of CSPM motor.", + "texts": [ + " The comparison with the result by the conventional method (static E&S model) was examined in order to investigate features of the result by the dynamic analysis with dynamic E&S model, as shown in Fig. 19. The static E&S model has greatly evaluated the loss from actual result, since the effect of the distorted waveform is disregarded. It is found that magnetic field is suppressed by the effect of the eddy current. The static analysis excessively analyses the magnetic field. 4 Magnetic Characteristic Analysis of Permanent Magnet Motor Figure 20 shows the concentrated flux typed surface permanent magnet motor (CSPM motor), which was developed for high density machine by our group and have been obtained by Oita University as the patent. The gap field of this motor has the 1.5 times of he magnetization of permanent magnet. Figure 21 shows the magnetic flux distribution. From this result it was found that the magnetic path becomes short, therefore the inductance increases as this motor. Figs. 22 and 23 show the distribution of the magnetic field strength and the magnetic flux density, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000332_50003-4-Figure3.19-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000332_50003-4-Figure3.19-1.png", + "caption": "Fig. 3.19. Extended tyre brush model showing off set and deflection of carcass line (straight", + "texts": [ + " The force diagrams correspond reasonably well with the theoretical observations. The moment curves, however, deviate considerably from the theoretical predictions (compare Fig.3.18 with Fig.3.13). It appears that according to this figure M z changes its sign in the braking half of the diagram. This phenomenon can not be explained with the simple tyre brush model that has been employed thus far. The introduction of a laterally flexible carcass seems essential for properly modelling M z that acts on a driven or braked wheel. In Fig.3.19 a possible extension of the brush model is depicted. The carcass line is assumed to remain straight and parallel to the wheel plane in the contact region. A lateral and longitudinal compliance with respect to the wheel plane is introduced. In addition, a possible initial offset of the line of action of the longitudinal force with respect to the wheel centre plane is regarded. Such an off set is caused by asymmetry of the construction of the tyre or by the presence of a camber angle. With this model the moment M z is composed of the original contribution M z' established by the brush model and those due to the forces Fy and Fx which show lines of action shifted with respect to the contact centre C over the distances Uc and Vo + Vc respectively", + " The resulting calculations can be performed in a direct straightforward manner because the slip angle of the extended model is the same as the one for the internal brush model. Later on, in Section 3.3, the effect of the introduction of a torsional and bending stiffness of the carcass and belt will be discussed. The resulting model, however, is a lot more complex and closed form solutions are no longer possible. In Section 3.3 the technique of the tread element following method will be employed in the tread simulation model to determine the response. The combined slip response of the simple extended model of Fig.3.19 is given in the Figs.3.20 and 3.21. It is observed that in Fig.3.20 the aligning torque changes its sign in the braking range. This is due to the term in Eq.(3.51) with the compliance coefficient c. The resulting qualitative shape is quite similar to the experimentally found curves of Fig.3.18. In Fig.3.21 the effect of an initial off set of Vo- 5mm has been depicted. Here carcass compliance has been disregarded and only the last term of (3.51) has been added. We see that a moment is generated already at zero slip 116 THEORY OF STEADY-STATE SLIP FORCE AND MOMENT GENERATION angle", + "102) to avoid apparent changes in the effective rolling radius at camber. The actual lateral coordinate Yo (3.105) plus a term Yrr is used to calculate the aligning torque. With this additional term the lateral shift of Fx due to sideways rolling when the tyre is being cambered is accounted for. We have Yb, eff = Yb + Ye with yrr= eyrr bsiny with an upper limit of its magnitude equal to b. The moment M z causes the torsion of the contact patch and is assumed to act around a point closer to its centre like depicted in Fig.3.19. A reduction parameter e~ is used for this purpose. More refinements may be introduced. For details we refer to the complete listing of the program T r e a d S i m included in the Appendix. With the displacement vector As (3.101) established we can derive the change in deflection e over one time step. By keeping the directions of motion of the points B and P in Fig.3.34 constant during the time step, an approximate expression for the new deflection vector is obtained. After the base point B has moved according to the vector As we have: gi - - with g i = - 9 e i (3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002892_s11665-021-05931-w-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002892_s11665-021-05931-w-Figure1-1.png", + "caption": "Fig. 1 The AMB2018-01 sample geometry used in the study. (a) 3D model of the AMB2018-01 bridge structure benchmark geometry. (b) Three legs (lower left corner) were separated from the parts for microstructural analysis using wire electrical discharge machining (wire-EDM)", + "texts": [ + " Through a combination of 3D scanning, optical microscopy, finite element analysis (FEA), scanning electron microscopy (SEM) and electron backscatter diffraction (EBSD), dramatic microstructural variations within features of different thickness within the same component were observed. Overall, this work demonstrates that microstructural homogeneity within the as-printed stainless steel components that possess features of varying thickness can be achieved through local manipulation of LPBF process parameters. For this study, the AMB2018-01 bridge structure geometry developed by NIST (Fig. 1a) was used. The benchmark part is 75 mm long, 12 mm tall and 5 mm wide, with 12 \u2018\u2018legs\u2019\u2019 of three different thicknesses: four 5.0-mm legs, four 2.5-mm legs and four 0.5-mm legs. The samples were fabricated using a commercial SLM 125HL LPBF system (SLM Solutions, Lu\u0308beck, Germany). The chamber was kept in an inert nitrogen atmosphere, with oxygen level maintained at <100 ppm. Two samples were printed concurrently on the same annealed 316L stainless steel build plate. The build plate was preheated to and maintained at 200 C", + " The bridge structures were separated from the substrate using wire electrical discharge machining (wire-EDM). After separation from the substrate, the samples were imaged using an ATOS Core 200 3D scanner (GOM, Braunschweig, Germany). The 3D scans were processed with CloudCompare open source point cloud processing software, where the 3D scans were overlayed with the nominal CAD geometry, and cloud-to-mesh (C2M) distances were computed. Legs of 0.5, 2.5 and 5.0 mm in thickness were then separated from the bridge structure using wire-EDM (Fig. 1b). For examination under an optical microscope (OM) and scanning electron microscope (SEM), the legs were hotmounted in Konductomet conductive filled phenolic mounting compound (Buehler, Lake Bluff, IL, USA) with the y-z surface exposed. Metallographic grinding was performed with 120- 1200 grit silicon carbide papers using an AutoMet 250 grinder/ polisher (Buehler, Lake Bluff, IL, USA). The samples were then polished with 3 and 1 lm DiaLube diamond suspensions on White Label woven silk polishing cloth (Allied High Tech Journal of Materials Engineering and Performance Products, Inc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002804_j.matpr.2021.04.568-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002804_j.matpr.2021.04.568-Figure6-1.png", + "caption": "Fig. 6. Fatigue analysis of E-Glass/Jute/Epoxy.", + "texts": [], + "surrounding_texts": [ + "CAD models are designed in CATIA which contains special tools for the generation of traditional surfaces which will then be transformed into solid models, using traditional and composite monoleaf spring materials. The measurements of a spring leaf of a TATA SUMO vehicle are used for the design of the mono-leaf spring. Table1 displays the mono leaf spring configuration parameters and the planned mono leaf spring CAD model, as shown in Fig. 1. Finite Element Analysis (FEA) is a computational method for deconstruction into very small elements of a complex structure. In the simulated world, ANSYS offers an affordable way to examine the success of goods or processes. Digital prototyping is called this method of product growth. Users will iterate different scenarios using simulated prototyping techniques to refine the software even before the production is launched. This allows the probability and expense of failed designs to be reduced. In this study, a model was developed that was imported into the ANSYS workstation and FEA analysis. Meshing is the mechanism where the entity is divided into very small pieces called components. Elements. It is sometimes called a piece by piece. The leaf spring model is here meshed with a 10 mm brick mesh part scale. The front end is restricted and the rear end only in Y and Z is restricted; translational movement is permitted in X direction. Conditions of loading include applying a force on the middle of the spring of the leaf upward on the base of the leaf spring. The loading range is between 1000 N and 5000 N. The meshed model and limit and conditions of loading of a leaf spring are seen in Figs. 2 and 3." + ] + }, + { + "image_filename": "designv11_83_0003433_978-3-030-40667-7_6-Figure1.3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003433_978-3-030-40667-7_6-Figure1.3-1.png", + "caption": "Fig. 1.3 The \u201cUniversal Tyre-Testing Machine\u201d or \u201cUniversal Rig\u201d invented by V.E.\u00a0Gough in 1949 to test tyres at the Dunlop Rubber Company industries in England", + "texts": [ + " The two previous projects highlight two important aspects to be later find in hexapods: \u2013 The idea of a fixed base articulating with a distal mobile platform; \u2013 The use of hinged joints for articulating parts of parallel manipulators. In 1949, V.E.\u00a0 Gough came up with a new parallel manipulator: the variable length strut P. Sessa et al. octahedral hexapod [11]. This device greatly changed industry and led to numerous replicas. It was meant to test tyres at the Dunlop Rubber Company industries in England, where Gough worked as an engineer, and was called the \u201cUniversal Tyre-Testing Machine\u201d or \u201cUniversal Rig\u201d (Fig.\u00a01.3). This machine had to test tyres under different loads to simulate aero-landing loads. Surprisingly, the same Gough wrote in his papers that systems with six axis, i.e., hexapods, already existed, albeit the configuration of the six axis was different and made up of three horizontal and three vertical axis. These systems are known as MAST (Multi-Axis Simulator Table) and are still manufactured by many companies [10]. Gough needed to arrange the six axis in a octahedral hexapod configuration in order to test different loads on different space planes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003453_s10846-021-01454-7-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003453_s10846-021-01454-7-Figure6-1.png", + "caption": "Fig. 6 Swimming robot design. a SolidWorks\u00ae model, b Physical model", + "texts": [ + " (4) and (5), the ICC point can be found as [32]: ICC \u00bc x\u2212Rsin \u03b8s\u00f0 \u00de; y\u00fe Rcos \u03b8s\u00f0 \u00de\u00bd \u00f06\u00de At time t and \u03b4t the new position of the robot will be at: x\u0307 y\u0307\u03b8\u0307s h i \u00bc cos wt\u03b4t\u00f0 \u00de \u2212sin wt\u03b4t\u00f0 \u00de 0 sin wt\u03b4t\u00f0 \u00de cos wt\u03b4t\u00f0 \u00de 0 0 0 1 2 4 3 5 x\u2212ICCx y\u2212ICCy \u03b8s 2 4 3 5 \u00fe ICCx ICCy wt\u03b4t 2 4 3 5 \u00f07\u00de The diving module is consisting of one stepper motor and a movable/sliding mass that is able to slide on the lead screw of the motor. Motion commands are sent to the microcontroller unit via the HC-05 Bluetooth module. A set of four 1.5 V AA batteries were used to supply energy to both stepper (yaw) and servo (pitch) motors, communication unit and the microcontroller of the robot with the required energy as shown in Fig. 6. The robot\u2019s body is set to be first at neutrally buoyant state. Then after a specified period the pitch motor starts to rotate allowing the movable mass to be shifted toward the frontal part of the body. This will in turn pitches the robot down within an angle with the horizontal plane of the body. Since the yaw motor is triggered to operate synchronously, the robot will start to dive gradually. After a pre-determined period, the pitch motor starts to roll back, and consequently the movable mass will be shifted back towards the center of gravity (CG) point to allow the robot to floats gradually and then reaches to the neutrally buoyant state again" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002789_s00170-021-07207-y-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002789_s00170-021-07207-y-Figure1-1.png", + "caption": "Fig. 1 The structures of hydrostatic guideways: a typical hydrostatic guideway; b hydrostatic guideway with four oil storage holes", + "texts": [ + " The main heat source on hydrostatic guideway comes from the power loss, which results from the consumption of hydraulic power by pumping oil through bearing clearances. Thermal transmission from this internal heat source to critical components is inevitable. This paper optimizes the structure of hydrostatic guideway to reduce the thermal equilibrium time and acquire a uniform temperature distribution, in order to avoid the temperature variation and high thermal sensitivity of the structure to temperature variations under continuous usage of hydrostatic guideway. Figure 1a shows a typical hydrostatic guideway structure, wherein double-sided static pressure bearings are adopted, and multiple rails are wrapping the bearing parts to minimize the structural deformation. In this case, the table covers the rails, which are arranged on the guideway base, and bearings are located at lateral and top regions of the guideway base. During the machining processes, the temperature of the outer region of the base is higher than the inner region, and the temperature of top region of the base is higher than the bottom", + " A set of through holes in the bottom of guideway base intersects perpendicularly to four oil storage holes so that four oil storage holes interconnect with each other. Many vertical channels designed into the base make hydraulic oil heated by the power loss flow into four oil storage holes. The hydraulic oil converging in four oil storage holes is ultimately drained back to the oil tank through a common oil return pipe. When the heated oil flows throughmany channels designed into the base and converges in the four oil storage holes, the entire base is warmed up; then, a quick thermal equilibrium and a uniform temperature distribution are acquired. Figure 1b shows the hydrostatic guideway with four oil storage holes. In order to verify the effectiveness of the new base structure on equalizing the temperature distribution, a simplified model of the guideway base in the process of continuous uniform motion is established, as shown in Fig. 2. In this model, an ironcasting base is located on a natural granite bed. The geometry dimensions of the base are 900 mm in length, 140 mm in height, and 520 mm in width. The geometry dimensions of the bed are 1600 mm in length, 300 mm in height, and 1200 mm in width" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000468_20050703-6-cz-1902.01340-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000468_20050703-6-cz-1902.01340-Figure2-1.png", + "caption": "Fig. 2. Condition of mode switching", + "texts": [ + " However, it is necessary to actually consider the restraint condition concerning the stepping over excess of the obstacle, the center of gravity position of the robot and the operation speed of the robot. These restraint conditions are led in the next paragraph. foregoing paragraph, it explained the mode switch when the biped robot stepped over and exceeded the obstacle. Here, the constraint condition to change the mode is led actually. The appearance that changes from mode1 from mode2 and mode2 to mode3 is shown in Fig. 2. Tiptoe Q of the robot should exist at the right of upper right of the obstacleO in the state immediately before mode1 changes into mode2 from (a). Moreover, heel P of the robot should exist at the upper part of z\u03b5 (> 0) in the state immediately before changing from mode2 into mode3 from (b). These constraint conditions can be expressed as follows. xq \u2265 xo (1) zp \u2265 z\u03b5 . (2) The center of gravity position control because of no fall becomes important before the biped robot steps over and exceeds the obstacle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000360_4-431-27901-6_2-Figure2.6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000360_4-431-27901-6_2-Figure2.6-1.png", + "caption": "Fig. 2.6. An inclined plate in motion", + "texts": [ + " Continuity Equation The continuity equation for a small volume element in an incompressible fluid (assumption 3) can be written as follows: \u2202u \u2202x + \u2202 \u2202y + \u2202w \u2202z = 0 (2.12) Integrating this equation along the pillar-shaped element (dx \u00d7 dz \u00d7 h) in Fig. 2.5 in the film thickness direction from y = 0 to y = h gives: \u222b h 0 \u2202u \u2202x dy + \u222b h 0 \u2202w \u2202z dy + [ ]h 0 = 0 (2.13) This equation can be rewritten as follows by changing the order of integration and differentiation: \u2202 \u2202x \u222b h 0 u dy + \u2202 \u2202z \u222b h 0 w dy \u2212 ( u ) y=h \u2202h \u2202x \u2212 ( w ) y=h \u2202h \u2202z + [ ]h 0 = 0 (2.14) This is the continuity equation for a pillar-shaped element between the solid surfaces shown in Fig. 2.6. In deriving Eq. 2.14, the following mathematical formula must be used because the integration limit h in Eq. 2.13 is a function of x and z (only a formula concerning x is given here). \u2202 \u2202x \u222b b(x) a(x) f (x, y, z)dy = \u222b b(x) a(x) \u2202 \u2202x f (x, y, z)dy + f (x, b(x), z) \u2202b(x) \u2202x \u2212 f (x, a(x), z) \u2202a(x) \u2202x Equation 2.14 can be written as follows in terms of the surface velocities from the boundary conditions (assumption 7): \u2202 \u2202x \u222b h 0 u dy + \u2202 \u2202z \u222b h 0 w dy \u2212 U2 \u2202h \u2202x \u2212W2 \u2202h \u2202z + (V2 \u2212 V1) = 0 (2.15) e" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000961_s11668-007-9014-8-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000961_s11668-007-9014-8-Figure6-1.png", + "caption": "Fig. 6 Plastic deformation morphology in engaged regions", + "texts": [ + " 1, 4) and two adjacent teeth are plastically deformed (regions marked E and F in Fig. 1, 5). It is worth noting that some metal fragments were imbedded in the grooves between teeth in two regions of plastic deformation on the crankshaft gear and in one region on the camshaft gear. The length of the fragments corresponds to the face width. When the two gears were engaged according to their turning directions, the orientation of plastic deformation bands on the two failed gears was in correspondence with each other as shown in Fig. 6. Such an appearance may result from fractured teeth crushing onto the tooth faces and being embedded into the groove between teeth. Additionally, adjacent teeth fracture regions in the two gears are also in correspondence with each other and partly overlap. Visual inspection indicates that gear teeth fracture in the four regions of the camshaft gear and in the two regions of the crankshaft gear took place at the root area. Table 1 Main fabrication steps and related technical demands of gear Gear Process Specification Camshaft gear Quenching fi Tempering fi Nitriding Surface hardness: HV1 \u2021 460; Compound layer thickness: 10\u201325 lm Crankshaft gear Quenching fi Tempering fi Nitriding Surface hardness: HV1 \u2021 460; Compound layer thickness: 10\u201330 lm Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000239_iecon.2006.347435-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000239_iecon.2006.347435-Figure4-1.png", + "caption": "Fig. 4. Position relation between observation device and mark", + "texts": [ + " This paper assumes that (i) there are four cross candidate points on the search line, (ii) there are two pairs if the angle which the cross candidate point and the search center make, and (iii)there are four cross candidate points that the black pixel continues to the search center. It assumes that the region which fulfills these assumptions is the cross and if the cross is detected, we get center of the search region (uk, vk) as the mark position on the image. We explain the position measurement of the mark. Fig. 4 shows position relation between the observation device and the mark. First, we set up a coordinate system. We suppose that the XR, YR, ZR axis in Fig. 4 are the center of the observation device. This coordinate system is defined as the coordinate of the device and the position measurement of this paper is done based on this coordinate. This coordinate system is set up corresponds with the initial posture of the observation device. And the CCD camera and the laser range finder are installed as shown in Fig. 4. The ray axis of the CCD camera and the laser ranger are parallel to the ZR axis of the coordinate of the device. And the CCD camera and the laser range finder change the posture depending on the rotation of the coordinate of the device. We described the position measurement of the mark. This paper assumes that the mark has turned to the front to the observation device. The information given from the observation device is the rotation angle of each axis of the device \u03d5enc, \u03b8enc from encoders, the distance from the device to the mark l from the laser range finder. And the CCD camera gives the position of the mark (uk, vk) on the image which was mentioned above. First, a position of the camera center which is the position of the point C in Fig. 4 is calculated in consideration of the relation between the camera and the laser. Equation (4) express the distance between OR and C, and Equation (5) express posture angle \u03d5cam. l\u2032 = \u221a(a 2 )2 + (l + b)2 (4) \u03d5cam = \u03d5enc \u2212 arctan (a l ) (5) And the position of the camera center C is given by the following equation using the information of the distance and the angle given by Equation (4) and (5). \u23a1 \u23a3 xc yc zc \u23a4 \u23a6 = T\u03d5T\u03b8 \u23a1 \u23a3 0 0 l\u2032 \u23a4 \u23a6 (6) Where T\u03d5 and T\u03b8 are rotation matrix about \u03d5cam and \u03b8enc. Then mark position M is given by the following equations", + " Therefore, using the mark position on the image uk and vk, we calculate mismatched angles \u03d5img and \u03b8img. \u03d5img = arctan (s \u00b7 uk l ) \u03b8img = arctan (s \u00b7 vk l ) . (11) And considering the difference of the ray axis of the laser and the camera, following equations show error values \u03d5e and \u03b8e of mismatched angles \u03d5img and \u03b8img and desired values \u03d5d and \u03b8d and express as \u03d5e = \u03d5img \u2212 \u03d5d \u03b8e = \u03b8img \u2212 \u03b8d (12) \u03d5d = arctan ( 2a l ) , \u03b8d = 0 where a is the installation gap of each components which is shown in Fig. 4. We construct the control system in order to revise the mismatched. This control system applies the general PI control system, and uses the erroneous value which is calculated by equation (12) and integration value of error as feedback values. We calculate control angles \u03d5cmd and \u03b8cmd using following equations. \u03d5cmd = kP \u03d5e + kI \u222b \u03d5edt + \u03d5m \u03b8cmd = kP \u03b8e + kI \u222b \u03b8edt + \u03b8m (13) kP , kI are the feedback gain respectively. And \u03d5m and \u03b8m in equation (13) mean estimation of the variation of the mark based on extended Kalman filter and are calculated by following equations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001581_icma.2008.4798866-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001581_icma.2008.4798866-Figure3-1.png", + "caption": "Fig. 3. Detection of Occlusion Areas", + "texts": [ + " Occlusion areas are divided into two types - one caused by geometrical shape, and one caused by moving objects. Our method can handle both types of occlusion areas. The bad effects from moving oqjects are smaller than occlusion area, because the moving oQiect in rooms such as people and chairs are cOlll'aratively small in LRF scan. However, when a small oQiect passes near LRF, a lot of areas can be hided by it. The m;thod of detecting occlusion caused by environm;n tal shape described as follows. In the example of figure 3, The robot observed each point of A1, A2, and A3 at position R, but it is impossible for the robot locating position R\u2032 to observe them. If the robot is now locating R and the reference scan is obtained in R\u2032, these points become occlusion areas. It is assumed that accurate moved distance and the angle of traverse from R\u2032to R turn out. At this time, the change in the angle of A1, A2, and A3 in R polar coordinate system is clockwise in order. However, the angle of A1, A2, and A3 in R\u2032 polar coordinate system become anti-clockwise in order" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000061_iros.2004.1389871-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000061_iros.2004.1389871-Figure1-1.png", + "caption": "Figure 1. Schematic of the mobile manipulator", + "texts": [ + " A fuzzy membership function based neural network that incorporates a fuzzy-neural interpolating network 191 is proposed to solve this problem of the potential singularity of the feature Jacobian matrix. The proposed fuzzy membership function based neural networks are used to approximate a nonlinear mapping that transforms image features and their changes into the desired camera motion. The camera must he calibrated to reduce the complexity of the approximation to the mapping. This paper primarily adopts the uncalibrated eye-inhand vision system to provide visual information to support the control of the robot manipulator mounted on the mobile base, as shown in Fig. 1. The objective of this paper is to develop a strategy for controlling the manipulator to approach and pick up a workpiece. This vision-guided control strategy is based on selected image feat\" and an image-based look-and-move control structure. Using this strategy, the mobile manipulator is expected to travel between any two stations in the production Line without stopping in-between, and to perform pick-and-place operations on non-planar ground without any special lighting. n. CARTES~AN SPACE Let c~ he an m-dimensional relative position vector of a workpiece with respect to the camera frame C , and let F he an n-dimensional feature vector of the workpiece image" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000143_6.2006-6241-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000143_6.2006-6241-Figure3-1.png", + "caption": "Figure 3. The error \u2016e\u2217(tc)\u2016\u221e in computing the inverse input at the current time instant tc exponentially decays with preview time Tp = tf \u2212 ti. The lower bound on the decay rate \u03b3\u0302 depends on the minimum distance \u03b3 of the unstable poles of the linearized internal dynamics from the imaginary axis, and on the nonlinearity of the internal dynamics (quantified by the Lipschitz constant K1).", + "texts": [ + " (23) 4 of 17 American Institute of Aeronautics and Astronautics Proof The proof, given in the Appendix, follows the argument for general nonlinear systems as in Ref. 1,2. We note that the error in computing the inverse (Eq. (22)) has two terms: the first term reflects the effect of finite number of iterations m, and the second term reflects the effect of finite preview time Tp. Remark 2 Theorem 1 shows that the error in computing the inverse input exponentially decays as the preview time Tp = tf \u2212 tc increases, as depicted in Fig. 3 (where e\u2217ff(\u00b7) = limm\u2192\u221e eff,m(\u00b7)). The worstcase decay rate \u03b3\u0302 of the computational error depends on the locations of the unstable poles of the internal dynamics\u2013the smallest distance of the unstable poles (of the internal dynamics) from the imaginary axis (which equals to \u03b3 for the VTOL example). This relationship is depicted in Fig. 3. For linear time invariant (LTI) systems, such relationship reduces to our previous results24\u2014for LTI systems, the Lipschitz constant K1 = 0 and no iteration is needed to compute the preview-based inverse input. Therefore, the decay rate \u03b3\u0302 is arbitrarily close to the minimal distance of the unstable internal dynamics poles the from the imaginary axis (i.e., the smallest real part of the nonminimum-phase zeros). Remark 3 If the original nonlinear system is locally exponentially stable around the origin, then standard stability theory arguments (e" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003388_9781119725008.ch3-Figure3.2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003388_9781119725008.ch3-Figure3.2-1.png", + "caption": "Figure 3.2 Basic design of a biofuel cell.", + "texts": [ + " In this chapter, we discuss the biofuel cells, their types, and the benefits by judiciously utilizing the physiological fluids to generate electricity. Further, the factors that affect fuel production efficiency including the nature of materials used for the design and development of the electrodes for various biomedical applications as implantable energy sources is also highlighted. BFCs utilize a living organism or a physiological source of energy to generate electricity. The conventional biofuel cell is comprised of an anode and cathode, which are separated via a proton exchange membrane to avoid mixing (Figure 3.2). The two electrodes are immersed in an electrolyte and are connected with the help of an external wire. At the anode, fuel is oxidized to release an electron, which is then passed through a coil to reduce an oxidant at the cathode. The primary aim of the fuel cell is to supply power to low-power devices. Biofuel cells can be classified conventionally on the basis of the location of the enzymes. If the proteins are placed inside the living cells, then these biofuel cells are referred to as microbial fuel cells, whereas if the catalyst is located outside the living cells, then they\u2019re termed as enzymatic fuel cells" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000808_j.engfailanal.2008.11.002-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000808_j.engfailanal.2008.11.002-Figure2-1.png", + "caption": "Fig. 2. Double torsion geometry.", + "texts": [ + " Thermosetting polymers are usually brittle and show a linear elastic behaviour with very little plastic deformation therefore, they fulfil the basic requirements of linear elastic fracture mechanics (LEFM). Different testing methods are available to study the fracture behaviour of these materials. A common technique for fracture toughness determination uses the single edge notched in bending (SENB) geometry. This technique is standardized by ISO 13586 and ASTM D-5049 with the specimen dimensions as shown in Fig. 1. However, the crack propagation can be studied by the double torsion technique whose geometry is shown in Fig. 2. Although not standardized, a distinguishing feature of this loading configuration is that the stress intensity factor, KI, is independent on the crack length for a range of crack lengths in the test specimen [3]. Some studies on the fracture behaviour of thermosetting polymers using double torsion techniques indicate that they usually exhibit an unstable non-continuous type of fracture called stick-slip, as opposed to thermoplastic polymers that show continuous and stable crack propagation. These two types of behaviour can be easily distinguished in the plot of the load versus displacement, as shown in Figs", + " Superfin corresponds to the commercial name of a copolymer obtained from a monomer mixture of 50% ADC and 50% of an aromatic polyester oligomer of high molecular weight, terminated by allyl groups that copolymerized with ADC through a radical reaction. The double torsion test configuration consists of symmetric four-point loading around a crack or a notch on one end of a rectangular plate. This produces torsional deformation in the two plate halves. Rectangular plates of 63 33 6 mm were mechanized from the polymer preforms. Test specimen, as shown in Fig. 2, included a 2 mm longitudinal groove along the length of the plate, in order to guide the crack growth. Notch depth was 8 mm and both the roots of the notch and the groove were sharpened using a razor blade. Wn, the moment arm (the separation between the loading and support points of each arm), was 12.5 mm. All mechanical test were performed at room temperature in a Galdabini universal testing machine using a 1 kN load cell, unless noted. In double torsion test, plain strain KIC can be calculated [4] using KIC \u00bc PcWm 1 d3dnW\u00f01 m\u00deW " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001605_paciia.2008.230-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001605_paciia.2008.230-Figure1-1.png", + "caption": "Figure 1 Robotic system with flexible joint The corresponding state-space model is", + "texts": [ + " This problem is designed to converge to a local minimum of the condition number. For the problem of observer design for the class of nonlinear systems considered in this paper, algorithm in [9] has been modified. We firstly make the whole left-half complex plane as the allowable set, then the proved algorithm coded in matlab continuously changes the locations of the eigenvalues so as to keep reducing the condition number. Finally we get the gain matrix L satisfying the theorem 3. Consider a one-link manipulator with revolute joints actuted by a DC motor, as shown in Fig 1. The elasticity of the joint can be well-modeled by a linear torsional spring[11]. The elastic coupling of the motor shaft to the link introduces an additional degree of freedom. The states of this system are motor position and velocity, and the link position and velocity[1,11]. ( ) ( ) ( ) 1 1 1 1 1 1 1 1 sin m m m m m m m m m Kk B u J J J k mgh J J \u03c4 \u03b8 \u03c9 \u03c9 \u03b8 \u03b8 \u03c9 \u03b8 \u03c9 \u03c9 \u03b8 \u03b8 \u03b8 = = \u2212 \u2212 + = = \u2212 \u2212 \u2212 with mJ being the inertia of the motor; 1J the inertia of the link; m\u03b8 the angular rotation of the motor; 1\u03b8 the angular position of the link; m\u03c9 the angular velocity of the motor; and 1\u03c9 the angular velocity of the link" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001352_1077546307076901-Figure15-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001352_1077546307076901-Figure15-1.png", + "caption": "Figure 15. Definition of internal nodes on a flexible slider crank.", + "texts": [ + " Table 3 lists the CPU time and the number of degrees of freedom for cases 1 to 4, and the number of element from one to four. The results are based on the following information: Time step of integration: 4 10 4 s Time integration duration: 20 cycles at rotating speed 150 rad/s Hardware: Pentium 4 CPU 3.00GHz personal computer Software: MATLAB Version 5.3 (R11.1) As shown, there is a quasi-linear relationship between the CPU time and the number of degrees of freedom. To apply the r-refinement to a flexible slider crank mechanism, internal nodes are defined in Figure 15. Symbols xi represents the non-dimensional position of an internal node, ranging from 0 to 1. One performs an optimization problem by minimizing a specific error as: Min e xi where xi should satisfy 0 xi 1. The error function e xi is defined by equation (27). The results shown in this section were obtained using a MATLAB Toolbox command \u201cfminsearch\u201d, which applies the Nelder\u2013Mead nonlinear minimization. The objective of this optimization is to find the positions of internal nodes that perform a minimum error" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001815_isitae.2007.4409284-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001815_isitae.2007.4409284-Figure1-1.png", + "caption": "Figure 1. A typical integrated Webpage of the teaching material", + "texts": [ + " A mass of 2D and 3D animations imitate the actions of the teachers in the classroom vividly, lively and accurately. This allows the students to grasp the knowledge by audio-visual sense synchronously and understand the knowledge easily in a relaxed studying environment. It is gratified that great enthusiasm of the students for studying the curriculum has been built unexpectedly. In fact, the students are all attention to the lessons in classroom. The following pictures show the characters of the web based multimedia teaching material. The Figure 1 is a typical webpage example. Such compact, amicable and man-machine interactive architecture organized according to knowledge points is propitious to understand the principles and their applications quickly. As a result, the period the whole lessons needed is shortened evidently and the abundant time can be spent to exercise varity practice activities, just like to look around factory, to disassemble and assemble some thingamys accordingly. It is in favor of enhancing the ability to use the techniques, skills, and modern engineering tools necessary for an engineer" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001890_2007-26-073-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001890_2007-26-073-Figure2-1.png", + "caption": "Figure 2 : Locations and Directions of Accelerometers at Middle Passenger Contact Points", + "texts": [ + " These PDDs are compared for various vehicles to evaluate the ride comfort instead of using single digit given by the above explained ISO methods. This gives additionally the information about the amount of time spent by human in a comfort zone over a period of exposure time. Ten different makes of three wheelers with 3 and 6 seaters were identified and chosen for this exercise. Table- III shows the type of vehicles used for the exercise. Tri-axial accelerometers were mounted on the middle passenger seat, feet and one uni-axial accelerometer in longitudinal direction at back rest. Fig. 2 shows accelerometer mounting locations. The instrumented three-wheelers were driven on two types of roads which would represent typical \u2018smooth city road\u2019 (road A) and \u2018rough city road\u2019 (road B) conditions at a speed of 35-45 km/h and 30-35 km/h respectively. Vehicle were loaded to their GVW condition Initial statistical checks were carried out by analysing the time series data for its maximum, minimum and RMS. These readings were checked and for identified channels data were filtered through weighting curve as per ISO 2631" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002657_tmag.2021.3073155-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002657_tmag.2021.3073155-Figure1-1.png", + "caption": "Fig. 1. Topologies of the 6/17 C-MFSPM machine.", + "texts": [ + " In Section II, the topology of FP-MFSPM machines is introduced, and six FP-MFSPM machines with different rotor teeth are compared by finite element analysis (FEA) to find the optimal statorcoil/rotor-tooth combinations. Afterward, the performance comparisons between C-MFSPM and FP-MFSPM machines with optimal stator-coil/rotor-tooth combinations are compared in detail, including no-load back electro-motive force (backEMF), radial air-gap flux density, cogging torque, and electromagnetic torque. Finally, several conclusions are drawn in Section IV. The 6/17 C-MFSPM machine is presented in Fig. 1, where phase winding connection is that the two coils spatial radially distributed are serially connected into one phase winding, such as A1 and A2 for phase winding A. It should be noted that the armature windings adopt a concentrated structure, and coil conductors layout of phase A is A1+, A1-, A2-, A2+. By changing the winding configuration from concentrated winding to full-pitch winding, a 6/17 FP-MFSPM machine is formed, as is shown in Fig. 2, where coil conductors layout of phase A turns into A1'+, A2'+, A1'-, A2'-" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003365_s12666-021-02364-w-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003365_s12666-021-02364-w-Figure3-1.png", + "caption": "Fig. 3 Model mesh and boundary conditions for the melt pool simulation", + "texts": [ + " The notion of segregating the domain is to discretize the space with finer elements in the vicinity of the laser flux area and coarser in the rest of the volume. This approach thus saves computational time and cost by reducing the overall number of elements. The top of Domain 1 is subjected to laser irradiation. The high density laser beam of Gaussian distribution (as shown in Fig. 2) is focused at the focal position over the substrate top. The entire computational domain is discretized using a nonuniform unstructured grid of tetrahedral cells. The mesh consists of 274,628 elements with an average quality of 0.8063 (shown in Fig. 3). The SS316L material substrate is used in the present case. The properties of the material are shown in Table 1. The thermal conductivity and specific heat are considered to be temperature-dependent properties of the material and are defined using an analytic function (refer to Table 1). The following governing equations are used for the simulation of heat model [31]. The mass, momentum and energy conservation are as follows: oq ot \u00fer: qu\u00f0 \u00de \u00bc 0 \u00f01\u00de q ou ot \u00fe q u:r\u00f0 \u00deu \u00bc r: pI \u00fe l ru\u00fe ru\u00f0 \u00deT 2l 3 r:u\u00f0 \u00deI \u00fe F \u00f02\u00de qC p oT ot \u00fe qC pu:rT \u00bc r: krT\u00f0 \u00de \u00fe Q \u00f03\u00de Here q, u, and p are the density, velocity, and fluid pressure, respectively, and l is the fluid dynamic viscosity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002945_978-3-030-69178-3_15-Figure1.3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002945_978-3-030-69178-3_15-Figure1.3-1.png", + "caption": "Fig. 1.3 Digital model (left) and kinematic graph (right) of a robotic workcell [48]", + "texts": [ + " The DT should combine and align all relevant aspects of modelling the function, structure and behaviour of the robotic cell including the workers, together with capturing the symbiotic interplay of the human and robotic agents. This includes representing the multimodal and bidirectional channels of communication and control as well. In the HRC context, the basis of the digital twin can be a linkage mechanism, which is capable of representing the geometric and kinematic relations of all static and moving objects in the workcell [48] (see Fig. 1.3). As planning and execution proceed, the initial model is gradually enriched with new elements obtained with perception or deduction. Time and again, the DT should be efficiently calibrated to the real environment by using the actual measured data, so that both robot programming and operator instructions can automatically adapt to the actual situation (see also Chap. 10 [45]). The application of digital twins and shadows, together with predictive engineering, can lead to anticipatory rather than reactive planning and control in a HRC environment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002333_s00500-021-05640-5-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002333_s00500-021-05640-5-Figure7-1.png", + "caption": "Fig. 7 Stress distribution in the coil spring", + "texts": [ + " The low-frequency-energy-based durability prediction models were then established from the power relationship exhibited by the low-frequency energy and fatigue lives as shown below: where the A and B are used to denote the constant values. By correlating the predicted fatigue lives from the lowfrequency-energy-based models with the fatigue lives of the experimental strain data, the survival rate of the fatigue life correlation within the conservative fatigue life boundary of 200% was thus evaluated as a way of validating the accuracy levels of the fatigue life prediction models (Kong et al. 2019a). The coil spring was subjected to a static load of 3600 N as shown in Fig. 7 with a higher stress level on the inner surface of the coil spring (facing the spring axis) implied the susceptibility of the inner surface to fatigue cracks (Zhu et al. 2014). This was caused by the transverse and torsional shear stresses associated with the deflecting of coil spring (Pattar et al. 2014). Apart from the simultaneous occurrence of stresses, spring curvature effect was also found to have a significant influence on coil springs with a small spring index (ratio of mean spring diameter to coil diameter) (Dragoni and Bagaria 2011), where in this case had been the suspension coil spring" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003073_978-3-030-60990-0_9-Figure22.2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003073_978-3-030-60990-0_9-Figure22.2-1.png", + "caption": "Fig. 22.2 Vehicle model illustration. \u00a9 2019 IEEE. Reprinted, with permission, from [19]", + "texts": [ + " Specifically, the human-in-the-loop shared control system will be formulated as a set of differential-difference equations, where the delay in the state variable is produced by a driver\u2019s reaction time. The complete vehicle model consists of the lateral vehicle dynamics, the steering column and the vision-position model for the lane-keeping task. According to [34, 38], under the assumptions of small angles and constant longitudinal speed, it can be described by x\u0307v(t) = Avxv(t) + Bv ( u(t) + Td(t) ) + Dv\u03c1, (22.1a) y(t) = Cvxv(t), (22.1b) where xv = [ vy ra \u03c8L yL \u03b4 \u03b4\u0307 ]T , y = yc is the lateral offset at the vehicle center of gravity, and \u03c8L = \u03c8v \u2212 \u03c8r as illustrated in Fig. 22.2. In (22.1), the model assumes that we have a constant curvature \u03c1. In practice, the non-constant road curvature can be approximated by piecewise constant functions as in [19, 31]. In a detailed description, vy is the vehicle lateral velocity, ra is the yaw rate, \u03c8L is the heading angle error, yL is the previewed distance from the lane center at a look-ahead distance ls , \u03b4 is the steering angle, \u03b4\u0307 is the steering angle rate, and Td is the driver\u2019s torque. See Fig. 22.2 for details. The state vector xv can be measured using various sensors including a camera and the inertial measurement unit (IMU). The matrices Av , Bv , Cv and Dv are expressed as Av = \u23a1 \u23a2 \u23a2\u23a2\u23a2\u23a2\u23a2 \u23a3 a11 a12 0 0 b1 0 a21 a22 0 0 b2 0 0 1 0 0 0 0 1 ls vx 0 0 0 0 0 0 0 0 1 Ts1 Ts2 0 0 Ts3 Ts4 \u23a4 \u23a5 \u23a5\u23a5\u23a5\u23a5\u23a5 \u23a6 , Bv = \u23a1 \u23a2 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2 \u23a3 0 0 0 0 0 1 Is Rs \u23a4 \u23a5 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a5 \u23a6 , Cv = [ 0 0 \u2212ls 1 0 0 ] , Dv = [ 0 0 \u2212vx 0 0 0 ]T , where a11 = \u22122(C f + Cr ) Mvx , a12 = 2(Crlr \u2212 C f l f ) Mvx \u2212 vx , a21 = 2(Crlr \u2212 C f l f ) Izvx , a22 = \u22122(C f l2f + Crl2r ) Izvx , b1 = 2C f M , b2 = 2C f l f Iz , Ts1 = 2C f \u03b7t Is R2 s vx , Ts2 = 2C f l f \u03b7t Is R2 s vx , Ts3 = \u22122C f \u03b7t Is R2 s , Ts4 = \u2212 Bs Is , and l f is the distance from the center of gravity to front axle, lr is the distance from the center of gravity to rear axle, M is the mass of the vehicle, \u03b7t is the tire length contact, Iz is the vehicle yaw moment of inertia, Is is the steering system moment of inertia, Rs is the steering gear ratio, Bs is the steering system damping, C f is the front cornering stiffness, Cr is the rear cornering stiffness, and vx is the fixed longitudinal velocity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002925_01423312211016188-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002925_01423312211016188-Figure1-1.png", + "caption": "Figure 1. Rolling bearing structure and simplified dynamic model.", + "texts": [ + " Rolling bearing local fault vibration model In order to study the local fault influence on dynamic characteristics of rolling bearings, it is necessary to build a rolling bearing dynamic model. Considering the accuracy and calculation efficiency, the following assumptions are made: (1) both inner and outer ring of bearing are rigid and are installed on rotor and bearing seat respectively with interference fit; (2) roller sliding, mass, inertia force and cage are ignored, and the contact with inner and outer rings meets Hertz contact theory; (3) axial bearing structure and loads are ignored. As a result, rolling bearing structure can be shown as Figure 1(a). Rollers are simplified as springs and damping, and the simplified bearing dynamic model can be established as Figure 1(b). Under the action of radial load, the inner and outer ring axes have relative displacement. Some rollers are compressed and form support force. Therefore, the radial deformation of ith roller can be expressed as follows: di = xin xout\u00f0 \u00decosui + yin yout\u00f0 \u00desinui 0:5c df \u00f01\u00de ui =vct + 2p i 1\u00f0 \u00de nb + u0 \u00f02\u00de vc = 1 2 vin 1 dcosa D +vout 1+ dcosa D \u00f03\u00de where xin and yin are inner ring vibration displacements,xout and yout are outer ring vibration displacements, c is bearing radial clearance, ui is the phase of ith roller, nb is roller number, u0 is the initial phase of first roller, vc is cage speed, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002360_012005-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002360_012005-Figure2-1.png", + "caption": "Figure 2. Fishtail open turn with rectilinear reverse movement, which is not parallel to the field boundary - movement from right to left.", + "texts": [ + " The purpose of the present work is to derive analytical dependences for determining the length of fishtail turns with a rectilinear reverse move that is not parallel to the field boundary, as well as for determining the width of the headland required to perform them in an irregularly shaped field and to analyze the influence of the angle between the direction of movement and the field boundary on the length of the non-working move and the width of the headland at different direction of movement of the unit in the field. A fishtail open turn and a fishtail closed turn with a rectilinear reverse movement in different directions of their performance in a field with an irregular shape are considered. To determine the length of the turn, a geometric method is used, in which the turn is represented by straight lines and arcs of a circle of equal radius. The different types of turns are presented in Figure 1, Figure 2, Figure 3 and Figure 4. The length of the turn is defined as the sum of the lengths of its geometric elements. The width of the headland required to perform the turn is defined as the sum of the segments perpendicular to the field boundary and depending on the elements of the turn. The symbols used in the figures are as follows: \u03b1 is the angle between the direction of working move of the unit and the boundary of the field; p. A \u2013 the beginning of the turn; p. B \u2013 the end of turn; p. O1, p. O2 \u2013 the centers of the first and second curvilinear movement within the turn; \u03b21, \u03b22 \u2013 the central angles of the arcs described in the first and second curvilinear movements; R \u2013 the radius of curvilinear movement (turning radius of the machine-tractor unit); M \u2013 the width of the tractor measured from the outside of the wheels", + " sin \ud835\udefc (12) The transition between the modes of movement shown in Figure 1(a) and Figure 1(b) is performed at an angle \ud835\udefc = tan\u22121 ( 2\ud835\udc45+\ud835\udc3b.sin \ud835\udefd1 \ud835\udc3b .cos \ud835\udefd1+2\ud835\udc59\ud835\udc4e ), (13) in which the width of the headland in both variants is the same. In this way of movement after reaching the boundary of the headland, a curvilinear movement is performed until reaching the boundary of the field, after which a rectilinear reverse movement is performed and again a curvilinear movement forward to the boundary of the headland (Figure 2). At the end of the reverse movement, the tractor wheels may be outside the headland with width E, i.e. in the field. The length of the turn is \ud835\udc59\ud835\udc47 = \ud835\udf0b\ud835\udc45 + \ud835\udc421\ud835\udc422 \u0305\u0305 \u0305\u0305 \u0305\u0305 \u0305 = \ud835\udf0b\ud835\udc45 + 2\ud835\udc45+\ud835\udc35 sin \ud835\udefd2 (14) The central angle \u03b22 is determined by the dependence \ud835\udefd2 = tan\u22121 ( 2\ud835\udc45+\ud835\udc35 \ud835\udc35 tan \ud835\udefc \u22122\ud835\udc59\ud835\udc4e ) (15) After substitution in equation (14) for the length of turn is obtained \ud835\udc59\ud835\udc47 = \ud835\udf0b\ud835\udc45 + [(2\ud835\udc45 + \ud835\udc35)2 + ( \ud835\udc35 tan \ud835\udefc \u2212 2\ud835\udc59\ud835\udc4e) 2 ] 1 2\u2044 (16) ICTTE 2020 IOP Conf. Series: Materials Science and Engineering 1031 (2021) 012005 IOP Publishing doi:10", + " Series: Materials Science and Engineering 1031 (2021) 012005 IOP Publishing doi:10.1088/1757-899X/1031/1/012005 The length of the turn is \ud835\udc59\ud835\udc47 = \ud835\udf0b\ud835\udc45 + \ud835\udc421\ud835\udc422 \u0305\u0305 \u0305\u0305 \u0305\u0305 \u0305 = \ud835\udf0b\ud835\udc45 + 2\ud835\udc45\u2212\ud835\udc35 sin \ud835\udefd2 = \ud835\udf0b\ud835\udc45 + [(2\ud835\udc45 \u2212 \ud835\udc35)2 + ( \ud835\udc35 tan \ud835\udefc \u2212 2\ud835\udc59\ud835\udc4e) 2 ] 1 2\u2044 , (45) because the central angle \u03b22 determined by the triangle O1CO2 is \ud835\udefd2 = tan\u22121 ( 2\ud835\udc45\u2212\ud835\udc35 \ud835\udc35 tan \ud835\udefc \u22122\ud835\udc59\ud835\udc4e ) (46) The length of the turn has a minimum at sin \ud835\udefd2 = 1, i.e. at \ud835\udefd2 = 90\u00b0. The turn has a minimum length at \ud835\udefc = tan\u22121 ( \ud835\udc35 2\ud835\udc59\ud835\udc4e ) , (47) as well as in the open turn performed in the same direction (Figure 2). For \ud835\udefc > tan\u22121 ( \ud835\udc35 2\ud835\udc59\ud835\udc4e ) the angle \u03b22 is determined by the dependence \ud835\udefd2 = 180 + tan\u22121 ( 2\ud835\udc45\u2212\ud835\udc35 \ud835\udc35 tan \ud835\udefc \u22122\ud835\udc59\ud835\udc4e ) (48) The width E of the headland at a small angle \u03b1 (Figure 4, (a)) is determined by the dependence ICTTE 2020 IOP Conf. Series: Materials Science and Engineering 1031 (2021) 012005 IOP Publishing doi:10.1088/1757-899X/1031/1/012005 \ud835\udc38 = 0,5(\ud835\udc40 + \ud835\udc35) cos \ud835\udefc + (\ud835\udc3b + \ud835\udc59\ud835\udc4e). sin \ud835\udefc (49) When performing the fishtail turn, the tractor goes deeper into the field. The total width of the headland, taking into account the innermost point in the field reached by the tractor, can be defined as the sum of the lengths of the following sections: \ud835\udc4e = 0,5\ud835\udc40" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000130_2005-01-2911-Figure21-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000130_2005-01-2911-Figure21-1.png", + "caption": "Figure 21: Pressure vessel shell structure unfolded", + "texts": [ + " A potential configuration for modules in a mobility system (Figure 19) has been explored using Klikko geometrical construction set (Klikko n.d.), which allows for assembly of shell structures using panels similar to the Trigon system. We have shown that the kinematic functionality of Trigon edge panel manipulator mechanisms will allow modules to climb existing portions of the structure and travel to their own target destinations (Figure 20). The edge connectors allow two adjacent panels to lie flat against each other so that volumes can be collapsed for compact transport (Figure 21). We are also working on a concept where thin-walled pressure barriers can be folded into \u201cseed cores\u201d, with hatch hardware already attached. Trigons panels carrying radiation shielding payloads can then interface with the cores, and fill out remaining structure for habitable pressure vessels (Figure 23). Square Trigon panels can be formed into cubes, with specialty payload panels dedicated to assembly and manufacturing functions. These cubes can function as cassette factories that are self-assembled into factory lines for production, either to self-replicate, or create other useful products (Figure 24)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000828_3.43615-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000828_3.43615-Figure8-1.png", + "caption": "Fig. 8 Control force and angle-of-attack variation with control force location.", + "texts": [ + " As the control surface approaches the neutral point, the force would have to increase. Since a constant control surface moment is required, the control force for trim varies inversely as the distance between the neutral point and the control surface force lc'. When the control surface is forward of the neutral point, the direction of the control surface force has to be reversed. The control force plotted as a function of the control surface location along the body as measured from the neutral point is shown as the solid line in Fig. 8. For the constant weight, e.g., altitude, and velocity, the sum of the lift due to the control surface and to the rest of the configuration must be a constant equal to the weight. This is shown in Fig. 8 in terms of the lift coefficients of Fig. 7. The required lift coefficient of the configuration less the control surface is shown as the dashed line. Since CL^C \u2014 (CLa) -coi, the variation of a is also represented by the dashed line provided the vertical scale is changed accordingly. The linearizing assumption of small angles is violated when the control surface is near the neutral point. The actual force (NP).C Fig. 7 l /2 />V 0 S lnm Control surface location: aerodynamic relationships at the neutral point", + " Once the angle of attack a has been determined, the control surface deflection can be found for a given surface whose relationship CLC = f(ot, 5) with positions along the body is known. If the neutral point of a configuration is located well forward, such as a slender body with small stabilizing surfaces, control surfaces located well forward could mean that the distance between the neutral point and the control surfaces would be rather small. The qualitative force system (or a similar one based on the unbalance moment opposite to the direction assumed for Fig. 8 or on the e.g. aft of the neutral point) of Fig. 8 shows that the forces required for trim, and consequently the trim angle of attack and control surface deflection, would be large. This is probably why configurations with forward control surfaces have very large stabilizing surfaces. The neutral points are quite far aft of the control surface location. The aerodynamic forces and moments during a maneuver are dependent on the pitching (or yawing) velocities, as well as the angle of attack and the control surface deflection. An indication of the maneuverability of the configuration is the elevator angle per gravity acceleration normal to the flight path required to maintain the vehicle in a steady turn" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000184_6.2004-4872-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000184_6.2004-4872-Figure1-1.png", + "caption": "Figure 1 The ICE vehicle", + "texts": [ + " Of particular importance to the design will be the ability of the aircraft to sustain significant system failures/damage, here to be represented by variations in control surface actuator characteristics and vehicle dynamics. As will be seen, the particular design approach to be espoused will not require failure identification/isolation nor control law restructuring. As such, it can be considered as an alternative to reconfigurable design approaches that do typically require such identification and restructuring. The vehicle that is the subject of the challenge is the Innovative Control Effector (ICE) vehicle shown in Fig. 1. A detailed description of a linear model of the vehicle is given in the Ref. 1 The control system response types for the design challenge are as follows: For longitudinal control, a pitch-rate command system is desired with no requirement for attitude hold. For lateral/directional control, roll-rate and sideslipcommand systems are desired, again with no requirement for attitude hold in the roll-rate system. Commands to the pitch-rate and roll-rate systems are assumed to be generated by cockpit inceptor commands, e" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002560_tec.2021.3069096-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002560_tec.2021.3069096-Figure4-1.png", + "caption": "Fig. 4 Flux path of 18/30 combination. (a) \ud835\udc3c\ud835\udc5e excitation (b) \ud835\udc3c\ud835\udc51 excitation.", + "texts": [ + " Three different slot/pole combinations for the rotor-end saliency are compared: 1) the machine M1 with integral slot distributed winding (18 slots/6 pole PM rotor, 30 pole rotor end); 2) the machine M2 with fractional slot distributed winding (18 slots/8 pole PM rotor, 16 pole rotor end) and 3) the prototype machine (12 slots/14 pole PM rotor, 10 pole rotor end). For fair comparison, the rotor geometry is set to the same size for all the combinations such that all air-gaps are 0.5mm and the rotor tooth width is half of the pole pitch. 1) M1 Machine The field flux maps under \ud835\udc3c\ud835\udc5e and \ud835\udc3c\ud835\udc51 stator current excitation are illustrated in Fig. 4 (a) and (b) respectively. It is seen that the \u201cleakage\u201d of the stator excited flux becomes much more serious when the carrier current vector is oriented on the d-axis of the reference frame defined by the rotor end. In this scenario, maximum overall reluctance takes place accounting for minimum inductance \ud835\udc3f\ud835\udc5a\ud835\udc56\ud835\udc5b. In Fig. 4(b), the flux path is expected to cross the rotor slot. However, due to the large rotor pole number and the narrow pole tooth width, the stator flux leaks through the adjacent teeth. The overall flux linkage under d-axis current excitation is almost the same as the situation in Fig. 4(a). Hence the rotor end saliency is not apparently reflected as the stator winding flux linkage. Consequently, the saliency modulation signal strength will be very weak due to the large pole number of the rotor end. 2) M2 Machine The M2 machine has fractional slot number. The magnetic paths of the stator excited flux are unevenly distributed in the rotor as shown in Fig. 5 and Fig. 6. In Fig. 5, the 18 slot distributed winding is configured with 8 magnetic poles and the number of slot per pole per phase \ud835\udc5e is 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000257_pedes.2006.344383-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000257_pedes.2006.344383-Figure4-1.png", + "caption": "Fig. 4. Phasor diagram of stator flux, rotor flux and stator currents.", + "texts": [ + " Both, stator and rotor fluxes can be expressed in terms of stator and rotor currents as follows: Hw HTE S(j) SP) SP3) S(4) S(5) S(6) 1 V2 V3 V4 V5 V6 VI 1 0 vo V7 vo V7 vo V7 -1 VI V2 V2 V3 V4 V5 1 V3 V4 V5 V6 VI V2 -1 0 V7 Vo V7 Vo V7 Vo -1 V5 V6 Vl V2 V3 V4 Eliminating derived: Lm Ir from (2), the following equation can be -fft -~~~~~~~~~~~~~~~~~~~~~~~~~~~~ s Lr r )s where Ls =LsLr-Lm2 Thus, stator current can be expressed as I L= Ks _-rny (5)s LS( Lr )s(5 By substituting (5) into (1), the following expressions for the torque can be derived: 3P Lmn Te = 2 L , tYr x Yfse2 LrLs Thus, the torque modulus is: T 3P Lm' -Isn,Te =- m,LL r si 2LrLs (6) (7) where)y is the angle between stator and rotor fluxes, Fig. 4 shows the phase diagram corresponding to (7). Fig.3. Flux trajectory and VSI voltage vectors and the corresponding flux increment. The electromagnetic torque in the three-phase induction motor can be expressed as follows: 3 Te =- ts x s (1) 2 where y's. is the stator flux, Is is the stator current and p is If the rotor flux remains constant and the stator one is changed incrementally by means of the stator voltage Vs, and the corresponding change of y is Ay, the incremental torque ATe can be expressed as follows: 3 L ATe =p m+A sn,tTe2= LrLs |r *+s+ sin)y (8) 2 LrLs The estimator block calculates stator flux and torque feedback signals" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002274_j.precisioneng.2021.01.004-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002274_j.precisioneng.2021.01.004-Figure1-1.png", + "caption": "Fig. 1. Experimental setup: (a) photo, (b) illustration of the robot.", + "texts": [ + " In a parallel link mechanism, the heavy actuator parts are located on the body base, and the leg parts become light-weight. The robot dynamics becomes similar to a concentrated body mass with a light-weight leg. As for the second point, a leg spring is set on the knee so as to become parallel to the line of the body and the tiptoe. By this setting, the elastic force becomes parallel to the line length of the body and the tiptoe, or leg length. These two points realize dynamics near a spring-mass dynamics. Fig. 1 shows a photo and a kinematic illustration of the experimental robot. The robot body is constrained to the vertical axis for the simplicity of experiments. The robot leg is not constrained. The robot employs a parallel link mechanism. The two motors are placed opposite on the sides of the body. An incremental encoder is attached to each motor. A linear spring is set near the leg joint as Fig. 1 (b). For measurement of the body height, the two-link mechanism as Fig. 1 (a) and an incremental encoder is attached. This paper uses three coordinates: the world coordinate, the base coordinate, and the polar base coordinate. The definitions of the coordinates are as follows. The world coordinate is an orthogonal coordinate in which the origin is the cross point of the ground and the vertical axis. The robot dynamics for simulation are calculated in the world coordinate. The base coordinate is an orthogonal coordinate in which the origin is the axis of the two actuators in the saggital plane. The base coordinate is used for control. The polar base coordinate is a polar coordinate in which the origin is the axis of the two actuators in the saggital plane. The polar base coordinate is used for the evaluation of leg spring. Jacobian matrixes in the world coordinate and the base coordinate are induced. From Fig. 1 (b), the tiptoe position in the world coordinate xtW = [xtW, ztW] is expressed as [xtW , ztW ] = [l(sinq1 \u2212 sinq2), zb \u2212 l(cosq1 + cosq2)], (1) where l is link length, q = [q1, q2] is the vector of the actuator angles, and zb is height of body in the world coordinate. Because the robot body is constrained to the vertical axis, the variables of the motion equation in Y. Abe and S. Katsura Precision Engineering 69 (2021) 36\u201347 the world coordinate are qw = [zb, q1, q2]. The Jacobian matrix in the world coordinate JacoW can be induced from partial derivative by qw as JacoW = \u2202xtW \u2202qw = \u23a1 \u23a2 \u23a2 \u23a2 \u23a3 \u2202xtW \u2202zb \u2202xtW \u2202q1 \u2202xtW \u2202q2 \u2202ztW \u2202zb \u2202ztW \u2202q1 \u2202ztW \u2202q2 \u23a4 \u23a5 \u23a5 \u23a5 \u23a6 = [ 0 lcosq1 lcosq2 1 lsinq1 lsinq2 ] ", + " (41) The amplitude A and the phase \u03c6 are calculated from the complex number \u03be. This section describes the experimental results and the analytical simulations. Firstly compensation controls are compared in free motions. Second, the experimental results of the proposed control system are presented. Third, the experimental results are analyzed by comparing with the ideal simulations. Finally, the effect of the parameter variation of the leg spring is confirmed experimentally. The experimental setup is shown in Fig. 1. The leg robot employs a parallel-link mechanism with a linear leg spring. The model number of the spring is TY-21. A softer spring TU-21 is also used for comparison in Section 5.4. The two actuators are placed opposite to each other. The model number of the motors is EC-i 40, 100W, 488607. As the option, the planetary gear head 6:1 is attached to each motor. The model number of the motor drivers is EPOS2 50/5. The output of the planetary gear head is connected to the harmonic gear head 21:1. The model number of the harmonic gear heads is HPG-14-21A" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000522_bfb0119424-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000522_bfb0119424-Figure1-1.png", + "caption": "Figure 1. The Melbourne Dextrous Finger, (a) a CAD model, (b) a schematic of the kinematic chain, and (c) its graph, which is extended in (d) to account for loads applied at the finger-tip and actuators.", + "texts": [], + "surrounding_texts": [ + "Davies' virtual power method [1, 11] for the analysis of loads in mechanical networks draws an analogy with current in an electrical network. The linkage can be represented as a graph, where the bodies axe nodes and the joints are arcs. A set of independent loops in the linkage axe identified and a circuit wrench is assigned to each. The circuit wrenches are unknown but must satisfy Kirchhoff's nodal law, which is essentially the condition of static equilibrium. The reaction in any joint is then the sum of the circuit wrenches passing through the corresponding arc in the graph. Davies' work should be consulted for the details of the method, but for brevity here we give only an application of it to the Melbourne finger, which is shown in Figure l(a). The kinematic chain of this device is shown in Figure l(b) and this is an intermediate step towards the graph shown in Figure l(c). The details of the drive have been omitted for simplicity. For each force or load applied to the linkage another path is added to the graph between the bodies acted on by the force, thus increasing the number of loops in the graph by one. The additional loops formed when finger-tip and actuators loads are included is shown in Figure l(d). The number of freedoms in the added paths is governed by the dimension of the space spanned by the applied force. For the actuators where a simple force is to be applied, the space is one dimensional and there must be (6 - 1 --)5 freedoms in the added path. Assuming a simple frictional point contact is made between the finger-tip and the grasped object, any pure force acting through the centre of the finger-tip is admissible and these forces span a three system. Consequently there are (6 - 3 --)3 freedoms in the added path around the finger-tip. There are seven independent loops in Figure l(d). Each freedom in the graph is assigned an arbi t rary direction and, using these, a loop matr ix B can be constructed with one column for each joint and one row for each loop (i.e. B is a 7 x 39 matrix). The elements of the matr ix indicate the freedoms traversed by each loop and whether the sense of the joint corresponds to the sense of the loop. We next need to form the motion screw matr ix M which describes the joint freedoms. The screw coordinates of the freedoms are expressed in a global coordinate frame and stacked to form M with one set of screw coordinates per row, (i.e. M is a 39 x 6 matrix). For the Melbourne Finger it is easy to write the screw coordinates symbolically so that they follow the movement of the linkage [12]. The freedoms in the added action loops about the actuators and finger-tip must be linearly independent and reciprocal [13] to the wrenches which are to be applied to the linkage at tha t point. The freedom matr ix F of the linkage can be assembled by taking each row of the loop matr ix B, writing it as a diagonal matrix, and multiplying it by the motion screw matr ix M. The freedom matr ix F is the concatenation of the resulting matrices. For the case being considered F is a 39 \u2022 42 matrix. An action screw matr ix A is formed from the unknown circuit wrench coordinates by writing them all in one column in the appropriate order. Kirchhoff's nodal law is then expressed as: FA = 0. (1) For our case, this gives 39 expressions to be satisfied by the 42 unknowns, and the three independent components of the circuit wrenches are directly related to the actuator forces. Reducing F to row echelon form gives expressions of which each has one dependent component of the circuit wrenches in terms of the independent ones. The reactions in each of the joints can then be found after forming a matr ix .4.* which has the coordinates of one circuit screw on each row. The joint reactions R are then found by: R --- BTA * (2) into which expressions for the dependent components of the circuit wrenches can be substituted to give R in terms of only the actuator efforts. The work so far has not considered the effects of friction, but to include it provides no conceptual difficulty, despite the increase in the size of the matrices involved. For each joint where the friction is to be considered another path must be added to the graph to form a loop around the joint of interest. The path contains screws that are reciprocal to the friction wrench in the joint, and their number depends on the number of freedoms permitted by the joint. A cylindrical joint requires four and a simple hinge needs five. The magnitude of the friction wrench is considered to be an independent quantity in the first stages, and we are able to calculate its effect on the joint reactions throughout the linkage. The reaction at a joint can be decomposed into pure force components which are orthogonal and aligned with the extent of the bearings. The magnitudes of these components are multiplied by a friction coefficient, determined from manufacturers' data sheets or tables of approximate coefficients. From this, the magnitude of the friction wrench can be determined and this is substituted into Equation 2. As a result, the joint reactions change and we need to iterate until the solution converges. One problem not yet addressed is the question of the sense that the friction wrench takes. When the linkage is in motion the sense of movement is determined and the sign to apply is clear. However, for the static case the direction of impending motion is needed, and here two alternative assumptions are possible. The first is that the friction wrenches oppose the forces applied by the actuators. In the experiments that follow, the finger-tip load is applied by a passive force/torque sensor, and we believe this first case to be appropriate. The other alternative is that the friction acts to assist the actuators in supporting the load. This would be common in the normal operation of the hand, when the grasped object is being pushed, by some fingers, into another and thus back-driving that finger. Interestingly, these two cases are not symmetrical about the frictionless results. The actual performance of the finger will lie between these two extremes and, for 'good' performance a narrow range is desirable. We assume that if free the finger-tip would move in a direction that is parallel to the finger-tip force, and hence presume this to be the impending finger-tip velocity. From the presumed finger-tip velocity the corresponding joint velocities are determined. 4. E x p e r i m e n t a l P r o c e d u r e The experiment observed the relationship between the static finger-tip forces and the actuated forces. The finger-tip forces were measured by constraining the spherical finger-tip in a specially designed ball and socket clamp fixed to a six-axis ATI FT3031 force/torque sensor. Since the connection between the finger-tip and sensor is equivalent to a ball and socket joint, only pure forces acting through the centre of the finger-tip were measured, see Figure 2. When connected in this manner, the combined finger and sensor have a mobility of zero, and consequently the actuator forces produce no motion and provide internal loading of the linkage only. It is the reactions to this internal loading that are measured by the force/torque sensor. Owing to its inherent complexity, the drive system was seen to distract from the specific questions of this work. Therefore, the drives of the finger were removed, and the finger and sensor were mounted vertically so that actuator loads could be applied by suspended weights hanging directly from the sliders 8 g n 8 o r . where they protrude from the back of the finger. This arrangement provided a direct measure of the forces applied to the linkage. The finger was set in a known configuration and the gravity loading of the finger linkage was zeroed out of the force/torque sensor signal. Known actuating forces were applied to the linkage and the resulting finger-tip loads were measured for several combinations of actuator loads. The finger-tip load was determined by sampling the force/torque sensor at 800Hz and averaging one second's worth of data. Three configurations were tested and, in all, 85 sets of measurements were taken. 5. R e s u l t s The results collected are shown in Figure 3(a) and (b) for the x-, y- and zdirections. The coordinate system is arranged with the positive z-direction vertically up in Figure 2, the positive x-direction is, more or less, towards the viewer, and the positive y-direction is to the right of Figure 2 so that a force in this direction presses straight into the force torque sensor without applying a moment. The graphs compare components of the measured finger-tip forces with the predictions of the model for the three configurations. The circles in Figure 3(a) indicate the predicted forces when no joint friction is modelled. The dashed line shows where predicted and measured finger-tip data are equal and it can be seen that the model overestimates the finger-tip force, which is as expected. The lines of best fit to the data are found and these have gradients of 0.82, 0.76, and 0.94, for the x-, y- and z-directions. From these we can infer an efficiency for the linkage of about 80% and, although the actual efficiency is dependent of the configuration, we see that it is reasonable consistent over the three configurations shown. The r 2 values for the regressions give an estimate of the spread from a strict linear relationship; they are 0.993, 0.927, and 0.995 which shows that there is a good linear relationship. With a passive force/torque sensor connected to load the finger-tip, one would expect the joint friction to reduce the predicted loads. Including friction in the model alters the predicted finger-tip loads as shown in Figure 3(b). Here the predictions with friction opposing the actuator efforts are marked with an %' and those with friction assisting the actuator efforts to support the fingertip load are marked with a 'A'. The dotted lines between the two sets indicate the predicted range in which the performance of the finger lies; on one edge the finger is driving the grasp and on the other it is being back-driven by the force applied through the grasped object. Figure 3(b) shows that the data with friction opposing the actuator forces is very close to the measured finger-tip forces. The gradients of the lines of best fits are 0.961, 0.945, and 1.000 for the x-, y- and z-directions, showing that the friction model now accounts for the measured finger-tip loads. The corresponding r 2 values are 0.983, 0.919, and 0.994 which indicates that the quality of the fit has dropped slightly but is still very good. The triangles in Figure 3(b) show the difference between the cases when friction opposes the actuators and when it assists the actuators in supporting the finger-tip load. The size of the difference determines the responsiveness of the finger when applying forces. If during the manipulation of an object, the finger-tip changes the direction of its motion from towards the object to away from it while always maintaining a squeezing force then a step change in the actuator effort is required to maintain the same grasping force. The model predicts that this effect is about 30% of the finger-tip load in the xand y-directions and, as such, this will be of some importance in strategies for synthesising and controlling the grasp. Design changes which reduce the disparity between the two cases will make the control task easier. For the graphs in Figure 3 there was no special tuning of the coefficients of friction to achieve the good fit shown. Thus these figures have been produced without information that is unavailable to the designer or cannot be reasonably estimated. Nevertheless, the results are sensitive to those coefficients which contribute most significantly to the reduction in the loads measured at the finger-tip. For one particular configuration and loading we can breakdown the contribution of the friction in each joint to the overall lowering of performance. Figure 4 shows a diagram of the finger with the approximate proportion of dif- ference between the frictionless and friction model. The bracketed percentages apply to the case where friction loads assist the actuator forces in supporting the finger-tip load. It can be seen that three joints account for about 80% of the performance loss owing to friction. An assessment of the reactions in these joints shows that the moment reactions lead normal forces that are an order of magnitude greater that the direct reactions, and this suggests one way in which the design of the linkage can be improved. Any effort to replace other journal bearings in the linkage with rolling element bearings appears to be of little benefit. 6. C o n c l u s i o n s We have applied Davies' virtual power method to analyse the role of joint friction in a multi-loop linkage. We found that static and finger-tip loads applied by a passive device are well described by the model when the frictionless values are adjusted with friction opposing the actuator loads. Moreover, the predictions are based only on information known by the designer before the linkage is constructed and no fine tuning of the friction coefficients has occurred. Therefore the model seems a useful tool for design. Comparing the frictionless predicted and measured finger-tip loads, an efficiency of the linkage can be inferred; for the Melbourne Dextrous Finger this efficiency is about 80%. An important property of the finger is its ability to be back-driven. In this case friction acts to assist the actuators in supporting the load on the finger-tip. The model can be used to predict this property, but as yet we have not been able to measure this effect and it is par t of our on going work. The normal operat ion of the hand is then expected to lie in a range between the case where the finger is driving the grasp and the finger is being driven by the grasp. The model predicts tha t the dead-band so formed is of the order of perhaps 30% of the finger-tip load, al though the exact amount is dependent on the configuration of the finger. We believe this has serious consequences for the way in which grasps are planned and executed by dextrous robot hands. To this point in t ime we are unaware of any a t t empts in the grasping l i terature tha t takes proper account of such an effect. The analysis of individual configurations and loading cases shows how the friction in individual joints contributes to the overall behaviour of the finger. We show tha t most of the friction is a t t r ibuted to just three joints and the analysis suggests a means to improve the design of the finger. Furthermore, for control purposes it seems sufficient to model only the friction in these three joints and this is a substantial saving in effort. R e f e r e n c e s [1] T. H. Davies. Circuit actions attributable to active couplings. Mechanism and Machine Theory, 30(7):1001-1012, 1995. [2] B. Armstrong-H~louvry. Control of Machines with Friction. Kluwer Academic Press, Boston, 1991. [3] C. R. Tischler and A. E. Samuel. Predicting the slop of in-series/parallel manip- ulators caused by joint clearances. In Advances in Robot Kinematics: Analysis and Control, pages 227-236, Dordrecht, 29th June-4th July 1998. Kluwer Academic Publishers. [4] J. Swevers, C. Ganseman, D. B. Tiikel, J. De Schutter, and H. Van Brussel. Optimal robot excitation and identification. IEEE Transactions on Robotics and Automation, 13(5):730-740, 1997. [5] M.R. Elhami and D. J. Brookfield. Identification of Coulomb and viscous friction in robot drives: An experimental comparison of methods. Proc. IMechE C, Journal of Mechanical Engineering Science, 210(6):529-540, 1996. [6] M. R. Elhami and D. J. Brook field. Sequential identification of Coulomb and viscous friction in robot drives. Automatica, 33(3):393-401, 1997. [7] C. Canudas de Wit. Experimental results on adaptive friction compensation in robot manipulators: low velocities. In Experimental Robotics I, pages 196-214, Montreal, June 19-21 1989. Springer-Verlag. [8] B. Armstrong and B. Amin. PID control in the presence of static friction: A comparison of algebraic and describing function analysis. Automatica, 32(5):679- 692, 1996. [9] K. Haln. Applied Kinematics. McGraw-Hill, New York, 2nd edition, 1967. [10] B. Paul. Kinematics and Dynamics of Planar Machinery. Prentice Hall, Eagle- wood Cliffs N J, 1979. [11] T. H. Davies. Mechanical networks--III: Wrenches on circuit screws. Mechanism and Machine Theory, 18(2):107-112, 1983. [12] C. R. Tischler. Alternative Structures for Robot Hands. Ph.D. thesis, University of Melbourne, 1995. [13] K. H. Hunt. Kinematic Geometry of Mechanisms. Clarendon Press, Oxford, 2nd edition, 1990." + ] + }, + { + "image_filename": "designv11_83_0001652_bf03177400-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001652_bf03177400-Figure2-1.png", + "caption": "Fig. 2. Laser vision sensor - internal structure.", + "texts": [ + " In this study, the sensor size needed to be large enough to measure the repair area with suitable resolution, but small enough to avoid contact with the internal wall of the reactor head. First, the maximum measurement area was determined, and then the sensor size and structure designed to obtain reasonable resolution. The maximum weight of the sensor was set at 1.2kg, based on the allowable load of the robot. The structure of the sensor was designed with a double laser due to the difference in the laser scan line in the left and right directions of the CRD nozzle. Fig. 2 shows the optimally designed laser vision sensor and Table I gives the design specifications. (a) -------~-,--.---- The image shown in Fig. 4(a) acquired by the CCD camera includes the laser line and noise due to double reflection. It is necessary to pre-process the image 2.4 Image Processing path through image processing, 3) transmitting the generated weld-path to the robot through serial communication and 4) performing the auto-welding. The laser vision sensor is comprised of a double diode laser, a CCD camera and an optic system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002460_978-3-030-59608-8_27-Figure8.8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002460_978-3-030-59608-8_27-Figure8.8-1.png", + "caption": "Fig. 8.8 Haptic device for precision tools grasping. Note some of the components used in the figure were adapted from [38]", + "texts": [ + " Although virtual simulation is being applied across a wide range of medical training applications, currently the majority of these applications focus on the cognitive aspects of a procedure only typically ignoring the technical components given the complexities associated with generating the haptic cues required to simulate them. The aim of this project is to improve the medical training simulation by providing haptic feedback in virtual medical skills training. We are focusing on simulating the grasping and manipulating of precision tools (e.g., a scalpel, etc.) similar to commonly available haptic gloves (see Fig. 8.8). We determined the force required to lift an object at rest with a known mass based on Newton\u2019s law of gravity. Grasping and manipulating precision tools such as a scalpel involves the thumb and the index finger. According to Nataraj [54], the difference between the magnitude of the forces applied by the thumb and index finger is negligible. Taking this into consideration, we distributed the forces equally between the thumb and the index finger. The haptic device then goes through a series of conversions of these forces to provide equal and opposing force" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003132_978-981-16-0119-4_20-Figure71.6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003132_978-981-16-0119-4_20-Figure71.6-1.png", + "caption": "Fig. 71.6 Exploded view of throttle pedal [8]", + "texts": [ + " The details of mass properties and the modified complexity factor of the consolidated design are shown in Table 71.1. From Table 71.1, the MCF value of the consolidated design is more than 44, so the design is suitable for manufacturing using AM [6]. A throttle pedal design is selected to demonstrate the proposed framework and validate the effectiveness of the new method with the Part Consolidation Candidate Detection (PCCD) algorithm used in [8]. The original design is taken from an opensource CAD database [11]. As shown in Fig. 71.6, the throttle pedal consists of 12 parts without counting the washers and fasteners. The product network is created by identifying the physical contact between the parts. From the network, the component with the highest centrality score is identified as the candidate for part consolidation. Based on the physical relationship between the components, the product network is drawn using the open-source network tool and is shown in Fig. 71.7. The centrality score of each node in the product network is measured and shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002241_j.actaastro.2021.02.001-Figure11-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002241_j.actaastro.2021.02.001-Figure11-1.png", + "caption": "Fig. 11. Two people facing each other with their heads oriented in the same direction, as seen from above. (a) F is exerted along a line which crosses \ud835\udc29\ud835\udc50\ud835\udc5c\ud835\udc5a,1- \ud835\udc29\ud835\udc50\ud835\udc5c\ud835\udc5a,2= \ud835\udc34\ud835\udc35. (b) F is exerted along a line which does not cross \ud835\udc34\ud835\udc35.", + "texts": [ + " The influence of the hand and arm locations on v\ud835\udc56 is not significant. When manipulating the model\u2019s arm, the upper arm is kept as close to the torso as possible. For the cases when |r\ud835\udc56| \u2248 30.5 cm\ud835\udc56 and 45.7 cm, v\ud835\udc5a\ud835\udc56\ud835\udc5b tilts about 1.3\u25e6 and 0.7\u25e6 from the AP axis, respectively. Thus F will produce a \ud835\udf49 closely aligned with v\ud835\udc5a\ud835\udc56\ud835\udc5b in either case. If we would like to initiate rotation about the same \ud835\udf4e direction for both \ud835\udc56 then r\ud835\udc56 \u00d7 F\ud835\udc56 must have the same direction. To satisfy this condition, the line on which F\ud835\udc56 lies must cross \ud835\udc34\ud835\udc35, see Fig. 11(a). In a situation such as this, we find that \ud835\udf491 and \ud835\udf492 are both oriented in the \u2212?\u0302? direction. Recall, that to ensure there is no translational movement of either \ud835\udc56, the F\ud835\udc56 vectors would need to be oriented from a point on \ud835\udc34\ud835\udc35. To cause 1 and 2 to rotate in opposite directions, we need to position F\ud835\udc56 such that r\ud835\udc56 \u00d7 F\ud835\udc56 produces \ud835\udf49 \ud835\udc56 vectors with opposite signs. This can be performed when ensuring F\ud835\udc56 are on a line which does not cross \ud835\udc34\ud835\udc35. This arrangement of vectors can be seen in Fig. 11(b). Note it is impossible to position F\ud835\udc56 on the line segment \ud835\udc34\ud835\udc35. Thus some F\ud835\udc47 component will be present. It may be that combining the primary rotation force with a secondary directing force may allow for rotations in which no translational component exists. I have not pursued that line of inquiry further at this time however. We have worked through some aspects of these maneuvers on aerial harnesses. When desiring rotation in the same direction, we have found that pulling rather than pushing is the most efficient and reliable method of force transfer", + " Attempting to push hand-to-hand against the palms and orient F\ud835\udc56 optimally to cause rotation is something akin to succeeding at a jumping-high-five. The maneuver is difficult to align and sliding past the other person\u2019s hand, even when contact does occur, is common. Instead, using the same arm (both left or both right), reaching around to grasp palm-to-palm, essentially grabbing the other person\u2019s thumb, offers multiple points of contact, and allows for subtle refinements in positioning as \ud835\udc56 reach agreement regarding the alignment. Consider the coordinate system shown in Fig. 11(b). This coordinate system is in the transverse plane which is mapped by unit vectors ?\u0302? \u2032 (along the AP axis) and ?\u0302? \u2032 (along the transverse axis) of \ud835\udc56 in the initial positions shown. Note this primed coordinate system is inertial unlike the original un-primed coordinates which were associated directly with the body. Imagine a unit circle surrounding each figure and tracing the direction the person is facing on the unit circle such that in Fig. 11(b), 1 is facing + ?\u0302? \u2032 and 2 is facing \u2212 ?\u0302? \u2032. If 1 and 2 have v\ud835\udc5a\ud835\udc56\ud835\udc5b parallel to ?\u0302? with \ud835\udc63\ud835\udc5a\ud835\udc56\ud835\udc5b,1 = \ud835\udc63\ud835\udc5a\ud835\udc56\ud835\udc5b,2, and they apply torques \ud835\udf491 = \ud835\udf492 on each other, they would begin rotating in the same direction with the same angular velocity |\ud835\udf4e2| = |\ud835\udf4e1|. If their initial orientations are those shown in Fig. 11(b), the ?\u0302? \u2032 component of 1 would initially decrease in magnitude, then follow sinusoidal oscillations along the axis. Starting from the same figure, the ?\u0302? \u2032 component of 2 would initially increase in magnitude then follow sinusoidal oscillations along the axis. These components for 1 and 2 are shown in Fig. 12(a) as the two top plots. The rotational velocity magnitudes are also plotted and are constant. Notice that when \ud835\udc651 and \ud835\udc652 are at their peak and trough, respectively, 1 and 2 are facing each other" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001904_2008-36-0289-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001904_2008-36-0289-Figure2-1.png", + "caption": "Figure 2 \u2013 Oil film pressure for (2a) flat regular bearing and (2b) axial profiled bearing.", + "texts": [ + " A modern four cylinders high speed engine was used to study the influence of axial profiled con-rod bearing. Details of the engine and bearing are given in the table I. Table 1. Engine main data for EHL simulation. Engine Diesel; 4 cylinder in line Capacity 1.560 liters Bore 75 mm Stroke 88.30 mm Connecting rod length 136.80 mm Lubricant SAE 5W30 (EZL1315) Oil gallery temperature 140\u00baC Rod material Steel Crankshaft material Steel Bearing diameter 47.033 mm Bearing length 16.750 mm Bearing/shaft clearance 50 \u03bcm Figure 2 presents the peak oil film pressure after 12o of top dead center for working condition of 156 Bar at 4,500 rpm, which represents the critical position for upper con-rod bearing. Note that picture 2a was generated with flat bearing while figure 2b takes in consideration axial profiling with maximum bearing wall thickness on the central region of the bearing length. The minimum bearing wall thickness is located on the edge regions which is 5 \u03bcm thinner the central region. The transition from the edges to central region was considered a flat ramp with length of 5 mm. Flat bearing (figure 2a) presented maximum oil film pressure of 176 MPa on the bearing edges with contact pressure of 74 MPa while the axial profiled bearing showed completely different shape for oil film pressure distribution with maximum value of 193 MPa localized on the central region of the bearing length and minimizing contact pressure. Another key feature considered for bearing performance is related to the oil film thickness presented on the figure 3. Once more is clear the advantage of the axial profiled bearing, even though it (3b) presented lower film thickness on central region when compared with regular flat bearing (3a) the axial profiled bearing (3b) showed minimum oil film thickness of 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002827_edunine51952.2021.9429151-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002827_edunine51952.2021.9429151-Figure1-1.png", + "caption": "Fig. 1. the number 1 corresponds to the board jetson NX, the 2 corresponds to the propellers, the 3 to the engines, the 4 to the stereo camera zed, the 5 corresponds to all the electronic components and the flight controller, the 6 to the batteries, for last the 7 to the drone GPS.", + "texts": [], + "surrounding_texts": [ + "As mentioned, the project to be implemented corresponds to a drone with artificial vision and autonomous movement capabilities to carry out object search and reconnaissance tasks. Taking into account that the undergraduate degree in Systems Engineering is mainly focused on software engineering, it has been a great challenge for students to acquire and apply knowledge in 3D printing, electronics, aerodynamics, radiocontrol, embedded systems and artificial intelligence[5]. Artificial vision and autonomous motion capabilities require that most of the data processing that the device constantly reads from the environment, be carried out inside it. Here it was necessary to evaluate different options for embedded systems such as Raspberry Pi and NVIDIA Jetson, opting for the latter, more precisely the NVIDIA Jetson Tx2, which offers the necessary computing power to support the operating system and artificial intelligence algorithms [6]. Video system uses a ZED stereo camera, which allows it to control its orientation and generate 3D maps of the environment. Complete system is monitored through a ground station developed with the Javascript programming language, which allows specifying the flight plan, as well as monitoring its variables and viewing the video stream. Student skills in Software Engineering were put to the test with the development of the artificial vision and autonomous motion components, which are managed by the Robot Operating System - ROS. In this way, the drone is able to recognize and avoid different obstacles present in the environment, such as trees, lighting poles and aerial cables." + ] + }, + { + "image_filename": "designv11_83_0002241_j.actaastro.2021.02.001-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002241_j.actaastro.2021.02.001-Figure3-1.png", + "caption": "Fig. 3. Visual representation of point model. Small blue dots represent joint locations. Larger red dots represent the center of mass of body segments. The black dot represents \ud835\udc29\ud835\udc50\ud835\udc5c\ud835\udc5a. In this position, \ud835\udc29\ud835\udc50\ud835\udc5c\ud835\udc5a occludes the center of mass of the lower torso.", + "texts": [ + " do not report the standard deviations of the measurements, giving no solid ground for reporting reliable standard deviations for the models which will be described shortly. The model developed for this paper was created by reconstructing a human figure from the Drillis data. The joints are linked in a \u2018\u2018parent\u2019\u2019 and \u2018\u2018child\u2019\u2019 manner [3]. The base of the torso is considered the parent of all other joints, and joints which are closer to the base are parents of joints which are further from the base. A visualization of a model, which will be referred to as \u2018\u2018the point model\u2019\u2019, is shown in Fig. 3. The model is built in three dimensions with rotational control over each joint through polar and azimuthal angles. The center of mass of the entire body can be calculated... p\ud835\udc50\ud835\udc5c\ud835\udc5a = \u2211 \ud835\udc56 p\ud835\udc56 \ud835\udc5a\ud835\udc56 (2) Since the body will always rotate around its center of mass p\ud835\udc50\ud835\udc5c\ud835\udc5a and the location of the center of mass relative to the rest of the body can change when the body is manipulated, knowledge about this location is extremely valuable when planning or executing movements. As will be shown in Section 4.1, the model generated center of mass calculations are consistent with measurement [12]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002881_s42235-021-0043-x-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002881_s42235-021-0043-x-Figure2-1.png", + "caption": "Fig. 2 Realization of linear reciprocating oscillation motion. (a) Photograph of the transmission mechanism; (b) top view of the transmission mechanism.", + "texts": [ + " Section 4 compares the formulated models of the wing stress of the oscillating-wing mode and flapping-wing mode. In section 5, the force measurement and observation of the wing movement are presented to verify the model and liftoff of the vehicle is demonstrated. The main conclusions drawn from the study are summarized in section 6. A combination of the crank-slider mechanism is designed and manufactured to realize the symmetrical reciprocating oscillating motion of the two wings in a straight line, as shown in Fig. 2a. A miniature brushless motor (KV rating: 2200) placed at the middle of the rack bottom drives the rotation of the crank and motion of two couplers and two output sliders after decelerating through a reduction gearbox (reduction ratio: 1:50). The two wings connected to the sliders move in opposite directions at the same instantaneous rate. The crank-slider mechanism has the advantage of a simple structure and is therefore commonly used in the transmission design of MAVs. The other end of the crank was also restrained to stabilize the transmission. The crank, couplers, and sliding tracks are all made of carbon fiber materials. The mass of the miniature brushless motor and reduction gear box is 15 g. The overall mechanism has the characteristics of light weight and good transmission performance. The top view of the transmission mechanism is shown in Fig. 2b. The inertial coordinate system is Zhou et al.: Liftoff of a New Hovering Oscillating-wing Micro Aerial Vehicle 651 represented by [X, Y, Z]T, and the body-fixed coordinate system is indicated by [x, y, z]T, with corresponding unit vectors [i, j, k], which is attached to the center of mass of the MAV. [m, n, p] is used to denote the unit vectors of the wing coordinate system. With the right wing being taken as an example to establish the model of oscillating motion, which can also be extended to the left wing in a straightforward manner, the wing displacement, X(t) is calculated as: 2 2 2 ( ) = cos(2\u03c0 ) + 1 sin (2\u03c0 )" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002563_012068-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002563_012068-Figure1-1.png", + "caption": "Figure 1. ROV Desain", + "texts": [ + " ROV is very useful in carrying out underwater tasks, especially in very deep and dangerous waters that cannot be carried out by humans such as exploration of hydrothermal sources, monitoring and maintenance of submarine pipes, construction and installation of marine platforms, exploration and study of marine habitats and military operations such as reconnaissance and investigation. In making the ROV, it must pay attention to the parts that must be owned by the ROV in carrying out its duties. In Figure 1 is a general description of the ROV design part as a requirement for making a design consisting of a frame, thruster, control system / box, and camera. OPERATED UNDERWATER VEHICLE (ROV) using an ARTIFICIAL NEURAL NETWORK (ANN) smart system. The way the system works is done by reading the input in the form of an Accelerometer and Gyroscope sensor which is then processed using a microcontroller and the output is PWM (Pulse Width Modulation) which is interpreted by the ESC (Electronic Speed Control) driver to move the motor according to the speed it should be" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002772_978-3-030-67750-3_14-Figure26.2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002772_978-3-030-67750-3_14-Figure26.2-1.png", + "caption": "Fig. 26.2 a Design 1. b Design 2", + "texts": [ + " These two designs are using the systems that use a barrel-like principle (Abdullah and Jamil 2015), as in Fig. 26.1a. Basic components in the designs can be termed as in Fig. 26.1b (Abdullah and Jamil 2015). The main advantage of this design is the possibility to separate the housing from the top and bottom cover that makes this concept easy to set the parameter of the device. By disassembling the part of the system, the coil and magnet can be modified, which made the parameter setting simpler. Both conceptual designs can refer to as in Fig. 26.2a,b, where Fig. 26.2a shows Design 1, while Fig. 26.2b shows Design 2. For design 1, a set of two coils is placed outside the magnet. All the linkages are connected to each top and bottom mount. The main shaft is from the bottom linkage connected to themagnet, and themagnet\u2019smotion is controlled by the spring,which is Parameters Design 1 Design 2 Magnet height, mm 200 200 Magnet radius, mm 100 100 Wound coil height, mm 220 220 Number of coil turn 20 10 Coil internal diameter, mm 140 140 Distance between magnet and coil, mm 10 10 Wire diameter, mm 2 2 attachedbelow themagnet" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002118_s11071-020-06158-5-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002118_s11071-020-06158-5-Figure1-1.png", + "caption": "Fig. 1 Cylindrical coordinates with origin \u2018O\u2019 at the centre of the middle surface z = 0 of the plate. \u2018P\u2019 is the material point on the top surface z = h/2, where h is the plate thickness", + "texts": [ + " The values of \u03b3(m,n) for a thin circular plate of radius R are read from Table 1, which is obtained by solving the free vibration equation, \u03c1hw\u0308 + Eh3 12(1 \u2212 \u03bd2) \u22074w = 0, (3) together with the free boundary conditions of the plate, given as [\u22072w \u2212 (1 \u2212 \u03bd) 1 r ( w,r + 1 r w,\u03c6\u03c6 )]r=R = 0 [(\u22072w),r + (1 \u2212 \u03bd) 1 r (1 r w,\u03c6\u03c6 ) ,r ]r=R = 0, (4) where (),r and (),\u03c6 denote derivative with respect to r and \u03c6. For a material point on the top surface of the plate, having coordinate (r, \u03c6, h/2), the acceleration components in the cylindrical coordinate system (Fig. 1) are given as, abr (r, \u03c6, t) = v\u0307br = \u2212h 2 w\u0308,r ab\u03c6(r, \u03c6, t) = v\u0307b\u03c6 = \u2212h 2 w\u0308,\u03c6 r abz(r, \u03c6, t) = v\u0307bz = w\u0308, (5) where abr , ab\u03c6 and abz denote material point accelerations along unit vectors er , e\u03c6 and ez , respectively. Also, the over-dot \u02d9( ) denotes time derivative. Here, subscript \u2018b\u2019 stands for the plate material point. The inplane position coordinate of the particle is denoted as {rp, \u03c6p} and is always assumed to be coincident with the undeformed plate coordinate {r, \u03c6} of the top surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002945_978-3-030-69178-3_15-Figure6.4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002945_978-3-030-69178-3_15-Figure6.4-1.png", + "caption": "Fig. 6.4 Scan areas combined and corresponding scanner placement", + "texts": [ + " The orange circle shows the maximum reach of the robot (1.65 m), and the turquoise circle shows the required safety distance calculated using the maximum speed of the robot (4.5 m/s). The red rasterised area displays the regions visible to the sensor within the required area. Table legs and other objects at the scan level cause the white areas in-between.While moving the scanner or any other object, the scan updates dynamically to match the new layout. Fig. 6.3 Scan areas of a scanner with gaps Figure 6.4 shows the simultaneous use of two scanners in the same locations as shown in Fig. 6.3, i.e., one scanner near the robot and a second one placed under the table in the middle. As the figure shows, the scan areas complement each other leaving almost no gaps. The area between the robot and the conveyor on top could be fenced from human access so there would be no need to scan it. Thefigures in the previous subsection describe the situationwhen the robot isworking within its full reach at its full speed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002906_iemdc47953.2021.9449573-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002906_iemdc47953.2021.9449573-Figure6-1.png", + "caption": "Fig. 6. Graphical representation of the thermal fault model in a quarter of a coil (3D)", + "texts": [ + " To get a better modeling of the local phenomenon in axial direction, the fault segment itself is split up into three different segments. The middle segment describes the fault segment, which can be seen in Fig. 5, and its height is chosen to a predefined value depending on the fault layer to determine a fault area. This fault area can be interpreted as the physical contact area of the copper of the shorted wires. The other two segments are used to get the same thickness of the remaining segments. Fig. 6 shows the separation of the fault segment in axial direction and their integration in the full model. For the sake of simplicity, just a quarter of the whole stator coil is shown, whereas the model considers the whole stator coil. This means that the model has a higher resolution in the fault region and models a physical realistic fault area. Shorting two layers inside one segment with full height would lead to a drastic overestimation of the fault area and would have a major effect on the maximum temperature rise as explained later on" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001650_pes.2007.386081-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001650_pes.2007.386081-Figure7-1.png", + "caption": "Fig. 7. Equivalent circuit for superimposed quantities under a forward fault", + "texts": [ + " The analyses in this section are built on an equivalent twosource system and the in-phase coefficient and angle \u03b1 depend on the angle between the sending and receiving end machines, but of course in a network with multiple machines and multiple interconnections between machines, it is often not possible to identify a simple two machine network like the one used above, so more researches would be made for such complex network. Consider the relay fixed at terminal M in the two-source system as shown in Fig.1. Rewrite magnitude comparison format of (2) to phasic comparison format as follows. oo && && 270 )1( )1( 90 |0| |0| < \u2032\u2212\u2212\u2212\u2032\u2206 \u2032\u2206\u2212\u2032\u2212+ < UkkU UUkk Arg RC RC (11) Fig.7 shows an equivalent additional fault circuit for forward fault in a two-source system, where ZY is the setting impedance. U&\u2206 and I&\u2206 are sudden-change quantities of voltage and current measured by the relay. Assume that the impedance of line MF is FZ , current distribution factor is gIIC && /\u2206= , then the measured impedance is CRZZ gF /+= . The fault components for forward fault can be described according to Fig.7 as follows. Then (11) can be rewritten as: oo 270 ])([)1( ])([)1( 90 < +++\u2212+ \u2212+\u2212+\u2212< SRCYRC SRCYRC ZkkZZkk ZkkZZkk Arg (12) In (4), 5.0=Ck and 15.0 << opk . In (5), 1=Ck . In (6), 5.0>== opRC kkk . Hence, 1>+ RC kk . Then o0 1 1 = \u2212+ +\u2212 RC RC kk kk Arg Let: \u23a9 \u23a8 \u23a7 +\u2212\u2212+=\u2032 \u2212+++\u2212=\u2032 )1/(])([ )1/(])([ CRSRCYY RCSRCYS kkZkkZZ kkZkkZZ (13) Simplify (12), it follows that: oo 27090 < \u2032\u2212 \u2032\u2212< Y S ZZ ZZ Arg (14) According to (14) and the detailed parameters of (4)-(6), the corresponding operation characteristics of (1) and (4)-(6) for forward fault can be drawn as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000237_pedes.2006.344346-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000237_pedes.2006.344346-Figure2-1.png", + "caption": "Fig. 2. Separate components for the core yoke and teeth in a 15hp, 4P, three phase induction motor.", + "texts": [ + " 3 shows the 60Hz B-H curve for the material EBG450 used in the simulations. Magnetization curve for teeth Magnetization curve for core 0.40 0.20 0.00 0 -DC magnetization curve 1000 2000 3000 4000 H (AT/m) 5000 6000 7000 Fig. 1. AC magnetization curves for the teeth and core yoke and the DC magnetization curve. Two components were created in the stator core, first one was for the core yoke and material property was set to be as core yoke saturation cure, and for the second one material property was set to be as teeth saturation curve [3] as shown in Fig. 2. 1.60 1.40 1.20- _ 1.00 a. m 0.80 Xm0.80 0.60 FEM analyses have been done for the two three phase, induction motors. Specifications of the analyzed motors are given in Table I and Table II. Fig. 4 shows the flux distribution in the mentioned 15 hp motor. Table IV shows the computational and tested line currents (A) for 5hp and 15hp 4P three phase induction motors at fullload. It is observed from the table that the percentage error between the computational and tested results reduces considerably in this case also if separate saturation curves for teeth and core yoke are used instead of one ac saturation curve for both the teeth and core yoke" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002654_s42417-021-00299-6-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002654_s42417-021-00299-6-Figure3-1.png", + "caption": "Fig. 3 Simplified torsional vibration model", + "texts": [ + " The main objective of this paper is to establish a simplified dynamic model of the DI-SO helical gear system with unsymmetrical inputs considering the time-varying meshing stiffness, the phase difference of the meshing stiffness between the two meshing pairs, time-varying backlash, meshing damping and comprehensive transmission error, and reveal the evolution law of the dynamic characteristics with different system parameters. The DI-SO helical gear system can be simplified as a torsional vibration model for two gear pairs which is shown in Fig.\u00a03 regardless of the deformation of the transmission shafts and the stiffness of the supports. There are three helical gears with two driving gears (i, k) and one driven gear (j), and they are represented by their base circles with radius rbi , rbj and rbk . In this system, Ti , Tk and Tj are the external torques acting on the three gears and Tk \u2264 Ti . i , k and j are the dynamic angular displacements of the three gears. Ji and Jj are the mass moments of inertia of the two driving gears, and Jk is the equivalent mass moments of inertia of driven gear ( Jk1 ) and the flywheel ( Jk2 )" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001851_978-3-540-74027-8_1-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001851_978-3-540-74027-8_1-Figure7-1.png", + "caption": "Fig. 7. Application of Viennese Waltz behaviour. Rotating on the spot", + "texts": [ + " There is no slip between the planet and the sun wheels. The planet wheel could be on the outside of the sun wheel, Fig. 5, in which case the rotation of the planet wheel is reversed. The robot is shown, Fig. 6, backing out of a corridor representing retreating motion, i. e. translating, and simultaneously rotating such that the head of the robot is ready to face an opponent in the direction of its exit path. If the radius of the virtual planet wheel is set to zero then the robot will turn on the spot, Fig. 7. If the radius of the sun wheel is set to zero and the planet wheel rotates on the outside of the sun wheel then the robot will rotate about a fixed point which is the sun wheel, Fig. 8. This occurs when the radii of the sun and planet wheels are set to infinite radius, Fig. 9. Viennese Waltz motion is programmed into an omnidirectional robot by using a model of a virtual sun and planet wheel, Fig. 10. The radii of each wheel depend on the required behaviour pattern. The contact point between the sun and planet wheels is an instantaneous centre of rotation, IC of R" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001982_20080706-5-kr-1001.01655-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001982_20080706-5-kr-1001.01655-Figure1-1.png", + "caption": "Fig. 1. System configuration", + "texts": [ + " Rather, the robot is composed of individually selected and obtained hardware components and is controlled by a software system arranged accordingly. This approach allows students using the equipment full access to all components, interfaces and source codes. Moreover, improvements and updates can be easily incorporated over the years. The robot arm provides a basis for the coupling of auxiliary components in order to develop a wide range of applications. The system configuration of the laboratory robot \u2013 mechanical and from a information processing point of view \u2013 is described in the following section. In Figure 1 the arrangement of the equipment is displayed. The central element is the robot portal. Beside the portal the control cabinet is arranged. The cabinet contains the complete control hardware including axis processors and robot control computer. The mechanical resp. kinematical composition is displayed in Figure 2. The axle drives constitute the link between robot control and mechanical components. They are connected via analog or digital interfaces resp. the implemented Controller Area Network (CAN) bus to the axis processors of the lower control level" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001352_1077546307076901-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001352_1077546307076901-Figure3-1.png", + "caption": "Figure 3. Flexible slider crank.", + "texts": [ + " The Lagrangian takes the form Le Te Ve Ye (8) Substituting equations (1)\u2013(6) into (8), and employing the Euler\u2013Lagrange equations, the governing equations of motion for a rotating and translating elastic beam can be expressed in the following matrix form: [Me] e [Ce] e [Ke] e Fe (9) where [Me], [Ce] and [Ke] are mass, equivalent damping, and equivalent stiffness matrices of a element, respectively {Fe} is a load vector of an element, and e is a vector of element variables. When formulating the mass matrix of the coupler, the mass of the slider should be taken into account. Transforming the governing equations of the elements such that they are expressed in terms of global variables, and then assembling the element equations, leads to the global governing equations of a slider crank mechanism. Figure 3 illustrates a set of five global variables ( i i 1 2 5) used in this paper when considering a flexible slider crank mechanism with each link modeled with one element of the type derived in Section 2.1. Based on this set of global variables, the base of the crank of the flexible slider crank mechanism is assumed to be rigid. Also, curvatures at all pins are set to zero, except at the base of the crank where the mechanism is driven. Considering one two-node element with quinitc polynomials for both the crank and the coupler, there are seven nodal variables: u1 (axial displacement) v1 and v2 (transverse displacements) 1 and 2 (rotations) m1 and m2 (curvatures). To construct the relationship between the nodal and global variables, several equations can be obtained by referring to Figure 3 as follows: X1 2 sin 2 3 (10) X2 2 cos 2 3 (11) X3 X1 tan 3 (12) where 2 and 3 are the angles of the crank and the coupler (see Figure 2). Also, the angle between the rigid body configuration and the reference position (see Figure 3) of the coupler can be expressed as coupler tan 1 X3 X2 R3 X3 X2 R3 (13) where R3 is the length of the coupler. Using equations (10) to (13), the relationship between the nodal of the ith element and the global variables can be expressed in matrix form: e i [Si ] (14) where e i [u1 1 1 m1 2 2 m2]T i (15) [Si ] 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 1 0 0 0 0 0 1 0 0 0 0 0 0 0 for the crank (16) [Si ] 0 sin 2 3 0 0 0 0 cos 2 3 0 0 0 0 [sin 2 3 tan 3 cos 2 3 ] R3 0 1 0 0 0 0 0 0 0 sin 2 3 tan 3 0 0 0 0 [sin 2 3 tan 3 cos 2 3 ] R3 0 0 1 0 0 0 0 0 (17) for the coupler [ 1 2 3 4 5]T column vector of the global variables (18) The above equations are based on the discretization, one element per link", + " Besides of five global variables ( i , i 1 2 5) defined earlier, six more global variables are defined: 6 (transverse displacement at middle node on the crank), 7 (rotation at middle node on the crank), 8 (curvature at middle node on the crank), 9 (transverse displacement at middle node on the coupler), 10 (rotation at middle node on the coupler), and 11 (curvature at middle node on the coupler). It is noted that all of them are with respect to the reference configuration (the dotted line for the crank and the dashed line for the coupler shown in Figure 3). Hence, the relationship between the nodal and the global variables shown as: 6 3 7 3 8 m3 for the crank (19\u201321) 9 X2 X3 2 3 10 coupler 3 11 m3 for the coupler (22\u201324) For the discretization of the number of elements greater than two, the relationship between the nodal and the global variables are similar to equations (19)\u2013(24), and then the matrix [Si ] can be obtained. Using the procedure outlined in Cleghorn (1980) the equations of the elements, equation (14), may be combined to obtain the global equations of a flexible slider-crank mechanism, which are [M] [C] [K ] F (25) where [M], [C], [K ] are global mass, damping and stiffness matrices, respectively {F} is a global load vector" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001002_1553-779x.1495-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001002_1553-779x.1495-Figure1-1.png", + "caption": "Fig. 1. Shunt and series capacitor excitation scheme of three-phase SEIG supplying single-phase load.", + "texts": [ + " The performances of the machine are computed by solving the non-linear first order differential equations of the developed dynamic model using famous fourth-order Runge-Kutta numerical technique of integration and also from the steady-state model. The simulated results have been compared with the experimental results to validate the developed model. 2 DOI: 10.2202/1553-779X.1495 Brought to you by | Purdue University Libraries Authenticated Download Date | 5/24/15 8:25 PM The circuit diagram of three-phase SEIG configured with one shunt and one series capacitor excitation is shown in Fig. 1(a). The dynamic equations for the hybrid model of the three-phase squirrel cage induction generator are given as: [vh] = [Rh][ih] + [Lh] p[ih] + \u03c9r [Gh][ih] (1) From which, the current derivative can be expressed as: p[ih] = [Lh]-1 { [vh] \u2212 [Rh][ih] \u2212 \u03c9r [Gh][ih] } (2) where, [vh], [ih], [Rh], [Lh] and [Gh] are defined in Appendix A. Also, from Fig. 1(a), the following equations can be written: )(11 sasc se L se se ii C i C pv \u2212== (3) )(1 sbsc sh sh ii C pv \u2212= (4) 0=++ scsbsa vvv (5) (6) Lsesa vvv += where, isa, isb and isc are currents in stator phases A, B and C respectively and vse, vsh and are voltages across capacitances CLv se, Csh and load respectively and p = (d/dt) is a time derivative operator. Since, the magnetization characteristics of the SEIG are non-linear due to saturation, the magnetizing inductance depends on the instantaneous value of the magnetizing current", + " In derivative form, the equation of speed can be written as: p\u03c9r = J P 2 (Tshaft \u2013 Te) (10) Hence, these eight differential equations (2), (3), (4), and (10) describe the dynamic model for the prediction of transient performance of the three-phase SEIG for single-phase power generation with shunt and series capacitor excitation scheme feeding both resistive and inductive loads. 4 DOI: 10.2202/1553-779X.1495 Brought to you by | Purdue University Libraries Authenticated Download Date | 5/24/15 8:25 PM Referring to Fig. 1(a), the following equations can also be written: (11) sascL iii \u2212= (12) sbscsh iii \u2212= (13) shsbsh Yvi = (14) seLsaL zivv \u2212= where, j\u03c9 C=shY sh. Using the symmetrical components transformation, we have the stator phase voltages as: (15) npsa vvvv ++= 0 npsb avvavv ++= 2 0 (16) npsc vaavvv 2 0 ++= (17) and the stator phase currents as: nnppsa YvYvYvi ++= 00 (18) nnppsb YavYvaYvi ++= 2 00 (19) nnppsc YvaYavYvi 2 00 ++= (20) where, \u2018a\u2019 is the unit complex operator given by 3 2\u03c0je . For the delta connected SEIG, 0 3 )( 0 = ++ = scsbsa vvv v ", + " From equations (22) and (23), we have (24) pL vkkkkkv }/){( 4324= 1 + Similarly, substituting symmetrical components for isc and isa in equation (11), we get )()( 2 nnppnnppL YvYvYvaYavi +\u2212+= (25) np vkvk 6= \u2212 5 \u2212 where, pYak )1(5 \u2212= . nYak )1( 2 6 \u2212= From equations (23) and (25), can be written as: Li (26) pL vkkkkki }/){( 4635= \u2212 4 + From equations (24) and (26), we have Z i v L L \u2212= where, 6354 3241 kkkk kkkk Z + + = (27) Hence, the single-phase SEIG using three-phase machine viewed from the load terminal can be represented by a single-phase circuit shown in Fig. 1(b), in which Xm and F are the unknown quantities. Applying Kirchoff\u2019s voltage law in Fig. 1(b), we get 0)( =+ LL iZZ where, is the load impedance. LLL jFXRZ += Under steady-state condition, can not be equal to zero and hence, Li (28) 0)( =+ LZZ This equation is solved by sequential unconstrained minimization technique (SUMT) in conjunction with Rosenbrock\u2019s method of rotating coordinates. Then, the air-gap voltage Eg is determined from the magnetization curve of the induction machine, and the current ip and the voltage vp are calculated. Subsequently, the phase voltages, phase currents, load voltage, load current and the performance of the SEIG at steady-state are determined using the equations (11) to (28)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002733_healthcom49281.2021.9398994-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002733_healthcom49281.2021.9398994-Figure5-1.png", + "caption": "Fig. 5. The 3D Helmholtz coil magnetic control system. {W} is the coordinates of the system.", + "texts": [ + " 4(b)), then A drop (about 5 \u00b5L) is dropped into the container which contains about 3 mL deionized water. Before the container is placed into the workspace, the dispersed nanoparticles are gathered together by placing a permanent magnet under the container (Fig. 4(c)). A backlight plate is placed under the transparent glass container as a light source. After the above steps, further magnetic control is prepared. A 3D Helmholtz coil system is used in our experiments to actuate the nanoparticle microrobots, as shown in Fig. 5. Controlled by a control software designed by Qt, which can send a current control signal to Sensoray S826 PCIe A/D IO card, the regulation of current input into the coil can be realized by three Maxon ESCON 70/10 motor drivers, consequently, the magnitude of the magnetic field can be changed. Thus, the approximately uniform magnetic field of 65 \u00d7 60 \u00d7 35 mm3 is generated in the workspace of 100\u00d7 68\u00d7 35 mm3, a uniform magnetic field of 20mT corresponds to a current of 8A. Furthermore, a monocular camera (PointGrey GS3-U341C6M), mounted on the top of the 3D Helmholtz coils, is used to capture the images as visual feedback for the closedloop control" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000430_ias.2005.1518684-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000430_ias.2005.1518684-Figure1-1.png", + "caption": "Figure 1. Rotor structure of the tested flux-barrier type SynRM.", + "texts": [ + " Since electrical parameter variations are expressed by continuous even functions, differentiable for the entire current and speed ranges, this state equation can be used for simulating the correct transient characteristics. As an example, steady-state characteristics of a conventional vector-controlled SynRM with constant d-axis current and transient characteristics of a maximum efficiency vector-controlled SynRM are calculated from the state equation, and are verified with experimental results measured by on-load test. II. DETERMINATION OF MOTOR PARAMETERS A. Measurement of d- and q-axes inductances Fig. 1 shows the rotor structure of the tested flux-barrier type SynRM with distributed stator winding. Table I shows the ratings and mechanical constants. IAS 2005 1754 0-7803-9208-6/05/$20.00 \u00a9 2005 IEEE Fig. 2(a) shows the test circuit used for measurement of the d- and q-axes inductances, Ld and Lq, taking cross-magnetic saturation into account. First by applying a low dc voltage to the a-b winding, the rotor is aligned to the d-axis position for ab winding. Afterward, the rotor must be locked in this position during this test" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002581_icees51510.2021.9383730-Figure9-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002581_icees51510.2021.9383730-Figure9-1.png", + "caption": "Fig 9. Higher Temperature at the Motor Rear Side", + "texts": [ + " The cylindrical housing was the initial geometry and for contours of temperature, the ANSYS simulation is analyzed for rotor, stator, shaft and other BLDC motor components. Fig.8 shows the temperature contours of the BLDC motor. The contours of temperature illustrate the flux of heat of these parts from winding, rotor, bearing and stator. The temperature is due to the flux linkage of the stator conductor and rotor permanent magnet flux linkage. During higher load condition the flux linkage will be more, the speed of the rotor increases and obviously the temperature of the BLDC motor also increases. Fig 9 shows the temperature increase on the housing and stator at the posterior side of the motor. Fig 10 explains this using the results of air flow profile. 120 2021 7th International Conference on Electrical Energy Systems (ICEES 2021) Authorized licensed use limited to: Carleton University. Downloaded on June 06,2021 at 01:21:49 UTC from IEEE Xplore. Restrictions apply. Since a casing on the outer housing is absent, the air gets mixed with the surrounding, therefore the air flow volume is minimized past the housing\u2019s rear side" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000534_s0065-2458(08)60221-1-Figure34-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000534_s0065-2458(08)60221-1-Figure34-1.png", + "caption": "FIG. 34. Principle of jet interaction amplifier.", + "texts": [ + "1 Free J e t Interaction Ampli f ier Based o n M o m e n t u m Control Free jet interaction amplifiers based on momentum changes appeal to the very elementary physical concepts. There is such an ease in explaining how amplification is achieved that this type of amplifier Is used as a model in cases where the underlying assumptions are far from being true and where additional effects are far more important. The main reason for all this is that hydrodynamics is almost not involved. The basic arrangement is shown in Fig. 34. Amplification occurs by adding momentum in the y-direction to the power jet (index p ) with the aid of a control jet (index c) . Thus the jet no longer leaves the amplifier through output terminal 1, but is displaced to impinge on receiver \"out 2.\" The changes in both flow rate and pressure available a t the output terminals may exceed the corresponding changes a t the control input a t the same time. Thus power amplification is easily achieved, and there are no difficulties in getting output pressure levels matched with input pressure levels", + " More optimistic predictions can be made for the modified version using walls and allowing high flow velocities : wall interaction type amplifiers run a t sufficiently high Mach numbers, or showing heavy cavitation-effects actually perform in the way of momentum control amplifiers. As far as \u201csaturation\u201d-effects due to the walls are neglected, their behavior is the same as that of an ideal free jet amplifier. 2.3.2 Turbulence Amplijiers The turbulence amplifier published by R. Auger [27] is one of the latest additions to the family of pure fluid elements. The arrangement is given in principle by Fig. 35. It resembles very much Figure 34, related to the free jet interaction amplifier. The underlying principle of amplification, however, differs remarkably. A laminar main jet is developed through the relatively long inlet tube. (This necessitates a Reynolds number well below 2000.) In the absence of any disturbance (no control flow) the jet remains laminar until it reaches the receiving tube, where an appreciable static pressure recovery becomes 202 DIGITAL FLUID LOGIC ELEMENTS possible. If the Reynolds number related to the main jet is maintained a t a proper value (approximately 300-1200) laminar flow can be converted into turbulent flow by a relatively weak flow signal through the control tube or wen by acoustic disturbances" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000388_iembs.2006.4397933-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000388_iembs.2006.4397933-Figure1-1.png", + "caption": "Fig. 1. Conceptual drawing of the hybrid swimming robot; robot body bonded to the propulsive element. Inset: magnified image of the propulsive element: An array of bacteria attached to a polymer micro-disk.", + "texts": [ + " Moreover, simple nutrient such as glucose is provided and ATP or ion gradients are generated by the cell. Most importantly, sensors are already present in the cell and integrated with the motor. Lastly, more complex organelles can be used hence more sophisticated motions can be produced [2], [3]. Therefore, in this research, we use flagellar motor inside the intact cell. Here, we propose a hybrid swimming microrobot which is propelled by helical flagella - only about 20 nanometers in diameter - of the bacteria attached to an inorganic robot body. Conceptual drawing of the robot is depicted in Fig. 1. The advantages of the hybrid robots include: (i) They run on a small amount of nutrient for an extended period of time (miniature and efficient), (ii) Components of the propulsive element, i.e. bacteria, self-replicate; therefore no microfabrication is required. However, there are numerous challenges associated with realization and characterization of these robots. Some of them are: (i) Repeatability and yield, B. Behkam is with Department of Mechanical Engineering, Carnegie Mellon University, Pittsburgh, PA 15213, USA behkam@cmu" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000130_2005-01-2911-Figure19-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000130_2005-01-2911-Figure19-1.png", + "caption": "Figure 19: Mobility system configuration studies", + "texts": [ + " The variety of stable geometries using equilateral triangles and squares are numerous. Using fold-out revolute actuators and powered wheels, mobility and suspension systems can be compactly stowed into triangular payload panels carried by Trigon robotic modules. The mobility systems can be applied to a variety of structures to create construction equipment, mining equipment, pressurized and non-pressurized rovers, and material handling implements. A potential configuration for modules in a mobility system (Figure 19) has been explored using Klikko geometrical construction set (Klikko n.d.), which allows for assembly of shell structures using panels similar to the Trigon system. We have shown that the kinematic functionality of Trigon edge panel manipulator mechanisms will allow modules to climb existing portions of the structure and travel to their own target destinations (Figure 20). The edge connectors allow two adjacent panels to lie flat against each other so that volumes can be collapsed for compact transport (Figure 21)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002838_012021-Figure26-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002838_012021-Figure26-1.png", + "caption": "Figure 26. Strain Distribution in Carbon Fiber", + "texts": [ + " The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.7.2. Stress Distribution The Max. And Min. Stress Distribution in Carbon Fiber is 97.378 MPa and 0.28227 MPa respectively shown in Figure 25. 3.7.3. Strain Distribution The Max. And Min. Strain Distribution in Carbon Fiber is 0.0024296 and 0.000021423 respectively shown in Figure 26. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.8. Analysing Testing Result of HSLA Steel 3.8.1. Total Deformation The Max. And Min. Total Deformation in HSLA Steel is 0.19371 mm and 0 mm respectively shown in Figure 27. The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. 3.8.2. Stress Distribution The Max. And Min. Stress Distribution in HSLA Steel is 181.96 MPa and 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002683_s11665-021-05761-w-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002683_s11665-021-05761-w-Figure2-1.png", + "caption": "Fig. 2 Specially designed and fabricated workpiece fixture with a facility to provide argon gas", + "texts": [ + " The present research work provides experimental guidance to design the workpiece fixture for providing shielding gas and its flow rate, to identify the most influencing process parameters, to understand processweld morphology, laser butt-welding procedure, and precau- Journal of Materials Engineering and Performance tions which must be taken care of during high energy density LBW process. To improve weld quality, obtained by previous researchers on laser welding of SS-316L, a particular type of fixture (Fig. 2) is designed and fabricated in the present study having facility to provide argon gas to minimize weld distortion and bead protection from environmental contamination during welding. To date, most of the reported investigations in this area are focused on SS-316L by using high power CO2 and Nd: YAG lasers only. Very few systematic investigations using fiber laser are conducted to study the individual as well as interaction effect of beam power, welding speed, and defocused position on bead features, microstructural characterization, welding defects, and mechanical properties of SS-316L weldments", + " Defocusing distance is defined as the distance of the focal spot position from the top surface of the workpiece. When the focal spot position is above the top surface of the workpiece, it is called positive defocusing (Fig. 4(a)). When the focal spot is just on the top surface of the workpiece, it is called zero defocusing distance (Fig. 4(b)). On the other hand, when the focal spot position is below the surface of the workpiece, it is called negative defocusing (Fig. 4c). The laser beam falls on the workpiece surface at an angle of 85 (Fig. 2) to protect the focusing lens as the laser beam is reflected from the workpiece surface. The diameter of the fiber core through which the laser beam is delivered is 50 lm. The specimen dimension of SS-316L plates is 120 9 100 9 3 mm3. To confirm the feasibility of the welding process on SS-316L alloy using a fiber laser, the beads on plate experiments are conducted at various process parameters under atmospheric conditions. It is found from the preliminary beads on plate experiments and available previous studies (Ref 12, 29, 41, 45) that the top and bottom surfaces of the workpiece become oxidized significantly in the molten state and it enhances the brittleness of the weldments", + " The weld bead s surface color is close to the color of the workpiece material of SS-316L, i.e., silver-white color. It indicates the excellent shielding of the weld pool from atmospheric contaminations. Under atmospheric condition, the air particles may trap inside the weld bead at high temperature, and it may try to escape from the weld pool during bead solidification. It may yield cracks inside the weld bead and reduces the weld quality. Therefore, in the present study, the workpieces are mounted on a special kind of fixture with a facility to provide shielding gas (Fig. 2) to inhibit the chances of distortion in the weldments and movement of workpieces during welding. Also, the flow rate of the shielding gas on top and bottom sides, nozzle angle, and their positions are suitably selected based on the literature survey and trial beads on plate experiments. During welding, the shielding gas (i.e., argon, 99.99% purity level) flow rate is fixed as 10 L per minute from both sides of the fixture. Also, the shielding gas is delivered both on the workpiece top surface through a co-axial nozzle as well as at 45 with the laser column surrounding it both at 5 L per minute (Fig. 2) for bead shielding from environmental con- tamination. The top side nozzle is moving with the welding nozzle. From the beads on plate trials, it is found that at a higher value of flow rate of shielding gas on the top side more violent flow of molten material takes place, which leads to the instability of the keyholes and forming porosity. Also, the disruption of molten metal on the top surface occurs at a higher value of flow rate. The excessive high flow rate could attenuate the incident laser beam" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001413_j.jsv.2008.02.056-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001413_j.jsv.2008.02.056-Figure8-1.png", + "caption": "Fig. 8. Forces on ring element with axial restraint at R2.", + "texts": [ + " (46) By proceeding in a manner similar to that used to find the equations for the H1 loading, we obtain the following equations for the ring coordinate displacements due to the T4 loading: d1T4 \u00bc T4RR2 EIz0 c\u00fe n\u20182 sin a cos a 12R nc\u2018 sin a 2R (47) d2T4 \u00bc T4RR2 EIz0 d n\u20182sin a cos a 12R nd\u2018 sin a 2R (48) d3T4 \u00bc T4RR2 EIz0 \u00fe 2T4R2\u2018 ERt3 (49) d4T4 \u00bc T4RR2 EIz0 nT4R2\u2018 sin a EIz0 \u00fe 4T4R2\u2018 ERt3 (50) dVT4 \u00bc T4RR2\u00f0R2 R1\u00de EIz0 1 n\u2018 sin a 2R . (51) We assume for each individual ring element that the applied pressure loading is uniform and acts at the ring element middle surface and that axial restraint against motion is developed at the edge of the ring associated with the radius R2 as shown in Fig. 8. For purposes of analysis, we replace the pressure loading by its components Hp applied radially and Vp applied normal thereto at the center of pressure as shown also in Fig. 8. ARTICLE IN PRESS T.A. Smith / Journal of Sound and Vibration 318 (2008) 428\u2013460444 The area of the middle surface of the ring element is given by G \u00bc p\u00f0R1 \u00fe R2\u00de ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u00f0R2 R1\u00de 2 \u00fe h2 q . (52) The loadings Hp and Vp applied at the center of pressure are given by Hp \u00bc Gp cos a 2pR3 \u00bc ph\u00f0R1 \u00fe R2\u00de 2R3 , (53a) V p \u00bc Gp sin a 2pR3 \u00bc p\u00f0R2 2 R2 1\u00de 2R3 , (53b) where Hp and Vp are the assumed equivalent uniform line loadings at the center of pressure due to the pressure loading p as shown in Fig. 8, and where R3 is given by R3 \u00bc 2\u00f0R3 2 R3 1\u00de 3\u00f0R2 2 R2 1\u00de \u00f0R1aR2\u00de, (54a) R3 \u00bc 0:5\u00f0R1 \u00fe R2\u00de \u00f0R1 \u00bc R2\u00de. (54b) The loading Vp2 developed at the edge of the ring element defined by the radius R2 in Fig. 8 is Vp2 \u00bc Gp sin a 2pR2 \u00bc p\u00f0R2 2 R2 1\u00de 2R2 . (55) The torque Tp per unit length about the centroidal axis of the cross section of the ring element due to the pressure loading p, considered positive for a clockwise rotation of the cross section as shown in Fig. 8, may be determined by consideration of the forces acting thereon in Fig. 8. ARTICLE IN PRESS T.A. Smith / Journal of Sound and Vibration 318 (2008) 428\u2013460 445 This development will be clarified if the terms HT p and V T p are defined as the radial and longitudinal loads, respectively, around the total circumference of the ring element and note that cos a \u00bc h=\u00bd\u00f0R2 R1\u00de 2 \u00fe h2 1=2, (56a) sin a \u00bc \u00f0R2 R1\u00de=\u00bd\u00f0R2 R1\u00de 2 \u00fe h2 1=2, (56b) cot a \u00bc h=\u00f0R2 R1\u00de. (56c) Thus, from Eqs. (53a) and (53b), respectively, we find HT p \u00bc 2pR3Hp \u00bc pph\u00f0R1 \u00fe R2\u00de, (57a) V T p \u00bc 2pR3V p \u00bc pp\u00f0R2 2 R2 1\u00de", + " ;N: (96b) The meridional stresses of interest for the boundary face of ring element N with axial load resisted at R02 may be found from s\u00f0N\u00dey\u00f0CC\u00de \u00bc 6TB4 t 2\u00f0N\u00de 2 L\u00f0N\u00deR \u00f0N\u00de 1 cos a\u00f0N\u00de R \u00f0N\u00de 2 t \u00f0N\u00de 2 p\u00f0N\u00de\u00bdR 2\u00f0N\u00de 2 R 2\u00f0N\u00de 1 cos a \u00f0N\u00de 2R \u00f0N\u00de 2 t \u00f0N\u00de 2 HB2 sin a\u00f0N\u00de t \u00f0N\u00de 2 , (97a) s\u00f0N\u00dey\u00f0DD\u00de \u00bc 6TB4 t 2\u00f0N\u00de 2 L\u00f0N\u00deR \u00f0N\u00de 1 cos a\u00f0N\u00de R \u00f0N\u00de 2 t \u00f0N\u00de 2 p\u00f0N\u00de\u00bdR 2\u00f0N\u00de 2 R 2\u00f0N\u00de 1 cos a \u00f0N\u00de 2R \u00f0N\u00de 2 t \u00f0N\u00de 2 HB2 sin a\u00f0N\u00de t \u00f0N\u00de 2 . (97b) It is emphasized here that the loadings L in Eqs. (95)\u2013(97) include the contributions Vp2 from preceding lower numbered ring elements as given by Eq. (55) and as shown in Fig. 8. The circumferential strain, Ax, may be determined for the points of interest on any ring from 2 \u00f0n\u00de x\u00f0AA\u00de \u00bc D\u00f0n\u00de1 D\u00f0n\u00de3 \u00bdt \u00f0n\u00de 1 =2 sin a\u00f0n\u00de R \u00f0n\u00de 1 \u00fe \u00f0t \u00f0n\u00de 1 =2\u00de cos a \u00f0n\u00de , (98a) ARTICLE IN PRESS T.A. Smith / Journal of Sound and Vibration 318 (2008) 428\u2013460456 2 \u00f0n\u00de x\u00f0BB\u00de \u00bc D\u00f0n\u00de1 \u00fe D\u00f0n\u00de3 \u00bdt \u00f0n\u00de 1 =2 sin a\u00f0n\u00de R \u00f0n\u00de 1 \u00f0t \u00f0n\u00de 1 =2\u00de cos a \u00f0n\u00de , (98b) 2 \u00f0n\u00de x\u00f0CC\u00de \u00bc D\u00f0n\u00de2 \u00fe D\u00f0n\u00de4 \u00bdt \u00f0n\u00de 2 =2 sin a\u00f0n\u00de R \u00f0n\u00de 2 \u00fe \u00f0t \u00f0n\u00de 2 =2\u00de cos a \u00f0n\u00de , (98c) 2 \u00f0n\u00de x\u00f0DD\u00de \u00bc D\u00f0n\u00de2 D\u00f0n\u00de4 \u00bdt \u00f0n\u00de 2 =2 sin a\u00f0n\u00de R \u00f0n\u00de 2 \u00f0t \u00f0n\u00de 2 =2\u00de cos a \u00f0n\u00de ; n \u00bc 1; 2; 3 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002581_icees51510.2021.9383730-Figure13-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002581_icees51510.2021.9383730-Figure13-1.png", + "caption": "Fig 13. Contours o f Temperature wih 10mm Depth Axial Fins", + "texts": [ + " Rectangular fins of 2 mm thick and 5 mm deep are introduced. This reduces the upmost temperature on the terminal windings by about 11%, as in fig. 11. The depth of the fins if further increased to decreases the windings temperature up to 17%. This is evident from fig. 12. However, in designing the fins, utmost care should be taken, as it should not comprise the rigid structure of the housing. 0,025 0,075 In the end windings there is no significant effect on the orientation of the fins on the temperature as seen in fig. 13. In this case, fins of 10 mm depth, with axial orientation were employed as contrary to the radial orientation considered previously. S.No T em perature and H eat transfer rate Fins (mm) Temperature(\u00b0C) Rate o f keat transfer(W) 1 5 mm (or) 0.005 m 59.42 \u00b0C 141.48 W 2 10 mm (or) 0.01 m 55.48 \u00b0C 204.33 W 3 20 mm (or) 0.02 m 49.13 \u00b0C 244.68 W VI. Co n c l u s i o n a n d Fu t u r e Sc o p e A. Conclusion The conclusion of this research has led to several conclusions to be made about this cooling of BLDC Permanent magnet motor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001340_10402000801926471-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001340_10402000801926471-Figure2-1.png", + "caption": "Fig. 2\u2014Inertial and bearing coordinates system.", + "texts": [ + " The first one is the local coordinates system used to describe the pressure field developed within the lubricant film, with axes (\u03b6, \u03b3, \u03be) and unit vectors ( e\u03b6 , e\u03b3 , e\u03be ), depicted in Fig. 1, together with h, the film thickness, and the velocity components (u j , vj , w j ) of a point Pj located on the outer surface of the journal, as well as the velocity components (ub, vb, wb) of a point Pb located on the inner surface of the bearing. In this study, cylindrical bearings with length L and diameter D will be considered, thus \u2212L/2 \u2264 \u03b6 \u2264 L/2 and 0 \u2264 \u03be \u2264 \u03c0 D. The system (Cb, Xb, Yb, Zb) moves with the bearing central plane and is shown in Fig. 2 together with axes (X, Y, Z), which define the inertial directions. A fourth system (Cj , Xj , Yj , Zj ) moves with the journal central plane, and it is assumed that there is no relative movement between Cj and Cb (journal and bearing geometric centers, respectively) in the axial direction. In order to describe the angular displacements of the bearing (identified by the subscript b) and the ones of the journal (identified by the subscript j) around the inertial axes, the following sequence of rotations is adopted: first, the rotations Ab, j around Z result in intermediary systems (C\u2032 b, j , X\u2032 b, j , Y\u2032 b, j , Z\u2032 b, j ); second, the rota- tions Bb, j around Y\u2032 b, j define (Cb, j , Xb, j , Yb, j , Zb, j ); and finally, the bearing and the journal rotate around their longitudinal axes Xb and Xj , respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001719_med.2008.4602276-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001719_med.2008.4602276-Figure2-1.png", + "caption": "Fig. 2. Flowfield in and out of ground-effect.", + "texts": [ + " Helicopters and other rotors when operating near the ground have significantly different dynamic characteristics from those operating away from the ground. The main difference is in the power required to sustain a given lift, which, as the rotor reaches closer to the ground, becomes proportionately more effective. Increased blade efficiency while operating in ground effect is due to two separate and distinct phenomena. First and most important, there is the reduction of the induced airflow velocity. Since the ground interrupts the airflow under the helicopter, the entire flow is altered as shown in Fig. 2. In [11] the ground plane influence is modeled by introducing a mirror image of the rotor, with equal and opposite strengths in terms of momentum, at equidistance below the ground. This ensures that there can be no vertical component of the downwash velocity on the ground\u2019s plane. The second phenomenon that affects lifting efficiency of the rotor when close to the ground is a reduction in the rotor tip vortex, also illustrated in Fig. 2. When operating in ground effect, the downward and outward airflow pattern tends to restrict vortex generation. Consequently, the outboard portion of the rotor blade becomes more efficient and overall system turbulence, caused by ingestion and recirculation of the vortex swirls, is reduced. An approximation for the overall effect on the thrust produced is given by the ratio between the thrusts Tg and T\u221e generated close and away from the ground, respectively Tg T\u221e = 1 \u2212 1 16 ( R zg )2 1 + ( V vi )2 \u22121 , (2) where R is the rotor diameter, V is the translational (typically forward) velocity, vi is the induced airflow, and zg is the distance to the ground given by zg = |[0 0 1]T IpB/(cos \u03c6B cos \u03b8B)|" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001291_amr.33-37.1011-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001291_amr.33-37.1011-Figure5-1.png", + "caption": "Fig. 5 shows the final bone forming under compressed loading condition with different bone keeping value. The compressed forces are set 100N in all cases. The results show that the bone formed as a solid pillar in all cases. There is much unwanted bone stack on it, just like strange grotesque, in all of the cases. The unwanted bone positions not only undefined during the bone forming process, but also changed before or after the convergence of the strain energy. However, the unwanted bone stacks is gradually decreased by decreasing the bone keeping volume and nearly close to fibula.", + "texts": [ + " (2) According to the stress distribution situation to calculate A and I, and then decide increasing or deleting elements and its locations. (3) According to the units increasing or deleting to decide the new model. (4) Judging whether the model\u2019s shape is convergent. If it does not convergent, returned to second step and continued to carries on the iBone optimization computation until the result is satisfacted. (a) Compression Boundary Condition (b) Compression and Bending Boundary Condition Fig.4 FEM Numerical Model for Two Different Boundary Condition Fig.5. Bone Final Formation under Compressed Load (F=100N) in Different Bone Keeping Volume Fig. 6 shows the final bone forming under compressed loading condition with different bone keeping value, and the compressed forces are set 200N in all cases. The results show that the bone formed largely changed with Fig.5, and bone shape as a pipe with hole in bone keeping volume are Vk=60%, Vk=40%, Vk=30% models. But in Vk=25% models, it is also appear the same results with Fig.5 results and show solid pillar. There is much unwanted bone stack on it too in all of the cases. Fx Fy F 1014 Advances in Fracture and Materials Behavior The unwanted bone positions not defined and go round and round along the pipe wall. The unwanted bone stacks gradually decreased by decreasing the bone keeping volume from 60% to 30% and nearly close to femur, in 25% close to fibula. Vk=60% Vk=40% Vk=30% Vk=25% Fig.6. Bone Final Formation under Compressed Load (F=200N) in Different Bone Keeping Volume Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001507_fie.2007.4418155-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001507_fie.2007.4418155-Figure2-1.png", + "caption": "FIGURE 2 EMULATED TRANSPORT AND SORTING LINE", + "texts": [ + " The simulated process shows parcels being withdrawn from the store register, being recognized at an identification unit, getting transported to the corresponding discharge station by a conveyor belt, and finally being poked from the conveyor belt to the discharge station by a pusher. Each of the real models used at the Lab had been already emulated. The Emulations of the models were created using LabVIEW\u00ae from National Instrument as the development tool, since it offers a graphical interface to the user, who can easily operate it. The emulated model of the Transport and Sorting Line is shown in Figure 2. As it can be seen, a top view was selected to emulate this system. The reason of this is that this view provides a complete visualization of the process. Also, the location of each Sensor is clearly identified. 1-4244-1084-3/07/$25.00 \u00a92007 IEEE October 10 \u2013 13, 2007, Milwaukee, WI 37th ASEE/IEEE Frontiers in Education Conference S3G-20 The real scale model of the Process Line with Machine tool and its corresponding Emulation are shown in Figures 3 and 4, respectively. At the moment, a diverse library of Third-Dimension (3D) Emulations is being developed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000088_j.jsv.2006.01.044-Figure9-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000088_j.jsv.2006.01.044-Figure9-1.png", + "caption": "Fig. 9. Decomposition of the exciting force.", + "texts": [ + " The following equation holds: a00 \u00fe b00o2 i \u00bc 2zioi, (14) where oi is the ith natural frequency of the system and xi is the damping ratio for the ith mode. Given the natural frequencies and damping ratios of two different modes, the coefficients a00 and b00can be determined by solving Eq. (14). The crankshaft load is mainly comes from the cylinder combustion and the reciprocation inertia of the piston and connecting rod. This load is transmitted through the piston and connecting rod to the crankpin. The force Fc that acts on the crankpin is shown in Fig. 9. It can be resolved into the tangential force Ft and radial force Fr. The equivalent mass of the connection rod are treated as lumped mass and attached to crankshaft body. In the rotating coordinate system, the crankshaft inertia load due to centrifugal is equal to m1ro 2, where m1 is the mass of the element, r is the distance of its center of gravity from the rotating axis of the crankshaft. As the total load {F(t)} acted on the crankshaft is periodic, it can be expanded into a series of harmonic force components by Fourier analysis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003453_s10846-021-01454-7-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003453_s10846-021-01454-7-Figure7-1.png", + "caption": "Fig. 7 Mass sliding mechanism", + "texts": [ + " Since the yaw motor is triggered to operate synchronously, the robot will start to dive gradually. After a pre-determined period, the pitch motor starts to roll back, and consequently the movable mass will be shifted back towards the center of gravity (CG) point to allow the robot to floats gradually and then reaches to the neutrally buoyant state again. Further details are discussed in next sections. The mass sliding approach is designed with a combinations of linear motion motor, 5 V stepper motor, and a sliding mass with a specified weight fixed onto linear motion shaft as shown in Fig. 7. During the sliding movement of the movable mass forwards or backward of the stepper lead screw, the proposed mass-gravity model of the whole body also goes forwards or backward, respectively. This way the robot pitches down or up; under the fins propulsion, the robot\u2019s body will be pushed downward or upward, respectively. The forces act on the sliding mass is as the following [33]: The weight of the moving mass P =Mmg (Mm and g are the movable mass and gravitational acceleration, respectively)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000360_4-431-27901-6_2-Figure2.8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000360_4-431-27901-6_2-Figure2.8-1.png", + "caption": "Fig. 2.8a-c. Mechanism of pressure generation. a wedge effect, b stretch effect, c squeeze effect", + "texts": [ + " Second, the right-hand side represents the causes of pressure generation and the three terms correspond to the following three mechanisms of pressure generation, respectively: 1. The first term represents the wedge effect: pressure generation due to the fluid being driven from the thick end to the thin end of the wedge-shaped fluid film by the surface movement. 2. The second term is the stretch effect: pressure generation due to the variation of surface velocity from place to place. 3. The third term is the squeeze effect: pressure generation due to the variation of surface gap (film thickness). The simplest examples of these three effects are shown in Fig. 2.8. Each example shows the case where the term is negative, i.e., the pressure is positive. Among these three, the wedge effect is most commonly seen, for example, in journal bearings. The squeeze effect plays an important role in the small-end bearings of crank rods and is fundamentally important also in animal joints, as pointed out by Reynolds in his paper of 1886 [3]. The stretch effect can be a problem when sliding surfaces are made of elastic materials such as rubber, but can usually be ignored" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002325_012023-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002325_012023-Figure2-1.png", + "caption": "Figure 2. The velocity and force acting on the blade element.", + "texts": [ + " Finally, we simulate and verify the self-tuning algorithm by comparing the ADRC control algorithm before and after optimization based on PSO algorithm in time and frequency domain. The quad tilt rotor with partial tilt wing in this paper is shown in figure 1, including four groups of propellers, front and rear wings, fuselage, elevator, motors, tilting mechanism, undercarriage and flight control system. Both ends of the front and rear wings are designed with a tilt nacelle. The tilt wing is connected to the nacelle and turns with the tilting of the propeller in the nacelle. The propeller is modeled according to the Goldstein vortex theory, and figure 2 shows the velocity and force acting on the blade element. R m is radius, r is the distance from hub center to any point of propeller profile, x is dimensionless value of r , is propeller solidity, /rad s is rotational speed of propeller, /V m s is inflow velocity, is inflow ratio, is blade element angle, T is blade element inflow angle at propeller tip, EV is resultant velocity, ,a t are axial and circumferential induced velocity. 4t Bw r (1) SAMDE 2020 Journal of Physics: Conference Series 1780 (2021) 012023 IOP Publishing doi:10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001825_iros.2008.4651119-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001825_iros.2008.4651119-Figure6-1.png", + "caption": "Fig. 6. TVMP: (a) when all contact points are F-points, (b) when contact point on Finger 3 closer to the base is S-point", + "texts": [ + " First, we consider Problem 1 for the case shown in Fig.5. Here, we consider to grasp an object by only Finger 1 and 3. \u03a3O is placed at [0.075 0 0.1]T . The contact position of Finger 1 is [\u22120.022 0 0.074]T . The contact positions of Finger 3 are [0.075 0 0]T and [0.171 0 0.072]T . \u03bcij is set to 0.3. RVDS is not used and REFS is set as follows (Here, we set mg = 1 [N]): Sref = {wex|wex = [0 0 \u2212mg 0 0 0]T }. (31) First, we consider the case when all contact points are Fpoints. The obtained TVMP is shown in Fig.6 (a). Denoting the components of r\u0307 by [x\u0307 y\u0307 z\u0307 \u03c6\u0307 \u03c8\u0307 \u03b8\u0307]T and letting \u03b2 = [0.697 0 \u2212 0.174 0 0.697 0]T , the TVMP is contained in the space expressed by [\u03b2 y\u0307 \u03c6\u0307 \u03b8\u0307]T . The TVMP is not full dimensional. Since the overall set can not be shown, the TVMP shown in Fig.6 (a) is the set mapped to \u03b8\u0307 = 0. From Fig.6 (a), it can be seen that the directions of generable velocities are limited for the geometrical constraints and large object velocity can be generated in the specific directions. Note that TVMM is zero but if calculating TVMM in the 4 D space, TVMM = 84.2. In the first case, the translational velocity can be generated only in y direction. Then, changing one of F-points to Spoint, we try to generate translational velocity in another direction. We consider the case when the contact point on Finger 3 closer to the base is S-point. HT s \u02c6\u0307pCs is set to [1 0 0 ]T (note that for easy to understand the direction of \u02c6\u0307pCs , HT s is added). Here, we only consider translational directions (we set RVDS is \u03c6\u0307 = \u03c8\u0307 = \u03b8\u0307 = 0). The obtained TVMP is shown in Fig.6 (b). From Fig.6 (b), it can be seen that velocity in another translational direction can be Finger2 Finger1 Finger4 Finger3 Fig. 4. Target system in numerical examples generated. The object velocity can be generated only in x\u0307 > 0 direction due to the direction of \u02c6\u0307pCs . Next, in order to see the effects of changes of configuration, friction and external force, we consider Problem 2 for the case shown in Fig.7 where the object moves in positive z direction. The contact points on Finger 1 and 3 are set to be F-point" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003132_978-981-16-0119-4_20-Figure19.7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003132_978-981-16-0119-4_20-Figure19.7-1.png", + "caption": "Fig. 19.7 Interpretation of the topology result", + "texts": [ + " There are several solver options within this tool. \u2018TOSS\u2019 is the solver option used as its optimal here for its breadth.5 The results are reasonable. 5See \u2018Page 14\u2032 of the \u2018Theory User Manual\u2019 of the software. Kindly translate online since this manual is written in German. See: https://download.z88.de/z88arion/V2/benutzerhandbuch.pdf. A topology optimization result provides the guidance for developing a frame. A frame representing \u2018material, geometry, and manufacturing\u2019 feasibility needs to be created (see Fig. 19.7). The material, for example, is a standard material normally used on bicycles (\u2018IS 3074:2005\u20196). Certain additional members are added to connect members laterally. These are positioned in a manner that they do not interfere with other aggregates in the product. For virtual verification, see Figs. 19.8, 19.9, and 19.10. The FE workbench within \u2018FreeCAD\u2019 is used. For meshing \u2018Netgen\u2019 and for solving \u2018CalculiX\u2019 are used (embedded inside the workbench). Two aspects Viz. \u2018stiffness\u2019 and \u2018propensity to yield\u2019 are assessed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002992_s13369-021-05752-y-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002992_s13369-021-05752-y-Figure3-1.png", + "caption": "Fig. 3 Models for mobile robot", + "texts": [ + " When packet losses happen, then cognitive controller A utilizes the latest control input saved in the memory 2, lets cognitive controller B and PID controller use the previous control signals saved in the memories 1 and 3, respectively. Otherwise, PID controller, cognitive controllers A and B use the current control signals. Thus, the cognitive control is to minish the effects of both the time delay and packet losses on the NCS. In this section, the state space equations of mobile robot in the NCS are depicted. The cognitive controller A, PID controller and cognitive controller B are designed, respectively. The typical mobile robot can be modeled as shown in Fig.\u00a03 [34\u201338]. Here, G and Gr show the real mobile robot and virtual mobile robot, respectively. , and , represent the linear speed, azimuth angle and angular velocity of G , respectively. The reference linear velocity, rotational angle and angular speed for Gr are depicted by r , r and r , respectively. The radius for mobile robot wheel is shown by r . Similarly, 2d is used for the distance from left wheel to right wheel of mobile robot. The mass of the mobile robot is described by M . ul and ur denote the driving inputs for left and right wheels of mobile robot, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002838_012021-Figure15-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002838_012021-Figure15-1.png", + "caption": "Figure 15. Total Deformation in S-Glass", + "texts": [ + " Stress Distribution in Kevlar 29 is 36.704 MPa and 0.26231 MPa respectively shown in Figure 13. 3.3.3. Strain Distribution The Max. and Min. Strain distribution in Kevlar 29 is 0.00054321 and 0.000005542 respectively shown in Figure 14. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.4. Analysing Testing Result of S-Glass 3.4.1. Total Deformation The Max. and Min. Total Deformation in S-Glass is 0.18974 mm and 0 mm respectively shown in Figure 15. The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. 3.4.2. Stress distribution The Max. And Min. Stress distribution in S-Glass is 59.887 MPa and 0.3646 MPa respectively shown in Figure 16. 3.4.3. Strain Distribution The Max. and Min. Strain distribution in S-Glass is 0.00069941 and 0.000005478 respectively shown in Figure 17. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003132_978-981-16-0119-4_20-Figure12.4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003132_978-981-16-0119-4_20-Figure12.4-1.png", + "caption": "Fig. 12.4 CAD model of the new water jar", + "texts": [ + " It is found that the relative importance of the strategies such as flat packing and ease of handling is exceptionally higher than the other strategies [14]. Flat packing strategy makes the shipment of jar easier and allows compact transport of products within the available area. From the scores, it can be observed that the customers and the delivery teamwere keenonmodifying the design for easy handling, to lift, carry, and to transfer without much effort and support. From the results of the QfD, the design team of the case organization developed a new alternate design for the water jars as shown in Fig. 12.4. Three handles were added to the new design. Two handles on one side and one on the lower bottom of the opposite side. The provision of handles will enable anyone to lift and carry the jars at ease. The head of the jar was made flat, with a knob to pour the water, precisely like a beak of a bird. The flat head helps to transport more units in a single trip by occupying the space available. The bottom of the jar was designed with bumps to add rigidity and durability of the jar, and this bumps also helps to lift the jar even from the lower position", + " Moreover, reduced production time coupled with faster delivery using the DfL characteristics decreased the lead time considerably gaining good customer satisfaction rate and reduced carbon foot print. The jars are majorly handled by women domestically in India have to depend upon others for lifting and changing them. But now, they find it extremely easy and comfortable to lift and changeover the jars. Thus, the reputation of the product manufactured by the company has also created an impact on its customers. Figure 12.4 shows the CAD model of the new jar developed based on the key inferences derived through QFD based on the HoQ. The case organization has determined to manufacture a water jar that should adopt the concept of design for logistics to increase the profit and efficiency of delivery. Therefore, this case study is brought to the spotlight for examination. There exists numerous literature that used QFD to address the redesigning of the product. But the research which considered logistical aspects is relatively few" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001431_12.774989-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001431_12.774989-Figure2-1.png", + "caption": "Fig. 2 Rigid wheel on deformable soil", + "texts": [ + " Each wheel is equipped with a planetary gear train and two motors in the wheel hub, one motor drive the wheel forward and the other drive the wheel to contract and extend accompanied with wheel diameter varying from 200mm to 390mm. The motor motion is controlled onboard according to the terrain in situ, when the rover is climbing a challenging obstacle or driving on slope terrain, the lower wheels can be extended to improve the mobility performance. As the rover is navigating on flat lunar regolith, the wheel is contracted and wheel tread is full, the titanium alloy wheel can be taken as rigid wheel relative to the soft lunar soil. As shown in Fig. 2 , according to Bekker\u2019s model for a rigid wheel, force balance equation can be written for the system by integrating the stress equations over the contact areas [4]: 0 0 ( cos sin )W rb p d \u03b8 \u03b8 \u03c4 \u03b8 \u03b8= +\u222b \uff081\uff09 The shear stress \u03c4 and normal stress p acting on a point along wheel rim can be defined by Bekker\u2019s model[4], the rover mass is given as input, then the angular value \u03b80 can be defined from Eq. (1) by numerical integration, the corresponding maximum wheel sinkage Z0 is 14 mm. The drawbar pull force DP is the force which is available to pull or push an additional payload until the maximum available traction is reached, DP can be expressed as: 0 0 ( cos sin )DP rb p d \u03b8 \u03c4 \u03b8 \u03b8 \u03b8= \u2212\u222b (2) The rover configuration parameters are shown in table 1, the geophysical properties of lunar soil from literatures [5, 6] are shown in table 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002360_012005-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002360_012005-Figure4-1.png", + "caption": "Figure 4. Fishtail closed turn with rectilinear reverse movement, which is not parallel to the field boundary - movement from right to left.", + "texts": [ + " The purpose of the present work is to derive analytical dependences for determining the length of fishtail turns with a rectilinear reverse move that is not parallel to the field boundary, as well as for determining the width of the headland required to perform them in an irregularly shaped field and to analyze the influence of the angle between the direction of movement and the field boundary on the length of the non-working move and the width of the headland at different direction of movement of the unit in the field. A fishtail open turn and a fishtail closed turn with a rectilinear reverse movement in different directions of their performance in a field with an irregular shape are considered. To determine the length of the turn, a geometric method is used, in which the turn is represented by straight lines and arcs of a circle of equal radius. The different types of turns are presented in Figure 1, Figure 2, Figure 3 and Figure 4. The length of the turn is defined as the sum of the lengths of its geometric elements. The width of the headland required to perform the turn is defined as the sum of the segments perpendicular to the field boundary and depending on the elements of the turn. The symbols used in the figures are as follows: \u03b1 is the angle between the direction of working move of the unit and the boundary of the field; p. A \u2013 the beginning of the turn; p. B \u2013 the end of turn; p. O1, p. O2 \u2013 the centers of the first and second curvilinear movement within the turn; \u03b21, \u03b22 \u2013 the central angles of the arcs described in the first and second curvilinear movements; R \u2013 the radius of curvilinear movement (turning radius of the machine-tractor unit); M \u2013 the width of the tractor measured from the outside of the wheels", + " when there is no rectilinear move before and after the turn in the headland. The length of this move is determined by the dependence \ud835\udc59\ud835\udc54 = (2\ud835\udc45\u2212\ud835\udc35).sin(\ud835\udefc+\ud835\udefd1) sin \ud835\udefc.sin \ud835\udefd1 + (\ud835\udc45\u22120,5\ud835\udc40).cos(\ud835\udefc+\ud835\udefd1) sin \ud835\udefc \u2212 \ud835\udc59\ud835\udc4e \u2212 (0,5\ud835\udc35 + \ud835\udc45) cot \ud835\udefc (41) This dependence can also be represented in the form \ud835\udc59\ud835\udc54 = (2\ud835\udc45\u2212\ud835\udc35) tan \ud835\udefc + (\ud835\udc45\u22120,5\ud835\udc40).cos \ud835\udefd1 tan \ud835\udefc \u2212 (\ud835\udc45 \u2212 0,5\ud835\udc35) sin \ud835\udefd1 + \ud835\udc59\ud835\udc4e (42) where sin \ud835\udefd1 = 2\ud835\udc45\u2212\ud835\udc35 [( \ud835\udc35 tan \ud835\udefc +2\ud835\udc59\ud835\udc4e) 2 +(2\ud835\udc45\u2212\ud835\udc35)2] 1 2\u2044 (43) cos \ud835\udefd1 = \ud835\udc35 tan \ud835\udefc +2\ud835\udc59\ud835\udc4e [( \ud835\udc35 tan \ud835\udefc +2\ud835\udc59\ud835\udc4e) 2 +(2\ud835\udc45\u2212\ud835\udc35)2] 1 2\u2044 (44) The way of movement is shown in Figure 4. In the headline, the unit performs a left turn, followed by a rectilinear reverse movement and again a left turn to the beginning of the next working move. ICTTE 2020 IOP Conf. Series: Materials Science and Engineering 1031 (2021) 012005 IOP Publishing doi:10.1088/1757-899X/1031/1/012005 The length of the turn is \ud835\udc59\ud835\udc47 = \ud835\udf0b\ud835\udc45 + \ud835\udc421\ud835\udc422 \u0305\u0305 \u0305\u0305 \u0305\u0305 \u0305 = \ud835\udf0b\ud835\udc45 + 2\ud835\udc45\u2212\ud835\udc35 sin \ud835\udefd2 = \ud835\udf0b\ud835\udc45 + [(2\ud835\udc45 \u2212 \ud835\udc35)2 + ( \ud835\udc35 tan \ud835\udefc \u2212 2\ud835\udc59\ud835\udc4e) 2 ] 1 2\u2044 , (45) because the central angle \u03b22 determined by the triangle O1CO2 is \ud835\udefd2 = tan\u22121 ( 2\ud835\udc45\u2212\ud835\udc35 \ud835\udc35 tan \ud835\udefc \u22122\ud835\udc59\ud835\udc4e ) (46) The length of the turn has a minimum at sin \ud835\udefd2 = 1, i.e. at \ud835\udefd2 = 90\u00b0. The turn has a minimum length at \ud835\udefc = tan\u22121 ( \ud835\udc35 2\ud835\udc59\ud835\udc4e ) , (47) as well as in the open turn performed in the same direction (Figure 2). For \ud835\udefc > tan\u22121 ( \ud835\udc35 2\ud835\udc59\ud835\udc4e ) the angle \u03b22 is determined by the dependence \ud835\udefd2 = 180 + tan\u22121 ( 2\ud835\udc45\u2212\ud835\udc35 \ud835\udc35 tan \ud835\udefc \u22122\ud835\udc59\ud835\udc4e ) (48) The width E of the headland at a small angle \u03b1 (Figure 4, (a)) is determined by the dependence ICTTE 2020 IOP Conf. Series: Materials Science and Engineering 1031 (2021) 012005 IOP Publishing doi:10.1088/1757-899X/1031/1/012005 \ud835\udc38 = 0,5(\ud835\udc40 + \ud835\udc35) cos \ud835\udefc + (\ud835\udc3b + \ud835\udc59\ud835\udc4e). sin \ud835\udefc (49) When performing the fishtail turn, the tractor goes deeper into the field. The total width of the headland, taking into account the innermost point in the field reached by the tractor, can be defined as the sum of the lengths of the following sections: \ud835\udc4e = 0,5\ud835\udc40. cos(\ud835\udefc+\ud835\udefd2) (50) \ud835\udc4f = \ud835\udc3b. sin(\ud835\udefc+\ud835\udefd2) (51) \ud835\udc50 = \ud835\udc45", + " If the rectilinear moves are not working moves, the headland with width \ud835\udc38\u2032 is cultivated. In this case their length must be determined, which is \ud835\udc59\ud835\udc54 = \ud835\udc38 \u2032\u2212\ud835\udc38 sin \ud835\udefc = \ud835\udc4e+\ud835\udc4f+\ud835\udc50+\ud835\udc51 sin \ud835\udefc = (\ud835\udc45+0,5\ud835\udc40) cos(\ud835\udefc+\ud835\udefd2)+\ud835\udc3b.sin(\ud835\udefc+\ud835\udefd2) sin \ud835\udefc + (\ud835\udc45 \u2212 0,5\ud835\udc35) cot \ud835\udefc \u2212 \ud835\udc59\ud835\udc4e (55) When the angle \u03b1 between the direction of movement and the field boundary increases the way of determining the width of the headland is changed. The boundary of the headland with width \ud835\udc38\u2032 'is determined by the front left wheel of the tractor, and the rear right wheel reaches the field boundary when reversing (Figure 4, (b)). The following segments are used to determine the widths E and \ud835\udc38\u2032. \ud835\udc4e = \u22120.5\ud835\udc40. cos(\ud835\udefc+\ud835\udefd2) (56) \ud835\udc4f = ( 2\ud835\udc45\u2212\ud835\udc35 sin \ud835\udefd2 + \ud835\udc3b) sin(\ud835\udefc+\ud835\udefd2) (57) \ud835\udc50 = \u2212(\ud835\udc45 + 0,5\ud835\udc40). cos(\ud835\udefc+\ud835\udefd2) (58) \ud835\udc51 = (\ud835\udc45 + 0,5\ud835\udc35) cos \ud835\udefc \u2212 \ud835\udc59\ud835\udc4e . sin \ud835\udefc (59) For the widths of the headlands and the length of the rectilinear move, it is obtained, respectively \ud835\udc38 = \ud835\udc50 + \ud835\udc51 = (\ud835\udc45 + 0,5\ud835\udc35) cos \ud835\udefc \u2212 \ud835\udc59\ud835\udc4e . sin \ud835\udefc \u2212 (\ud835\udc45 + 0,5\ud835\udc40) cos(\ud835\udefc+\ud835\udefd2) (60) \ud835\udc38\u2032 = 2\ud835\udc4e + \ud835\udc4f = \u2212\ud835\udc40. cos(\ud835\udefc+\ud835\udefd2) + ( 2\ud835\udc45\u2212\ud835\udc35 sin \ud835\udefd2 + \ud835\udc3b) sin(\ud835\udefc+\ud835\udefd2) = = 2\ud835\udc45 cos \ud835\udefc \u2212 2\ud835\udc59\ud835\udc4e sin \ud835\udefc \u2212 \ud835\udc40. cos(\ud835\udefc+\ud835\udefd2) + \ud835\udc3b. sin(\ud835\udefc+\ud835\udefd2) (61) \ud835\udc59\ud835\udc54 = \ud835\udc38 \u2032\u2212\ud835\udc38 sin \ud835\udefc = (\ud835\udc45\u22120,5\ud835\udc40) cos(\ud835\udefc+\ud835\udefd2)+\ud835\udc3b.sin(\ud835\udefc+\ud835\udefd2) sin \ud835\udefc + (\ud835\udc45 \u2212 0,5\ud835\udc35) cot \ud835\udefc \u2212 \ud835\udc59\ud835\udc4e (62) The length of the rectilinear move is \ud835\udc59\ud835\udc54 = 0, (i.e. \ud835\udc38\u2032 = \ud835\udc38) at an angle ICTTE 2020 IOP Conf. Series: Materials Science and Engineering 1031 (2021) 012005 IOP Publishing doi:10.1088/1757-899X/1031/1/012005 \ud835\udefc = tan\u22121 ( (\ud835\udc45\u22120,5\ud835\udc40) cos \ud835\udefd2+\ud835\udc3b.sin \ud835\udefd2\u2212\ud835\udc45\u22120,5\ud835\udc35 (\ud835\udc45\u22120,5\ud835\udc40) sin \ud835\udefd2+2\ud835\udc59\ud835\udc4e\u2212\ud835\udc3b.cos \ud835\udefd2 ) (63) At a higher value of the angle \u03b1 the tractor moves within the headland with width E, which is determined by dependence (60) (Figure 4, (c)). The reverse move of the tractor is parallel to the field boundary when \ud835\udefc+\ud835\udefd2 = 180\u00b0. When the angle \u03b1 increases above the value \ud835\udefc = tan\u22121 ( \ud835\udc45 \ud835\udc59\ud835\udc4e ) (64) the headland is limited by the front right wheel of the tractor (Figure 4, (d)). The width of the headland is determined by the dependencies: \ud835\udc4e = \u22120.5\ud835\udc40. cos(\ud835\udefc+\ud835\udefd2) (56) \ud835\udc4f = \u2212\ud835\udc3b. sin(\ud835\udefc+\ud835\udefd2) (65) \ud835\udc50 = \u2212\ud835\udc45. cos(\ud835\udefc+\ud835\udefd2) (66) \ud835\udc51 = (0,5\ud835\udc35 \u2212 \ud835\udc45). cos \ud835\udefc + \ud835\udc59\ud835\udc4e . sin \ud835\udc59\ud835\udc4e (67) \ud835\udc38 = \ud835\udc4e + \ud835\udc4f + \ud835\udc50 + \ud835\udc51 = = (0,5\ud835\udc35 \u2212 \ud835\udc45) cos \ud835\udefc + \ud835\udc59\ud835\udc4e . sin \ud835\udefc \u2212 (0,5\ud835\udc40 + \ud835\udc45) cos(\ud835\udefc+\ud835\udefd2) \u2212 \ud835\udc3b. sin(\ud835\udefc+\ud835\udefd2) (68) When the rectilinear reverse move is not parallel to the field boundary, there is no rectilinear nonworking move when entering and exiting the headland with width E, i.e. when the unit may goes outside from the headland when turning" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001076_detc2007-34665-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001076_detc2007-34665-Figure2-1.png", + "caption": "Figure 2. Non-rotating forced excitation test set-up", + "texts": [ + "org/about-asme/terms-of-use Forced excitation tests are widely used for characterizing the stiffness and damping of structures and mechanical Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/01/2016 components. Several past researchers [11-13] have used more sophisticated versions of the frequency domain based method presented here. For linear systems tests using forced response can be performed with an impulse, a periodic chirp, or even a random noise excitation. However when considering systems that exhibit strong nonlinear behaviors, test parameters require more careful control. The experimental test rig is shown in Figure 2. Originally developed for testing high-speed oil-free bearings [10], the rotating test rig in Figure 2 was used to support a soft mounted rotor system using wire mesh dampers. There are two sets of duplex mounted ceramic angular contact ball bearings supporting the rotor. Each bearing set is mounted in series with double-fold beam style squirrel cage, which is in parallel with the wire mesh damper donuts. Damping or energy dissipation is generated when there exists a relative motion between the inner and outer squirrel cage rims. Each copper damper was designed to have a 25% density mesh, a 4.8 in (122 mm) OD, a 2", + " The differences in the methods can be attributed to the different mode shapes for the two experimental setups in Figures 2 compared to Figure 6. Since the damping is frequency dependent, a direct correlation for damping cannot be made between the two methods, however interpolating the FRM results in very close agreement with the transient time domain method. 7 Copyright \u00a9 2007 by ASME fication results rms of Use: http://www.asme.org/about-asme/terms-of-use Download The rotating tests were conducted using a longer rotor than shown in Figure 2. The rotor model of the long rotor is shown in Figure 10. The imbalance was determined through a balance machine check and was used in the appropriate locations and phases. The synchronous response to imbalance (Figure 11) was calculated for station 8 (motion probe 3) using the damping and stiffness results from non-rotating testing and shows two well-damped critical speed locations at 11,000 rpm 8 ed From: http://proceedings.asmedigitalcollection.asme.org/ on 02/01/2016 Terms and 24,000 rpm. Since the stiffness and damping coefficients are functions of both frequency and amplitude, the synchronous response calculation required iteration between the vibration amplitude and force coefficient at each rotor speed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002772_978-3-030-67750-3_14-Figure28.3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002772_978-3-030-67750-3_14-Figure28.3-1.png", + "caption": "Fig. 28.3 Nozzle exterior design improvement", + "texts": [ + " The aim is to create a high pressure and velocity stream of the fine particles flow out through the outlet, thus creating a narrow spray distance. As for mode two, as presented in Fig. 28.2. Figure 28.2(b), the dry powder particles for this mode flow two direction through outlet 1 and outlet 2. For the current nozzle, outlet 2 has a bigger diameter with 52 mm. The working principle of these outlets requires the nozzle to turn counter clockwise to allow the particles travel through two pathways. This is to allow the fluid to pass through a more significant diameter outlet (outlet 2) that create a wider spray angle with a slow velocity. Figure 28.3 illustrate the evolution of the bi-nozzle design to the existing one. The design begins with the setting up the project in SolidWorks. In the study on the nozzle, an internal analysis type was selected to simulate an internal flow. A list of pre-defined fluid already exists in SolidWorks but ammonium dihydrogen phosphate (Senthilkumaran et al. 2012) was not included in the library. This needs to be added manually. Thewall was assumed to be adiabatic. The roughness of the internal nozzle is also assumed to be zero" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002026_ettandgrs.2008.291-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002026_ettandgrs.2008.291-Figure1-1.png", + "caption": "Figure 1. The anchor system", + "texts": [ + " Another linearization method is the inverse system approach[48], which use the concept of mapping, and is easy to master by engineers, Figure 3. In this study, the inverse system approach is used to linearize the windlass motor, and linear controllers are designed to fulfill constant tension control of the anchor chain under see wind. The tension of the anchor chain, Hchain, is described by: Hchain = Hwindlass + Hwind, (1) where Hwindlass is the force from the windlass, and Hwind is the force generated by sea wind, Figure 1 & 2. The wind acts on the ship, and the ship then acts on the anchor chain through the windlass system. The constant tension control framework is shown in Figure 2. H*chain, the desired tension of the chain, is compared with the real tension, Hchain. The error signal, H*chain - Hchain, is sent to the controller. The controller generates the input signal of the windlass, and Hwindlass is generated to counteract the wind disturbance, Hwind. 978-0-7695-3563-0/08 $25.00 \u00a9 2008 IEEE DOI 10.1109/ETTandGRS" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002408_tmag.2021.3064023-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002408_tmag.2021.3064023-Figure3-1.png", + "caption": "Fig. 3. 3-D diagram of the half-quasi-squirrel cage winding.", + "texts": [ + " Toroidal winding shown in Fig. 2(b) in which the coil is wound around stator yoke is proposed to shorten the length of the useless end, but the utilization of the coil is very low [8]. As for the concentrated winding shown in Fig. 2(c), although the end length is shorter, a large eddy current loss will be caused as this winding type has more harmonic components. To deal with the mentioned problems, in this article, a novel winding structure, named half-quasi-squirrel cage winding, is proposed. As shown in Fig. 3, the main features of the proposed structure are as follows: one conductor is inserted into each stator slot to improve the utilization rate of slot, and connected by the end ring at one end. The other end is directly connected to the modular driving circuit. In this way, a high integration of machine and controller can be realized, which can further reduce the system volume too. At the same time, to reduce the copper loss of the conductor, the litz wire having many thin wires in parallel is used as the conductor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000035_cdc.2004.1430366-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000035_cdc.2004.1430366-Figure1-1.png", + "caption": "Figure 1: Engagement geometry - Collision triangle", + "texts": [ + " It is shown in this paper that uniform ultimate boundedness of the missile-target system state (LOS and LOS rate) is obtained in the case of highly maneuvering targets provided certain conditions are met. For the case of a nonmaneuvering target, asymptotic stability is demonstrated. The three-dimensional engagement of a missile for the intercept of a target is relatively complex. A more tractable two-dimensional engagement can be studied by assuming that the lateral and longitudinal planes are decoupled, as can typically be achieved by means of roll control [13]. The two-dimensional missile-target engagement geometry is shown in Figure 1, where vm and am are missile speed and acceleration, respectively, and vt and at are the target speed and acceleration, respectively. The range r between the missile and the target is related to the closing velocity vcl as vcl = dr dt (1) and the sine of the LOS angle (t), which is the angle between the LOS and the fixed reference, is given by y(t), which is the relative separation between the missile and the target perpendicular to the fixed reference x-axis, over the range r(t) : sin( (t)) = y(t) r(t) ", + " In general, a reasonable assumption is to define the smallest miss distance rm to be equal to half of the largest target dimension [16]. The second and third inequalities in (6) relate to limitations in velocities and accelerations of both the missile and the target. For the missile, maneuverability is constrained by its aerodynamics, for instance the control deflections are constrained. The objective of the guidance law is to drive the missile to the so-called intercept or collision triangle [1], as shown in Figure 1, where the following collision condition is satisfied vtsin( t) = vmsin( m). (7) Equation (7) comes from the LOS rate equation given by r \u2022 = vtsin( t) vmsin( m) (8) which expresses target velocity with respect to missile velocity orthogonally projected unto the LOS. Equation (7) corresponds to the equilibrium d /dt = 0 for system (4). The guidance command ag is synthesized such that (3) and (4) are asymptotically stable at equilibrium ( = o, \u2022 = 0). Figure 2 shows the feedback structure for the guidance system, which includes the dynamics of a maneuvering target, the missile flight control system (missile in closedloop with autopilot) and an estimator" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002687_j.seta.2021.101240-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002687_j.seta.2021.101240-Figure6-1.png", + "caption": "Fig. 6. Flux lines Trends of conventional prototype Two Poles-PMDC Motor by Infolytica MotorSolve Software.", + "texts": [ + " Furthermore, the projected ABO technique effective performance is compared with the conventional techniques namely Finite Element Analysis (FEA), Slot and Pole Configuration (SPC) model, and Analytical Model (AM) for Cogging Torque computation. Flux lines of conventional PMDC motor provide high armature current at no load torque, flux lines are clearly indicating their trend in comparison with R. Ullah Khan et al. Sustainable Energy Technologies and Assessments 46 (2021) 101240 the conventional PMDC Motor is shown in Fig. 6. The modified PMDC motor has small air gap between the stator and rotor position. Moreover, this modified PMDC motor net magneto force will be reduced. The modified pole PMDC motor is designed with Infolytica MotorSolve Software is illustrated in Fig. 7. Flux lines Trends of Modified Pole PMDC Motor has been shown with the help of \u2018Infolytica MotorSolve Software. After cutting the pole magnets into pieces, it has been observed that, the concentration of flux lines are more towards the gap in magnets of main poles i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002859_s12046-020-01541-9-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002859_s12046-020-01541-9-Figure4-1.png", + "caption": "Figure 4. Cooling curve comparison of SGI with varying section thickness.", + "texts": [ + "3) SGI is preferred since it has a lower tendency towards shrinkage defects. As given in table 1, the CE (%C?1/3%Si) value of around 4.5 was used for simulation and for melting as well. The solidification cooling curves of the different sections of a stepbar were obtained by simulating SGI of varying section thickness. The cooling curves of section thickness 5 mm are represented as point 1, and the cooling curves of sections 10 and 15 mm are at points 2 and 3, respectively. The respective cooling curve of different section thicknesses is given in figure 4. The liquidus temperature of this composition was found to be 1167 C, whereas the solidus temperature was 1163 C. It can be observed from figure 4 that as the section of casting increases the eutectic thermal arrest of SGI increases. The eutectic arrest time in the 5-mm section cooling curve is about 50 s, whereas for 10 and 15 mm it is around 70 and 120 s, respectively. This variation in the eutectic arrest with section thickness can be explained by the fact that the precipitation of graphite during eutectic reaction leads to the release of energy due to latent heat, which balances the heat losses of the melt. Therefore the higher the extent of graphite precipitation higher will be the heat released, which leads to longer thermal arrest during the eutectic reaction. Consequently, the time of eutectic arrest is an indication of the amount of graphite formation at the eutectic reaction. It can be observed from figure 5 that the cooling curves at three different points on the same section thickness do not show any legitimate variation in the cooling curve pattern. The thermal arrest is similar in each cooling curve given in figure 5. It can also be observed in figure 4 that the degree of undercooling is higher in the thin section, whereas the 15 mm section shows almost no undercooling with a comparatively flatter cooling curve as against those of 5 and 10 mm thickness. It is generally accepted that undercooling and grain size are inversely related to each other; therefore it can be said that the grain size in the thin section will be smaller; consequently, the number of grains in thin section sections will be more compared with thick sections of the casting. Thus, it is important to note here that the section thickness of castings will have a substantial impact on the final microstructure and amount of graphite" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000676_978-1-4020-8829-2_9-Figure11-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000676_978-1-4020-8829-2_9-Figure11-1.png", + "caption": "Fig. 11. Initial and optimal designs for the isotropy maximization of the Hunt platform", + "texts": [ + " (34) The objective function f of this 6-dimension problem is finally: f (zc, LI, LS,RB,RP,H, \u03b1, \u03b2, \u03c8, LC) = 1 32 32\u2211 i=1 1 \u03ba (( J\u0303f ) i ) . (35) Also for practical reasons, design parameters of this robot are limited by bounds (see minimum and maximum values in Table 3). The same restrictions are imposed on the assembly problem to avoid \u201cinside\u201d configurations as in Fig. 7. This leads to the final extended function g (see (8a)). The results obtained with the SQP method are presented in Table 3, and initial and optimal design can be compared in Fig. 11. The graphical results are represented in Fig. 12. It takes 94 iterations and 110 evaluations of the objective function to optimize the 10 parameters. In Fig. 5.2, we can observe the evolution of the objective function in three main steps. These steps correspond to \u201csequences\u201d of the SQP method for which active constraints are defined. For example, we can observe in Fig. 5.2 that the active bound constraint on variable \u03b2 is relaxed in iteration 22 from which \u03b2 begins to increase. L en g th s (m m ) The \u201cvalidation\u201d of this result is made by applying stochastic optimization algorithms on the same optimization problem: on the one hand, we apply a classical genetic algorithm and on the other hand, an evolutionary strategy is used" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002128_s13369-020-05100-6-Figure10-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002128_s13369-020-05100-6-Figure10-1.png", + "caption": "Fig. 10 Demonstrates the safety factor", + "texts": [ + " In addition to the deformation, the concentrated stresses that were exposed to the machine were calculated in more than one location to know the stresses on the shafts and screws due to the load resulting from the welding torch and the gantry. The results showed that the greatest stresses were placed on the screws as the responsibility for the motion in addition to the shafts as shown in Fig. 9d\u2013f. Moreover, the safety factor was calculated for the machine designed when exposed to a load higher than the load used and was within the save side aspect of the user\u2019s design as shown in Fig. 10. The CADWDM system includes the development of a 3D object depended on the CAD drawing, planning and slicing of patterns in standard thickness layers. The file that is drawn by CAD is imported or brought to the online programing technique. Therefore, the desired geometry to be classified according to the CAD specification is sliced into layers with a regular height. The deposition process was created using an interface between the deposition tool and modeling data. After validating the model, it is sliced from bottom to top with regular layer thickness" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002874_s10846-021-01410-5-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002874_s10846-021-01410-5-Figure8-1.png", + "caption": "Fig. 8 a visualisation of various coordinate frames associated with the proposed benchmark system. b illustration of the function of the conveyor system. Object are placed haphazardly at one end of the prototype, where they may be moved into vicinity of the desired coordinate frame by the conveyor. Red, green, and blue coordinate lines express the x, y, and zaxes, respectively", + "texts": [ + " Initially, objects placed on the conveyor belt are translated to an approximate x-axis mid position within top-view image space by utilising the beam sensor. At this stage, objects are discoverable through vision and grasp configurations may be computed by the deployed grasp synthesis methodology. Top-view image space It and robot workspace G may be treated as though they occupy the same space by translating the object a set amount between frames. Effectively, conveyor space is ignored and treated as an intermediary consolidation step prior to an interaction. Figure 8b illustrates the function of the conveyor system. Note that the intermediate conveyor translate step introduces a fixed error, shown in Table 5. Each load-cell unit was calibrated individually. Note that a load-cell unit consists of an HX711 amplifier and TAL220 load-cell pairing. Calibration was specific to each unit and varied slightly between units. 50 initial measurements were taken with no weight present to tare the unit. A calibration tool was bolted to the load-cell and used for calibration", + " Errors associated with this subsystem are tabulated in Table 6. Because an object is supported by a fixed number of loadcells\u2014and the total weight is known\u2014the ratio of weight distribution between axes can be calculated. This ratio may be used in conjunction with known distances between loadcells to compute a relative COG position within vision. The load-cell coordinate frame LC is quantified by 4 sensitive components\u2014the conveyor platform rests on these components. Load-cell space LC and the top-view camera frame It share two axes\u2014Fig. 8a and Fig. 8. To formalise load-cell coordinate space in terms of vision x- and y-axes, the output of each component is measured. Ratio coefficients are used to estimate the COG position of an object xtCOG\u00bd ; ytCOG in image space It directly: xtCOG \u00bc acoeffx\u2212b \u00f05\u00de ytCOG \u00bc ccoeffy\u2212d \u00f06\u00de where a, b, c and d are found through experimentation. Ratio coefficients are defined by: coeffx \u00bc weightRF \u00fe weightRR weighttotal \u00f07\u00de coeffy \u00bc weightRF \u00fe weightLF weighttotal \u00f08\u00de where weighttotal is computed as the sum of all weight measurement inputs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000469_50009-6-Figure8.7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000469_50009-6-Figure8.7-1.png", + "caption": "Figure 8.7. Lateral view of drag components encountered by a swimming sea lion.", + "texts": [ + " Both of these characters are lacking in phocids (King, 1983). 8.2.5.3. Mechanics of Locomotion Among modern pinnipeds, terrestrial and aquatic locomotion are achieved differently. Three distinct patterns of pinniped swimming are recognized, yet all create thrust with the hydrofoil surfaces of their flippers. When swimming, these hydrofoils are oriented at an angle (the angle of attack) to their direction of travel, producing thrust parallel to the direction of travel and generating lift perpendicular to that direction (Figure 8.7). One of these patterns, pectoral oscillation (forelimb swimming), is seen in otariids. Sea lions and fur seals move their forelimbs to produce thrust in a manner similar to flapping birds in flight. Observations indicate that the hind limbs are essential in providing maneuverability and directional control but play little role in propulsion (Godfrey, 1985). The larger pectoral flippers, with nearly twice the surface area of the pelvic flippers move in unison, acting as oscillatory hydrofoils in a stroke that includes power, paddle, and recovery phases (Figure 8" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002655_iaeac50856.2021.9390842-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002655_iaeac50856.2021.9390842-Figure1-1.png", + "caption": "Fig. 1. Phasor diagrams of five-phase induction motor\u2019s magnetic linkage and EMF", + "texts": [ + " HARMONIC CURRENT CONTROL METHOD OF FIVEPHASE INDUCTION MOTOR BASED ON ASYNCHRONOUS ROTATION TRANSFORMATION By dual-plane asynchronous rotation transformation, project the fundamental and third harmonic subspace variables to the rotary coordinate systems d1-q1, d3-q3, where \u03c9e1 is not equal to \u03c9e3, and orientate the fundamental and third harmonic rotor flux linkage to the d1, d3 axis respectively, and then adopt the dual-plane magnetic field orientation control for the fundamental and third harmonic rotor flux linkages \uff0c namely, r j rjd \uff0c r j 0q . Further, the phasor diagrams of five-phase induction motor under dual-plane magnetic field orientation control with the third harmonic injection are obtained, as shown in Fig. 1. 1086 Authorized licensed use limited to: University of Prince Edward Island. Downloaded on July 03,2021 at 12:09:49 UTC from IEEE Xplore. Restrictions apply. In Fig.1, 2 2 mj m j m j mj m j m j m jarctan( / ) d q d q d (4) According to Fig. 2, the expression of the fundamental and third harmonic air gap flux linkage angle can be obtained: rj s j mj rj s j arctan q d l i L i (5) Substitute (5) into (3), the following expression between the fundamental and third harmonic current components can be following: r3 s 3 r1 s 1 e1 e3 r3 s 3 r1 s 1 arctan 3( ) 3arctan q q d d l i l i L i L i (6) Assume that s 3 r1 s 1r3 I3 e1 e3 s 3 r3 r1 s 1 tan[3( ) 3arctan ] q q d d i l iL k i l L i (7) In combination with (4), (6), and (7), the expressions of the third harmonic components can be obtained as follows: e1 m1 s 3 2r3 I3 e3 m3 r3 I3 e1 m1 s 3 2r3 I3 e3 m3 r3 18 1 ( ) 18 1 ( ) d q i l k L L k i l k L L (8) Since the electromagnetic torque generated by the third harmonic current is very small, the electromagnetic torque can be approximately expressed as follows: 2 m1 em t p s 1 s 1 r1 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003537_s40430-021-03132-z-Figure15-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003537_s40430-021-03132-z-Figure15-1.png", + "caption": "Fig. 15 A phase of a chamber in GEROTOR unit", + "texts": [ + " Appendix: 2 Geometric volume displacement of\u00a0GEROTOR Unlike the cylinder piston machines where volume displacement rate is proportional to the rate of change of stroke (10)\ufffd\ufffd\ufffd\ufffd\ufffd\ufffd\u20d7OM = \ufffd\ufffd\u20d7Co \ufffd\ufffd\ufffd\ufffd\u20d7+Ao + r\u20d7m (11) R0 r0 = Z (Z \u2212 1) i.e., R0( R0 \u2212 r0 ) = Z (12) Xm =Ao cos + 1 Z cos Z + rm cos ( + ) Ym =Ao sin + 1 Z sin Z + rm sin ( + ) (13) = tan\u22121 ( sin (Z \u2212 1) A0 + cos (Z \u2212 1) ) length per unit time, volume displacement rate in epitrochoid generated ROPIMAs depends upon the time rate of change of area of the chambers in azimuthal direction (Fig.\u00a015). Using equation (14), it is also possible to find the volume variation in working chambers in time due to instantaneous flow rate of the pump. For calculation of the area of the GEROTOR pump working chamber, the method that considers the analysis of influence of infinitesimal values of rotation angles of the pump\u2019s working elements is used. The area of working chamber is calculated by integration. For calculation of the variation in the pump\u2019s working chamber, the equivalent system with fixed axis of gear pair elements is considered" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000602_978-3-540-74764-2_52-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000602_978-3-540-74764-2_52-Figure1-1.png", + "caption": "Fig. 1. Different modules of the final robot.", + "texts": [], + "surrounding_texts": [ + "navigation, planning, simulation and user interfacing. For the integrated onboard vision system, a camera, which can be mounted on mobile micro-robots, has been developed into a working prototype. This system can generate 3D data of micro manipulation scenarios. The computer vision system developed offers a broad range of stable recognition algorithms for micro handling applications. In the field of object localisation, the global localisation system represents a major achievement. This system has reached final prototype state with a position resolution of about 4\u00b5m over the complete size of the workspace. By using phase shifting algorithms the resolution of the system is about 1\u00b5m.\nThe platform consists of a locomotion system allowing a velocity of about 0.5 millimetres per second comA\u0302bined with a large working area (220 \u00d7 200 mm2) and nanometer resolution. The robots navigate on a flat, horizontal surface using holonomic movements. The main tasks of the locomotion system are to bring the handling-tools (e.g. grippers) to the region of interest or to transport micro objects. The hardware platform incorporates advanced, beyond state-of-the-art technological solutions. The system design enhances the performance achievable by each individual micro-robot by suitably distributing various mechanical degrees of freedom (DOF) among the micro-robots to enable complex object handling and manipulation. Piezo actuators have been incorporated due to their excellent resolution. An onboard IR module assures the communication with a host computer. The onboard electronics controls the robot motion, generates and amplifies the driving signals for all actuators and tools and pre-processes the signals from the on-board sensors. Power management was a critical issue to ensure autonomous operation of the robot for the longest time possible. A power consumption of significantly less than 1 W per robot was aspired. In particular, heat management was important from both an energy saving point of view and to avoid detrimental effects resulting from overheated system components or manipulated objects.\nThe wireless power transfer is implemented by an external magnetic field generator, the so-called Power Floor. It provides a travelling magnetic field throughout the working area, powering the robot via a miniaturised robot coil. The operating frequency is defined as 500 kHz. The coil power pack with resonant, rectifying and filtering circuit is designed in a volume of 11, 5 \u00d7 11, 5 \u00d7 4mm3. In order to reduce the skin effect, the coil is wound with 200 turns of Litz-wire, which is composed of 30 twisted \u00d8 0,03 mm enamel copper wires. The output voltage depends directly on the intensity of the external magnetic field, which is tuned to transfer a power output of 330 mW / 3.3 V from the coil power pack.", + "The local position sensor is based on an AFM scanner mounted on a rotor high motion positioning actuator. The AFM sensor consists of one position scanner and one cantilever. The scanner is made of 4 PZT stack actuators that permits movements with 3 DOF (x,y,z) and the cantilever contains a force sensor based on a piezoresistance (see Fig. 2).\nThe AFM tool consists of three main components: 1.) the AFM probe with the integrated piezoresistance, 2.) an AFM holder for easy probe exchange and 3.) the XYZ scan stage.\nA rotational drive has been developed, evaluated and redesigned to be a suitable interface between robot and tool. Five rotational drives with specific rotors", + "for each tool have been built (see Fig. 3). Furthermore, the rapid prototyping technique for making multilayer piezoceramics, developed in the beginning of the project, has been utilised to make the drivers for the grippers.\nThe maximum torque of the motor is 80\u00b5N at a drive voltage of 50 V and a spring force of 1.2 N. The rotational actuator has extremely good motion resolution (0, 1\u00b5rad) for driving frequencies up to about 80 Hz (0.1 rpm). Fast transport can be achieved in the frequency range 3\u20136 kHz (4 rpm). The power consumption is about 1 mW and 80 mW respectively.\nA micro gripper was developed which consisting of a piezo-electric actuator with gripper arms of stainless steel machined by Electro Discharge Machining (EDM). Although in this work, U-shaped actuators were used for the steel grippers, in theory bar shaped actuators could also be used. The stainless steel gripper tips are machined using wire EDM. The total gripper length is approximately 12 mm and the tip thickness of the gripper is approximately 60\u00b5m.\nOne project goal was the development of a position sensing system with a resolution of about 5 \u00b5m. This has been successfully realised and works now in a stable prototype status. A partly-virtual Moire\u0301-based interferometer was developed to reach this demand. This measurement tool is needed to get the precise positions of the micro-robots. The software communicates via TCP/IP with the robot control software and sends the coordinates of the Moire\u0301-marks.\nOne test setup for the measurement system in a real environment that contains optical obstacles is shown in Figure 4 (left). There are several objects in this image: a.) micro-robots with equipped Moire\u0301-marks, b.) chessboard-like distortions on the working floor, and c.) other black objects which are not a aprt of a micro-robot. These objects have to be distinguished from the Moire\u0301-marks, so that a successful measurement cycle is guaranteed (see result in Fig. 4, right).\nA good practice for increasing the measurement quality is the usage of phase shifting. A common phase shifting method using four 90\u25e6 phase shifts has been used. The result is an error surface, which shows a statistical measurement error" + ] + }, + { + "image_filename": "designv11_83_0001320_msf.580-582.85-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001320_msf.580-582.85-Figure4-1.png", + "caption": "Fig. 4 The jig for compressive test", + "texts": [ + " Impact and absorption energy were calculated. The impact energies were 30J, 50J, 70J and 90J. The jig was manufactured in-house because the specimens were too large to be fixed with the existing jig. Four sides of the specimen were fixed. Fig.3 The specimen of impact test Compressive test of the honeycomb sandwich panel Specimen of 240 x 120mm size were constructed according to ASTM C-364 and the test was conducted with INSTRON 8502. The velocity of head was set up for 5mm/min. A jig with a guide was used to prevent its collapse. Fig.4 shows the jig, manufactured for testing. To measure the highest compressive strength, the load in the edgewise direction, was added until breaking point. Result and discussion Evaluation of the mechanical properties of face material Because of increasing fiber\u2019s number by 12K, the faces, CF1263/Epoxy were stronger than CF3327/Epoxy having 3K(tensile stiffness 5.2%, tensile strength 45%, bending stiffness 7.2%, bending strength 0.5%, inter planular strength 6.7%). Low velocity impact test of honeycomb sandwich panel When the impact energy was 30J, there was no failure on the sandwich panel" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002953_tia.2021.3089662-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002953_tia.2021.3089662-Figure3-1.png", + "caption": "Fig. 3: In (a), scheme of the test bench with the Motor Under Test (MUT), the Master Motor (MM), and the Torque Meter (TM). In (b), picture of the test bench.", + "texts": [ + " fr = 0 Hz, the current supplied to the motor is only the magnetizing current, thus the rotor current is zero and the stator current vector is aligned to d\u03bb-axis. It should be noted that in this situation the motor is not producing torque and the q\u03bb-axis current is zero. From this condition, increasing the rotor frequency, i.e. reducing \u03c9me, the stator current angle is modified moving from the d\u03bb-axis toward the q\u03bb-axis. At the end, when the rotor is locked \u03c9me = 0 rad/ sec, i.e. fr = fs the current vector is approximately aligned to q\u03bb-axis. The test bench scheme adopted for this measurement is reported in Fig. 3a, including the electrical machines involved, namely induction Machine Under Test (MUT) and Master Authorized licensed use limited to: University of Liverpool. Downloaded on July 04,2021 at 08:30:43 UTC from IEEE Xplore. Restrictions apply. 0093-9994 (c) 2021 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. Motor (MM). Fig. 3b shows the test bench picture. As shown in Fig. 2, the purpose is to modify the stator current angle of MUT by varying fr for a given stator current amplitude. MUT is operated by the control system at fixed reference angular frequency \u03c9e,ref and reference current amplitude iref. MM is controlled by the mechanical speed reference \u03c9m,ref. This allows a precise rotor frequency control of MUT. It is worth to notice that the torque is not controlled, the torque (Tm) that is produced at the different rotor frequencies is measured by a torque meter installed in the system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002738_s0263574721000539-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002738_s0263574721000539-Figure2-1.png", + "caption": "Figure 2. Experimental environment for simulating planar robot motion in microgravity on ground: (a) overview, (b) air-floating testbed (from two viewpoints), (c) eccentric motor unit, and (d) elastic body.", + "texts": [ + " In contrast, on rough or rocky terrain, the robot utilizes the rotational hopping mode by taking advantage of the reaction torque. Thus, the robot enables more adaptive mobility to unstructured terrain under microgravity, such as on asteroids. Such bimodal mobility by a single actuation is expected to contribute to the enhancement of robotic exploration for space applications, whose available resources are limited. The experimentation for demonstration of the bimodal locomotion under microgravity is outlined below. Figure 2(a) shows the experimental environment that can simulate motion under planar microgravity. The inertial coordinate system, I{X, Y , Z}, is defined and fixed on the stone plate having 1 \u00d7 1 m in the X-Y plane. An air-floating testbed is used to simulate planar robot motion in microgravity. The testbed can freely move in the X-Y plane on the plate, and microgravity works in the negative direction of the Y -axis. The level of the planar microgravity is adjusted by changing the inclination angle of the stone plate", + " The sampling frequency of the tracking system is 100 Hz, and its position detecting accuracy is less than 0.1 mm in the following experiments. Additionally, a https://www.cambridge.org/core/terms. https://doi.org/10.1017/S0263574721000539 Downloaded from https://www.cambridge.org/core. Carleton University Library, on 27 May 2021 at 17:03:12, subject to the Cambridge Core terms of use, available at high resolution digital camera is used to observe the microscale behavior of the elastic body during the experiments. The frame rate and resolution are set to 30 fps and 3840\u00d72160 pixel, respectively. Figure 2(b) shows the air-floating testbed that simulates planar motion under microgravity. The testbed has a weight of 4.32 kg, a moment of inertia of 1.54 \u00d7 10\u22122 kgm2, and a size of 15 \u00d7 15 cm in the X-Y plane. The motor embedded in the testbed is a brushed DC motor (DCX32L, produced by maxon https://www.cambridge.org/core/terms. https://doi.org/10.1017/S0263574721000539 Downloaded from https://www.cambridge.org/core. Carleton University Library, on 27 May 2021 at 17:03:12, subject to the Cambridge Core terms of use, available at motor ag.) whose shaft is connected to an eccentric weight via clamp-coupling, as shown in Fig. 2(c). The eccentric weight has a diameter of 65 mm and a height of 23 mm. Its weight, moment of inertia, and eccentric distance from the rotational axle can be physically changed by adding small weights into the spaces. Moreover, the testbed has a micro-controller with a battery. Therefore, we can remotely control it via wireless communications. Throughout the experiments, we obtain time-series data of the motor rotational angle. Regarding surface elastic bodies, two polypropylene wire rods, whose diameter is 0.5 mm, are used. They are attached onto the side of the testbed as shown in Fig. 2(a). According to our previous work [13, 14, 15], as a benchmark of feasible design parameters of an elastic body, their inclination angle and natural length are, respectively, determined to be 70\u25e6 and 20 mm, as shown in Fig. 2(d). In the following experiments, the microgravity in the Y -axis direction is simulated such that the relative error from the desired value becomes less than a \u00b15%. With respect to the microgravity in the X-axis direction, it is set to approximately 10\u22125 G to not militate against the testbed motion. This section first presents motion analyses of the vibration propulsion mode based on the microgravity experiments. Then, it models the vibration propulsion. Furthermore, comparative analyses of the experiments and numerical simulations are elaborated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000226_icit.2003.1290227-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000226_icit.2003.1290227-Figure1-1.png", + "caption": "Fig. 1, Diagram of inverted pendulum", + "texts": [], + "surrounding_texts": [ + "For several decades, inverted pendulum systems have served as excellent test beds for control theory. Because they exhibit nonlinear, unstable, non-minimum phase dynamics, and because the full-state is not often fully measured, control objectives are always challenging. In [4] an inverted pendulum is taken as an example to demonstrate the fuzzy controller TSK design method. In [8] comparison of different fuzzy control laws of an inverted pendulum was presented. In [7] fuzzy logic controller with two sets of rules: first for swinging up the pendulum, and second for balancing the pendulum in upper position was studied Experimental robustness properties of fuzzy controllers remain theoretically difficult to prove and their synthesis is still an open problem. In this paper the user friendly method for tuning of fuzzy controller is used. The tuning process is making by using evolutionary algorithms." + ] + }, + { + "image_filename": "designv11_83_0002031_inmic.2008.4777752-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002031_inmic.2008.4777752-Figure1-1.png", + "caption": "Fig. 1 Geometry of the boundary-value problem", + "texts": [ + " In the , 2 4a' 2 4a case at hand, J.l = G(m + M) == GM , because m\u00ab M . Based on this theorem a formu-Iation to calculate transfer time and velocity vector at any instant during the boost phase is developed. This formulation is termed as the Lambert scheme. The symbol,f: is used to express functional relationship in the above equation (do not confuse with true anomaly). Therefore, one notes that the transfer time does not depend on true or eccentric anomalies of the launch point or the final desti-nation. Fig. 1 illustrates geometry of the problem. gyroscopes. (12) VI. THE INVERSE-LAMBERT SCHEME which says that v (current velocity), , (current position) and '2 (position of destination) are co-planar. For no cross-range error, the angle of inclination, \u00a2, between the velocity vector, v, and normal to the trajectory, n, must be 90\u00b0. In the Lambert scheme a subroutine computes devi-ation of \u00a2 from 90\u00b0. Extended-cross-product steering [10], dot- product steering [11] and ellipse-orientation steering [12] could be used to eliminate cross-range error" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002324_012099-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002324_012099-Figure1-1.png", + "caption": "Figure 1. Schematic diagram of the whole view mirror principle.", + "texts": [ + " Therefore, the harness recognition system should have high adaptability and robustness, and can still correctly recognize characters with certain noise and distortion. Instead of the three cameras in the industry for full-view image acquisition, it is difficult to obtain synchronization, image fusion algorithms are complicated, and real-time performance is poor. Innovatively propose a full-view mirror, and determine the appropriate angle based on the refraction and reflection modeling analysis of light. As shown in Figure 1. The principle of establishing the angle of the lens: modeling is based on geometric optics. It can be seen from the reflection theorem that an object formed by reflection can be equivalently regarded as a virtual image of the object directly imaging in the lens. Since the line diameter is small compared to the objective lens distance and lens size, and the image-to-lens distance is small compared to the image-tooptical axis distance, coaxial optics is used as the basic theory. The boundary conditions are established" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002646_s11517-021-02347-5-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002646_s11517-021-02347-5-Figure1-1.png", + "caption": "Fig. 1 The position and dimensional constraints of the robot inside the MRI bore: (a) the relative position between the patient and the robot, (b) the dimensional constraints", + "texts": [ + " The robot, which is operated inside a standard closed MRI bore during the prostate intervention, must satisfy the requirements of confined space, rigorous magnetic field environment and intervention operation. The considerations are described as follows. The diameter of the closed MRI bore is about 600 mm while the distance between the top and the scanner bed is about 470 mm. The patient lies on the scanner bed in the supine position with legs spread. The confined space in the MRI bore constrains the overall size of the proposed robot. The robot should be secured on the bed and near the perineum of the patient as shown in Fig. 1. It\u2019s critical to maintain a certain distance among the patient, the robot and the equipment used in the prostate intervention to avoid the mutual interference and guarantee the patient\u2019s safety. The overall size constraints of the robot are 320 mm \u00d7 270 mm\u00d7 200 mm. The MR compatibility is an important issue in the development of the robot. Based on the American Society for Testing Materials (ASTM) standard F2503-13, the term \u2018MR compatible\u2019 is redefined and classified into MR safe, MR conditional andMR unsafe" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002838_012021-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002838_012021-Figure8-1.png", + "caption": "Figure 8. Strain Distribution in Al 6061 T6", + "texts": [ + "The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.1.2. Stress Distribution The Max. and Min. Stress Distribution in Al 6061 T6 is 63.7 MPa and 0.3898 MPa respectively shown in Figure 7. 3.1.3. Strain Distribution The Max. and Min. Strain Distribution in Al 6061 T6 is 0.00096883 and 0.0000076477 respectively shown in Figure 8. 3.2. Analysing Testing Result of Structural Steel 3.2.1. Total Deformation The Max. and Min. Total Deformation in Structural Steel is 0.18644 mm and 0 mm respectively shown in Figure 9. The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.2.2. Stress Distribution The Max. and Min. Stress Distribution in Structural Steel is 183" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000790_s1560354708040072-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000790_s1560354708040072-Figure6-1.png", + "caption": "Fig. 6. Just before a collision of the pencil with the plane, the center-of-mass velocity v is perpendicular to the radius to the old instantaneous axis of motion. Vector v\u22a5 is the component of v perpendicular to the new axis of rotation. The momentum impulse P, assumed to occur at the new axis of rotation, has component P\u2016 along the radius that intercepts this axis.", + "texts": [ + " A more consistent analysis was given in [2]. 4.1. Asymptotic Velocities at the Beginning and End of a 1/N Turn If the pencil has N faces, then the angle between adjacent major radii is \u03b2 = 2\u03c0/N , as shown in Fig. 5. If a major radius has length a, then the width of a face is b = 2a sin(\u03b2/2) = 2a sin(\u03c0/N). At the end of a 1/N turn the center of mass of the pencil has velocity vector v perpendicular to the major radius from the center of mass to the edge that is the (old) instantaneous axis of rotation, as shown in Fig. 6. As the face of the pencil collides with the plane the instantaneous axis shifts to the adjacent major radius, and an impulsive force is exerted over the colliding face. REGULAR AND CHAOTIC DYNAMICS Vol. 13 No. 4 2008 An idealization is that this impulsive force is concentrated along the edge that becomes the new instantaneous axis. This assumption is not too plausible, but it leads to an analysis that has no free parameters (for given a and N of the pencil and angle of inclination \u03b8). Since the impulsive force acts by assumption only on the new axis of rotation, angular momentum L is conserved about this axis", + " Hence, the angular momentum of the pencil about the new axis just before the collision is L = ICM\u03c9 + mav\u22a5 = (k \u2212 1 + cos \u03b2)ma2\u03c9 = L\u2032 = kma2\u03c9\u2032, (32) invoking conservation of angular momentum about the new axis. Thus, \u03c9\u2032 \u03c9 = v\u2032 v = k \u2212 1 + cos \u03b2 k = 1 \u2212 1 \u2212 cos 2\u03c0 N k . (33) For example, \u03c9\u2032/\u03c9 = 11/17 for a solid hexagonal pencil (cos \u03b2 = 1/2), \u03c9\u2032/\u03c9 = 8/11 for a hollow hexagonal pencil, and \u03c9\u2032/\u03c9 = 1 in the limit of a circular pencil whether solid or hollow. We can confirm the result (33) by an analysis that includes the impulse on the edge newly in contact with the inclined plane. As shown in Fig. 6, we let P\u2016 be the component of the impulse along the line from the edge to the axis of the pencil, and P\u22a5 be the component transverse to this line. The changes in linear momentum of the pencil caused by these impulses are P\u2016 = mv sin \u03b2 = ma\u03c9 sin \u03b2, (34) P\u22a5 = mv cos \u03b2 \u2212 mv\u2032 = ma(\u03c9 cos \u03b2 \u2212 \u03c9\u2032), (35) while the change in angular momentum about the axis of the pencil is given by (k \u2212 1)ma2(\u03c9 \u2212 \u03c9\u2032) = ICM (\u03c9 \u2212 \u03c9\u2032) = \u2212aP\u22a5 = ma2(\u03c9\u2032 \u2212 \u03c9 cos \u03b2), (36) which leads again to Eq. (33). The solution given in [1] seems to be based on the assumption that P\u22a5 = 0, in which case Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003259_978-981-16-1769-0_32-Figure15-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003259_978-981-16-1769-0_32-Figure15-1.png", + "caption": "Fig. 15 Equivalent stress\u2014opposite sense", + "texts": [ + " When the surgeon actuates the handle, the inner rod moves to the left, resulting in the jaws being opened, as shown in Fig. 14b. The outer sleeve is capable of sliding over the outer tube of the forceps. It has two extreme positions, which are defined by the grooves in the connector, as shown in Fig. 8. The locking pin fixes the outer sleeve into either of these grooves, ensuring that the outer sleeve cannot move during operation. When the outer sleeve completely covers the jaws, it is in the S-I mode, as shown in Fig. 15a. When the sleeve is retracted to the other extreme position, and the jaws are uncovered, the device is in the Forceps mode, as shown in Fig. 15b. The locking pin depicted in Fig. 12a is inserted within the groove in the outer sleeve and the connector. Then the pin is rotated to lock the outer sleeve in either of the positions described above. This arrangement ensures that the outer sleeve does not move and disrupt the surgery. Challenges in the Design of a Laparoscopic Surgical Forceps 447 The surgeon can change the orientation of the jaws as and when required by rotating the knob clockwise or anticlockwise, as shown in Fig. 7. operating in the forceps mode" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000332_50003-4-Figure3.34-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000332_50003-4-Figure3.34-1.png", + "caption": "Fig. 3.34. The isotropic tread element with deflection e in two successive positions i- 1 and i. Its base point B moves with speed Vb and its tip P slides with speed Vg.", + "texts": [ + " Here, we will restrict the discussion to steady-state slip conditions and take a single zone with length equal to the contact length. In Section 2.5 an introductory discussion has been given and reference has been made to a number of sources in the literature. The complete listing of the simulation program TreadSim written in Matlab code is given in Appendix 2. For details we may refer to this program. Figure 3.33 depicts the model with deflected belt and the tread element that has moved from the leading edge to a certain position in the contact zone. In Fig.3.34 the tread element deflection vector e has been shown. The tread element is assumed to be isotropic thus with equal stiffnesses in x and y direction. Then, when the element is sliding, the sliding speed vector Vg, that has a sense opposite to the friction force vector q, is directed opposite to the deflection vector e. The figure depicts the deflected element at the ends of two successive time steps i- 1 and i. The first objective is now to find an expression for the displacement g of the tip of the element while sliding over the ground", + " The moment M z causes the torsion of the contact patch and is assumed to act around a point closer to its centre like depicted in Fig.3.19. A reduction parameter e~ is used for this purpose. More refinements may be introduced. For details we refer to the complete listing of the program T r e a d S i m included in the Appendix. With the displacement vector As (3.101) established we can derive the change in deflection e over one time step. By keeping the directions of motion of the points B and P in Fig.3.34 constant during the time step, an approximate expression for the new deflection vector is obtained. After the base point B has moved according to the vector As we have: gi - - with g i = - 9 e i (3.110) e i ei-1 + g i A s i e -1 i-1 Here, ei_ 1 denotes the absolute value of the deflection and gi the distance P has slided in the direction of -ei_~. In case of adhesion, the sliding distance gi = O. When the tip slides, the deflection becomes: 138 THEORY OF STEADY-STATE SLIP FORCE AND MOMENT GENERATION ,Hi qz,i ei - (3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001278_s1068798x08110087-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001278_s1068798x08110087-Figure8-1.png", + "caption": "Fig. 8. Threaded joints tested by Birger and Iosilevich, Klyachkin, and Repin (a), Klyachkin (b), and Kovgan (c).", + "texts": [ + " The accuracy of Eqs. (4)\u2013(6) is also verified by comparison with the experimental data in [1, 5\u20138]. The forces in six M12 screws (n = 6; pitch P = 1.75 mm) connecting annular flanges (thickness h1 = 18 mm and h2 = 36 mm) were measured in [1]. Experiments with thinner flanges are not considered, in order to ensure compliance with the St Venant principle. For the steel flanges, the external diameter D = 115 mm, and the internal diameter d = 25 mm. The screws are inserted over a circle with a diameter Ds = 84 mm (Fig. 8a). The joint is subjected to a tensile force F = 0\u2013 100 kN. The preliminary tightening force of the screws is 4.2, 13, 21.5, or 26.3 kN. Joints with combined loading by tensile force F and tipping moment M = 106F N mm are also conducted. In that case, Fti = 12 kN. The points in Fig. 9a correspond to the sum of the forces Fti and FsF loading the screw under the action of tensile force, and those in Fig. 9b to the force Fs1 loading screw 1 under the action of the tensile force and tipping moment. The curves are calculated from Eqs", + "25 mm) connecting a rectangular flange (thickness h1 = 32 mm) to a base (h2 = 0) were measured in [5]. In the tests, steel 45 flanges (E = 2.1 \u00d7 105 MPa) and SCh18 cast-iron flanges (E = 1 \u00d7 105 MPa) are subjected to tensile force F, and a steel 45 flange is subjected to tensile force F in combination with tipping moment M = 80F N mm. 1068 RUSSIAN ENGINEERING RESEARCH Vol. 28 No. 11 2008 IVANOV The flange dimensions are as follows: l = 200 mm; b = 100 mm; x1 = 80 mm; x2 = 40 mm; x3 = 0; x4 = \u201340 mm; x5 = \u201380 mm (Fig. 8b). The mean microirregularity height of the contact surfaces Ra1 = Ra2 \u2248 0.16 \u00b5m (finish class 10). The preliminary tightening force on the pin Fti = 6.25 kN. The results of these tests and calculations are shown in Fig. 9d, where the dark points and curve 1 correspond to the steel flange, while the crosses and curve 2 correspond to the cast-iron flange; and in Fig. 9e, where the pin number i is shown. The calculations assume that c0 = 500; \u03b5 = 2.64 for steel flanges and \u03b5 = 2 for cast-iron flanges; \u2206 = 4 \u00b5m; Wmax = 0.5 \u00b5m; l0 = 14 mm; l1 = 18 mm. It is found that \u03c7F = 0.04 and \u03c7M = 0.039 for the steel flange, while \u03c7F = 0.073 for the cats-iron flange. The forces in ten M16 pins (n = 10; P = 2 mm) connecting rectangular flanges (thickness h1 = h2 = 30 mm) were measured in [7]. The test data correspond to steel flanges (E = 2.1 \u00d7 105 MPa) under the action of tensile force F. The flange dimensions are as follows: l = 1000 mm; b = 40 mm (Fig. 8c). The mean microirregularity height at the contact surfaces Ra1 = Ra2 \u2248 2.5 \u00b5m (milling). The preliminary tightening force of the pins is Fti = 10, 20, 30, or 40 kN. The results of the tests and calculations are shown in Fig. 9f. It is assumed in the calculations that c0 = 500; \u2206 = 50 \u00b5m; Wmax = 0.007 \u00b5m; l0 = 32 mm; and l1 = 28 mm. Correspondingly, it is found that, for the four Fti values, \u03b5 = 2.5, \u03c7F = 0.334; \u03b5 = 2.25, \u03c7F = 0.247; \u03b5 = 2.11, \u03c7F = 0.205; and \u03b5 = 2.01, \u03c7F = 0.180. The inclined line beginning at the origin is the force acting on the pin when Fti = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002990_01423312211015124-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002990_01423312211015124-Figure2-1.png", + "caption": "Figure 2. Schematic of the electromechanical system.", + "texts": [ + " Moreover, the fixed settling time is bounded by T \u0142 2 k1e 1 g\u00f0 \u00de + 2 k2e d 1\u00f0 \u00de : \u00f047\u00de According to Equation (46), the small regions about the origin can be made arbitrarily small if the design parameters ki and hi are appropriately adjusted. This completes the proof. To ensure that readers have a clear understanding about the whole control design procedure, the block diagram of the proposed fixed-time adaptive fuzzy BC scheme is introduced in Figure 1. In this section, two simulation examples are performed to validate the proposed fixed-time adaptive fuzzy BC scheme. Example 1: The first example is a practical one borrowed from Dawson et al. (1994). Consider the following electromechanical system depicted in Figure 2: D\u20acq+B _q+Nsin q\u00f0 \u00de+D1 q, _q, t\u00f0 \u00de= t, L _t =V Rt KB _q+D2 q, _q, t\u00f0 \u00de, \u00f048\u00de where q is the angular motor position, t is the motor armature current, V is the input control voltage, D= J Kt + mL2 0 3Kt + M0L2 0 Kt + 2M0R2 0 5Kt , B= B0 Kt , N = mL0G 2Kt + M0L0G Kt , and D1 q, _q, t\u00f0 \u00de and D2 q, _q, t\u00f0 \u00de are the modelling errors. Here J is the rotor inertia, m is the link mass, M0 is the load mass, L0 is the link length, R0 is the radius of the load, G is the gravity coefficient, B0 is the coefficient of viscous friction at the joint, Kt is the coefficient which characterizes the electromechanical conversion of armature current to torque, L is the armature inductance, R is the armature resistance, and KB is the back electromotive force coefficient" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000948_026635108786260965-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000948_026635108786260965-Figure2-1.png", + "caption": "Figure 2. Wrinkle analysis schematic drawing.", + "texts": [ + " For given width of h (120mm), the strain at the membrane center is (1) Using the constitutive equation (2) \u03c3 \u03c3 x y xy E v v v v\u03c4 = \u2212 \u2212 1 1 0 1 0 0 0 1 2 2 \u03b5 \u03b5 \u03b3 x y xy \u03b5 \u03b5 \u03b3 \u03b4 \u03b4 x y xy h h = 0 2 1 the stresses at the membrane center are obtained by (3) Therefore, the minor principal stress is (4) That is (5) From the tension field theory, if the minor principal stress is zero, (6) Wrinkle occurs and the critical shear distance for wrinkle is given by (7) The wrinkles will appear on whole membrane surface when \u03b41 > \u03b41cr. It can be seen from the Equation. (7) that the critical shear distance is in proportion to the initial tensile distance, which confirms that increasing the initial pretension can effectively suppress the formation of the wrinkle. A wrinkling mode schematically plotted in Fig. 2 shows that the rectangular membrane generates the initial inner stress under the action of \u03b42, and the \u03b4 \u03bd \u03bd \u03b41 2 2 1cr = \u2212 \u03c3min = 0 \u03c3 \u03b4 \u03b4 \u03b4 min = \u2212 \u2212 + + E h v v2 1 1 2 1 2 2 2 \u03c3 \u03c3 \u03c3 \u03c3 \u03c3 \u03c4min = + \u2212 + +x y x y xy2 2 2 2 \u03c3 \u03b4 \u03c3 \u03b4 \u03c4 \u03b4 x y x Ev v h E v h E v h = \u2212( ) = \u2212( ) = +( ) 2 2 2 2 1 1 1 2 1 184 International Journal of Space Structures Vol. 23 No. 3 2008 transverse shearing under \u03b41. A local coordinate system (\u03be, \u03b7) is defined, where \u03be is the wrinkle direction, tension force per unit along the local axis is T (\u03be) and b (\u03b7)", + " Young\u2019s modulus of the membrane material is E, membrane width is h, the thickness is t, the half-wavelength of the wrinkle is \u03bb, the out-ofplane displacement function is w and \u03b8 is the included angle between the wrinkle stretching wire and bottom margin. To define the out-plane deflection of the membrane, the Von Karman large deflection formula is introduced, (8) where F is the stress function of wrinkled membrane, D is bending rigidity, and the operator L is (9) For the center regions of rectangular membrane, the Equation(8) turns to (10) We assumed that \u03be is parallel to the wrinkle direction, and \u03b7 is perpendicular to it. Based on the wrinkling configuration shows in Fig.2, we configure the out-ofplane displacement function as (11) where mh is the wrinkle length (m sin\u03b8 = 1). Substituting the Equation(11) into (10), then (12)b D D mh T mh D m \u03b7 \u03c0 \u03bb \u03c0 \u03be \u03bb \u03c0( ) = + + ( ) ( ) ( ) + 2 2 2 2 4 2 22 h( )2 w A mh \u03be \u03b7 \u03c0\u03be \u03c0\u03b7 \u03bb , sin sin( ) = D w T w b w \u2207 \u2212 ( ) \u2202 \u2202 + ( ) \u2202 \u2202 =4 2 2 2 2 0\u03be \u03be \u03b7 \u03b7 L x y x y x y y x = \u2202 \u2202 \u2202 \u2202 \u2212 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 + \u2202 \u2202 \u2202 \u2202 2 2 2 2 2 2 2 2 2 22 D w L w F\u2207 = ( )4 , The critical compression stress vertical to the wrinkle direction is (13) Then the half-wavelength can be written as (14) Therefore (15) where (16) For the membrane center regions, (17) Substituting Equation (1) into Equation (17), therefore (18) where (19) Along wrinkle direction, the stress is \u03c3\u03be, (20) Corresponding tension force T (\u03be) is, (21) Substituting Equation (21) into Equation (15), the expression of the wrinkle wavelength is given by (22) The strain along \u03b7 direction is supposed to be \u03b5\u03b7, which can be divided into two parts" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002993_s00170-021-07405-8-Figure10-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002993_s00170-021-07405-8-Figure10-1.png", + "caption": "Fig. 10 Simulation result and measured method. a Simulation result. b Tooth profiles. c Measured method", + "texts": [ + " Due to the sun gear has two kinds of tooth shapes (external gear and internal spline), and the inner metal flow in the forming process is extremely complex, whichmakes the teeth accuracy of sun gear difficult to control. In order to obtain an in-depth understanding of the deviation distribution laws of finished sun gear, the simulated results of sun gear formed by precision sizing operation were evenly divided into nine section planes. Then, the tooth profiles obtained from numerical simulation were imported into procedure UG to measure the deviations of internal-external teeth, as shown in Fig. 10. The maximum of single profile deviations (f\u03b1) was assumed to the total profile deviation, the maximum of differences between the target and finished tooth thickness (left and right surface) was the total helix deviation, and the maximum ofM value errors in the reshaped tooth and target tooth was defined as the total M value deviation. The influences of key process parameters on forming accuracy of external gear and internal spline were investigated through FE simulations to determine the optimal process parameters of precision sizing operation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000171_detc2004-57029-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000171_detc2004-57029-Figure3-1.png", + "caption": "Fig. 3 Mechanism diagram of 5-UPS/PRPU 5-axis PMT", + "texts": [ + " The constraint motion of the PRPU limb, the inverse solution equation and the Jacobian Matrix are presented. Finally, the workspace [15] and the dexterity [16] of the PMT are studied. The novel 5-UPS/PRPU 5- DOF parallel mechanism has been declared(No. 1371786) in CPO(the Patent Office of China). The prototype of the PMT is shown in Fig. 1 and the computer modeling is shown in Fig. 2. 1 Copyright \u00a9 2004 by ASME ms of Use: http://www.asme.org/about-asme/terms-of-use The parallel mechanism is shown in Fig. 3. The moving platform and the stationary platform are connected by five UPS(Universal-Prismatic-Spherical)limbs and a PRPU(Prismatic-Revolute-PrismaticUniversal) limb. The PRPU is set in the center of the mechanism. Like the Stewart platform, the mechanism is a close-loop mechanism. According to Kutzbach Grubler Formula, the DOF of the mechanism is given by \u2211 = +\u2212\u2212\u2217= g i ifgnM 1 )1(6 (1) where M is the DOF of the mechanism\uff0c n is the number of the component\uff0c g is the number of the kinematic pairs\uff0c is the DOF of the ith kinematic pair. if As shown in Fig. 3, \uff0c then 35,19,15 1 === \u2211 = g i ifgn Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/09/2016 Te So the DOF of the mechanism is 5. ( ) 53511915*6)1(6 1 =+\u2212\u2212=+\u2212\u2212\u2217= \u2211 = g i ifgnM 2. KINEMATIC PERFORMANCE OF THE MECHANISM 2.1 Constrained motion of the moving platform The PRPU limb is passive and contains no actuators. In the PRPU branch, the axis of the first grounded prismatic pair and the second revolute pair is set parallel to the base plane; the following prismatic pair is perpendicular to the second revolute axis; the first revolute axis of the universal joint is parallel to the second revolute axis; the second revolute axis of the universal joint is parallel to the base plane in the initial configuration", + " Thus the closed-loop control for the PMT may be realized and the machining accuracy of the PMT will be improved. When the position and orientation of the moving platform is determined, according to equation (6), the inverse solution of PRPU branch will be given by ( ) ( ) \u23aa \u23aa \u23aa \u23aa \u23a9 \u23aa \u23aa \u23aa \u23aa \u23a8 \u23a7 \u2212= \u2212= \u239f\u239f \u23a0 \u239e \u239c\u239c \u239d \u239b \u2212 \u2212 = \u2212+\u2212= += \u03b2\u03b8 \u03b8\u03b1\u03b8 \u03b2\u03b1 \u03b2\u03b1\u03b8 \u03b2\u03b1\u03b2\u03b1 \u03b2 4 23 4 4 2 2 4 2 42 41 ccaX csaYarctg csaYccaXa saZd BO A BO A BO A BO A BO A (7) nloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/09/2016 T As shown in Fig. 3, we establish the reference frame {A} on the stationary platform and the moving frame {B} on the moving platform. Given position and orientation parameter of the moving platform, the inverse solution can be obtained. Let ),,,,( \u03b2\u03b1BO A BO A BO A ZYX R denote orientation tranform matrix and [ ]TBO A BO A BO A BO A ZYX=P denote the vector of the origin of {B} expressed in {A}. The five joints of the stationary platform are arranged as shown in Fig. 1. One joint U1 is along the Y-axis of {A} and its coordinate value is W" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002925_01423312211016188-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002925_01423312211016188-Figure2-1.png", + "caption": "Figure 2. Additional displacement path of outer ring local fault.", + "texts": [ + " roller revolution speed,vin and vout are inner and outer ring rotation speed, respectively, generally vout = 0, d is roller diameter, D is pitch circle diameter, D= Din +Dout\u00f0 \u00de=2, Din and Dout are inner and outer raceway diameters, respectively, and a is contact angle. For cylindrical roller bearings and deep groove ball bearings without axial load, a= 0. Here df is the additional roller displacement. The additional displacement generated by local defects excites the system and makes bearing vibration signal carry pulse components. As shown in Figure 2, the additional displacement path cause by the outer ring fault can be divided into two stages. First, the roller continues to move after contacting with the local fault, and its centre makes the first stage circular motion around the fault front edge. For a small size fault, the roller will not touch the bottom. When the roller contacts with the fault front and back edges at the same time, the additional displacement df reaches the maximum value. The roller impacts the back edge and produces impact force excitation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000636_bfb0119382-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000636_bfb0119382-Figure1-1.png", + "caption": "Figure 1. The 4 phases of binding (wrapping a rope around a box). The RWI robots are denoted by disks. Bonnie is denoted by B; Clyde is denoted by C. The Pioneer helper is denoted by the gripper icon. Translation: From the state obtained in Phase 4, translations can be effected by having Bonnie and Clyde move with parallel, synchronous velocities. For example, from Phase 4, the box can be translated 'southwest' by commanding the same motion to B and C. For pure translations, B and C should be equidistant from the box. That can be accomplished using a ratchet (regrasp) step.", + "texts": [ + " The goal of this skill is for a team of two robots to surround an object with a rope to generate a rope grasp. The ends of the rope are tied along the \"waistline\" of each robot. The basic idea of the algorithm is to keep one robot s tat ionary while controlling the other robot to move around the object with the rope held taut . At some point, the moving robot has to cross over the rope in order to complete the grasp; this step requires coordination with the stationary robot. We have developed a control algorithm for this skill and implemented it using two RWI B14 robots and one Pioneer robot (see Figure 1). In the first phase, the B14 robots axe on the same side of the box. Clyde, the active robot (the right robot in Figure 1) moves so as to keep the rope taut around the box. Bonnie, the left robot in Figure 1, is passive and stays in place. The control for encircling the box is implemented as hybrid control of a compliant rotation. The robot starts with a calibration step to determine the level of power necessary to achieve rotational compliance to the box. The robot tries to move in a polygonal approximation to a circle. Due to the radial force along the rope, the robot's heading complies to remain normal to the radial force, thereby effecting a spiral. The robot continues this movement until it gets close to having to cross the rope", + " The algorithm for the second phase starts with Ben locating the corner of the object closest to the Bonnie. Ben uses its gripper to push the rope against the box, thus securing the rope grasp. Then Bonnie gives some slack in the rope. In the third phase, Clyde crosses over the rope. In the fourth phase, Bonnie tightens the rope. The tension in the rope signals the helper to let go and move away. In the translation skill, two robots translate a bound object by moving with parallel, synchronous velocities (see Figure 1). The basic idea of this algorithm is to regrasp start by regrasping a bound object so that (1) the two robots are equidistant from the box along a line called the rope line and (2) the rope line is perpendicular to the desired direction of translation. These two properties are achieved by using the active part of wrapping algorithm: keeping the rope taut, both robots rotate counter-clockwise about the box until the rope line becomes perpendicular to the desired direction of translation. At this point, a regrasp step is necessary to ensure that both robots are equidistant along the new rope line", + " Occasionally, Ben misses the corner and thus does not get a good grasp of the rope. The failures in flossing and ratcheting were due to signaling errors. Occasionally, the passive robot stops prematurely. We believe that these errors are due to low battery levels during those specific experiments, but more investigation is necessary. 4. A n a l y s i s In this section we provide a brief analysis of the manipulation grammar and the non-holonomic nature of the rope manipulation system. B. Binding (Section 2.2). T. Translation (Figure 1, post-Phase 4, and Section 2.2), O. Orientation (Section 2.2), and G. Regrasping (when the rope slips along the box in Section 2.2). Let us call these primitives B, T, O, and G, respectively. If O is an orientation step, let O R be the reverse step, performed by reversing the robots' direction. In this case, flossing can be written as (O, OR) *, the repeated iteration of O followed by O R. Note that no direct communication is required to switch between O and oR; the signal is transmitted through the task (the rope)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002836_tmag.2021.3083418-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002836_tmag.2021.3083418-Figure6-1.png", + "caption": "Fig. 6. (a) Slotless and (b) slotted shaftless PM motor geometry.", + "texts": [ + " It was shown in [11] that due to the nature of matrices LP and LR, their expression in (51) can be further simplified by eliminating the first term and rewriting the second by replacing the apparent permeability with the differential permeability, defined as: \u00b5\u2032r = 1 \u00b50 \u2202B \u2202H = \u00b5r + \u2202\u00b5r \u2202H H (52) = \u00b5r + \u2202\u00b5r \u2202B B (53) This results in the following Jacobian matrices, depending on the chosen formulation: Jnode = [ -L\u2217H L (12) F L\u2032P -L\u2217\u03a8 ] Jloop = [ -L\u2217B L (13) F L\u2032R -L\u2217f ] (54) with L\u2032P and L\u2032R the permeance and reluctance matrices respectively, computed with the differential permeability instead of the apparent permeability. Previous research [10]\u2013[12] has shown for MEC that the loop formulation exhibits better numerical properties compared to the node formulation. This is attributed to the fact the Jacobian matrix tends to have very large condition numbers in the node formulation, in addition to the shape of the residual functions hindering the convergence [12]. For illustration, the HAM is applied on slotless and slotted, shaftless permanent-magnet (PM) motors shown in Fig. 6. Dimensions tm, ta, tw and ts are 3.5, 1, 2 and 1mm, respectively. All regions have an axial length of 50mm. All regions are modeled using Fourier description, except the stator yoke that is made of NO20 steel and the slotted winding from Fig. 6 (b) that are both modelled using MEC. To analyze the convergence of the different methods and formulations under different saturation levels of the stator yoke, the strength of the PM is varied through its remanent magnetic flux density Brem. Authorized licensed use limited to: BOURNEMOUTH UNIVERSITY. Downloaded on July 03,2021 at 19:02:58 UTC from IEEE Xplore. Restrictions apply. 0018-9464 (c) 2021 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www", + " This error is related to the Dirichlet limit condition in FEA that is set at 5 times the stator outer radius, while in HAM the Fourier representation accurately extends to infinity. The error is increasing with Brem due to the impact of saturation of the stator yoke. The cogging torque is computed using the Maxwell stress tensor method by defining an integration path in the airgap across the whole physical domain [23]: Tcog = R2 magLm \u00b50 \u222b qp 0 BqBp \u2223\u2223 p=Rmag dq (58) with Rmag the mean airgap radius and Lm the motor axial length. The magnetization angle of the PM is then varied over a complete cogging torque period, i.e. \u03c0/3 for the motor represented in Fig. 6 (b). The result is shown in Fig. 11, where a good agreement is shown between HAM and FEA. Subsequently, the impact of the HAM mesh size is investigated on the cogging torque. Table IV gives, for different HAM mesh sizes, the mean squared relative error on the cogging torque compared to FEA and compared to the finest HAM mesh. It confirms the HAM model converges when the mesh size is increased, but not to the same exact value as the FEA model. The reasoning behind this is related to the Dirichlet limit condition in FEA, as explained in previous paragraph, and difference in implementation of the same BH curve" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002116_s12541-020-00407-8-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002116_s12541-020-00407-8-Figure4-1.png", + "caption": "Fig. 4 Error due to poor adjustment of contact point", + "texts": [ + " But in presence of resistance force, slippage and elastic deformation happen and cause the backward movement of the roller (Fig.\u00a03b, c). In this case, sum of backward movements should be less than forward movement, else the roller doesn\u2019t move or moves in opposite direction. 3. Accuracy of roller angle: It is almost impossible to guaranty an accurate and fixed roller angle. It depends on bearings, mechanism structure, etc. 4. Misalignment of the contact point: When applied preload, shaft center and roller center are not aligned, the contact point between roller and shaft will change 1 3 (Fig.\u00a04). Under mentioned condition, effective angle between roller and shaft will be less than nominal value. 5. Manufacturing and assembly: Surface roughness, accuracy of parts, unwanted deflection of parts. In this section, three topics are discussed: the lead of mechanism, the preload applying methods, and the pressure at the contact point. The lead of the mechanism ( \u0394x ), is the linear displacement during one revolution of the shaft and is related to the angle between axes of the roller and shaft, , and the diameter of the shaft, d , as follows: According to Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000732_j.isatra.2007.11.002-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000732_j.isatra.2007.11.002-Figure4-1.png", + "caption": "Fig. 4. Two stand model of roller section.", + "texts": [ + " Meantime, the supervisory control observed the actual speed of the drive and the output of the neuro-fuzzy controller which send the parameter to the GA for searching parameters. These searched parameters are fed to the neuro-fuzzy controller to generate the torque, and to control the drive. The proposed scheme optimizes the loop formation effectively. It also circumvents the usage of tension meter. Therefore, the proposed methodology reduces the complexity of the mill and ensures the safety operating environment. The two stand model of a steel rolling mill with the rollers is shown in Fig. 4. This model is used to predict the strip tension between the inter-stands by the known parameters like strip modulus of elasticity, strip thickness, speed difference between stands, distance between stands and strip temperature. The strip tension observer can be implemented based on the drive torque components, related to the tension between the inter-stands. This estimator would be dynamic and provide constantly updated observer computation of the strip tension between inter-stands. It is assumed that no slip occurs between the rollers and the steel strip" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001426_detc2007-34051-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001426_detc2007-34051-Figure1-1.png", + "caption": "Figure 1 Vibration model", + "texts": [ + " In this paper, authors continue to use the concept of measuring vibration and applying accurate transfer function measured experimentally to assess the vibration excitation correspondent with the common part of tooth surface form. The objective of this paper is to show the possibility of evaluating geometrical situation between tooth surfaces of mating gears in operation from measured vibration signal. Rotational vibration of a gear system can be simply considered as a one degree of freedom system as shown by the model in Fig.1. Gear blanks and shafts are treated as rigid components, and meshing part are modeled as a spring and a damper. The spring corresponds to the variation of tooth stiffness during meshing that is the cause of meshing vibration components. In addition, vibration also comes from gear geometrical errors attributed to the irregularities during manufacturing process or the assembling procedure. These kinds of gear excitations are modeled as the displacement errors e in the meshing part. From this model, the equation of motion can be expressed along the line of action as, WtextKxCxM =\u2212++ ))()((&&& " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000636_bfb0119382-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000636_bfb0119382-Figure2-1.png", + "caption": "Figure 2. The 2 phases of rotating an object by \"flossing\". The robots are denoted by disks. V~ denotes the velocity of the active robot. Vp denoted the velocity of the passive robot. Fp is the force felt through the rope by the passive robot.", + "texts": [ + " At this point, a regrasp step is necessary to ensure that both robots are equidistant along the new rope line. From this position, the robots can translate by commanding synchronous velocities. In the flossing skill, two robots rotate an object in place by coordinating movements of the rope between them. This type of manipulation is similar in flavor to the pusher/steerer system described in [6]. Flossing is preceded by a binding operation. We have developed a control algorithm for this skill and implemented it using two RWI B14 robots (see Figure 2). In the first phase the active robot translates forward. The passive robot follows along, feeling force Fp and keeping the rope taut. This action causes the object to rotate. This phase terminates when the passive robot stops in place, thus increasing the tension in the rope. The active robot senses this impediment and stops. At this point, the robots can reverse the passive/active roles and repeat the first phase. When Bonnie is the active robot, the object rotates clockwise; when Clyde is the active robot, the object rotates counterclockwise" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001722_ifost.2008.4602951-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001722_ifost.2008.4602951-Figure4-1.png", + "caption": "Fig. 4. Bifurcation of mode.", + "texts": [ + " If the rod remains rectilinear and is exposed to the further compression, a square of frequency becomes negative that signals that such configuration is unstable. The rod loses stability and displacement of variable end point appears. In a point where displacement becomes equal to length of a rod (buttresses have coincided, rod geometry is a loop), at 2.18 CLP P\u2248 \u22c5 , first own frequency becomes equal to zero, that could be interpreted as a bifurcation of mode of balance of a looplike configuration of a rod, loop can rotate (fig.4.). Then, at increase in compressing force, displacement becomes more lengths of a rod; the square of first own frequency becomes negative, that testifies to instability of the obtained configuration of a rod [5]. Such configuration is physically unrealizable. If was possible to keep a rod from primary loss of stability it can lose stability again, under other form. Figure 5 shows dependences of squares of own frequencies and displacement on value of compressing force at loss of stability on the second critical force ( )2(2) 22CL EJP \u03c0= " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000878_j.tsf.2008.09.048-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000878_j.tsf.2008.09.048-Figure1-1.png", + "caption": "Fig. 1. Illustration of Delaunay's method. f indicates the focus of an ellipse. A sphere (a), neck-like structure (b) and tube with constant mean curvature (c) are constructed.", + "texts": [ + " (7) can be neglected, this equation can be modified into H\u2212 \u03b3 \u00fe \u03b32\u22122g0k11 1=2 2k11 \" # H\u2212 \u03b3\u2212 \u03b32\u22122g0k11 1=2 2k11 \" # \u00bc 0 \u00f014\u00de therefore, the surface can be constructed by linking two kinds of constant mean curvature (HN0 and Hb0). Among such surfaces, the rotationally symmetric hypersurface can be constructed intuitionally by Delaunay's method [10]: by rolling a given conic section on a straight line in a plane to obtain the trace of a focus, and then rotating this trace around the line, one can get a undulated surface which satisfies Eq. (14), as illustrated in Fig. 1. Here, the conic section is assumed to be an ellipse with length semiaxes a and b (aNb), where a \u00bc k11 \u03b32\u22122g0k11\u00f0 \u00de1=2\u2212\u03b3 : \u00f015\u00de A beaded structure is obtained by the Delaunay constructionwith a thin ellipse section (Fig. 1a). If the ellipse is very flat, then the beaded structure becomes vesicles of necklace-like structure (Fig. 1b). On the other limit, if the ellipse becomes a circle the resulting surface is a cylindrical tube (Fig. 1c). Indeed, all these shape transitions are observed in the experiment by varying CI of the dipeptide. Moreover, our theory can even make a rigorous prediction on the geometric characteristics of the necklace-like structure, as below. Denoting the maximum and minimum radii of the undulated surface as \u03c1max and \u03c1min, one can show that \u03c1max=a+(a2\u2212b2)1/2 and \u03c1min=a\u2212 (a2\u2212b2)1/2, thus the average radii of the undulated cylinder is (\u03c1max+\u03c1min)/2=a. On the other hand, a spherical surface of radius 2a also satisfies Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002051_pime_conf_1964_179_063_02-Figure13.2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002051_pime_conf_1964_179_063_02-Figure13.2-1.png", + "caption": "Fig. 13.2. Types of fluid-film plain bearing", + "texts": [ + " In this way in theory the two metal surfaces never contact one another, and as a result wear should be negligible and a very long life should be obtained. Since the two surfaces are being pressed together by the applied load, a pressure has to be produced in the film of fluid in order to resist this effect and keep the surfaces separated. There are two different ways by which this can be done, Val I79 Pr 3D at UNIV NEBRASKA LIBRARIES on June 9, 2016pcp.sagepub.comDownloaded from 6 iU. J. NEALE as shown in Fig. 13.2. In the first method, commonly called a hydrostatic bearing, the space between the surfaces is fed with fluid from an external pumped supply at a pressure sufficient to maintain the fluid film. The second method of obtaining a pressure in the fluid film between the surfaces is to generate it hydrodynamically, by making the surfaces of a suitable shape and flooding them with a viscous fluid. The principle is a well-known one and involves making the bearing surface in the form of one or more taper wedges" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003537_s40430-021-03132-z-Figure10-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003537_s40430-021-03132-z-Figure10-1.png", + "caption": "Fig. 10 Meshed model for three characteristic positions and stress at all active contacts of star and ring", + "texts": [ + " Therefore, to estimate deformations, gaps and contact stresses at active contacts in GEROTOR unit, a separate FEM analysis than that adopted in case of ORBIT motor (Roy et.al. 2019) is required. The CAD models of star-ring set of GEROTOR unit have been developed in SolidWorks by importing geometric data from MATLAB. The geometric generation of starring profiles is briefed in Appendix 1. The geometric design parameters of the considered GEROTOR pump are detailed in Table\u00a01. For structural analysis of the starring set, this CAD model has been imported into Ansys\u00ae environment referring to the feature in Fig.\u00a010. The static structural module has been used in Ansys\u00ae for the present FEM analysis. SOLID187 element which is a higher order 3-D, 10-nodes (each node has three degrees of freedom) element has been used for the present FEM model. It has a quadratic displacement behavior and is well suited for modeling irregular meshes (such as those produced from various CAD/CAM systems). The element has plasticity, hyperelasticity, creep, stress stiffening, large deflection and large strain capabilities. It also has mixed formulation capability for simulating deformations of nearly incompressible elastoplastic materials, and fully incompressible hyperelastic materials", + "1125 Nominal width of star-ring (b) (mm) 20 Radius of outer centrode (base circle of the envelope) (R0) (mm) 26.4145 1 3 1 3 in terms of grid independence/convergence test, refinement technique under mesh control feature has been used. The refinement technique is applied at the edges of the star and ring lobe portion. Using the above said technique, mesh has been generated for different rotational angle of star w.r.t its own axis, out of them, only the meshes for \u2212 30\u00b0, 0\u00b0 and + 30\u00b0 of star rotational angles are shown in (Fig.\u00a010a, e and i). A mesh independence study has also done to get close to the accurate results. This study consists of refining elements and comparing the refined solutions to the coarse solutions. If further refinement (or other changes) does not significantly change the solution, then these nodes and elements numbers can be finalized. Journal of the Brazilian Society of Mechanical Sciences and Engineering (2021) 43:429 Page 13 of 20 429 It has been considered that both the metal to metal active contact surfaces of the star and ring are constrained from penetrating each other", + " The star-ring set remains in a state of balance because the torque on the shaft will be equal to the torque generated by fluid pressure. The linear displacement of the model must be fixed along Z axis, and it can only revolve along this axis. The ring part GEROTOR depicted in Fig.\u00a03 rotates along its own axis O which is also aligned with Z axis. Referring to Fig.\u00a03, the star-ring positions have been considered from \u03b8 = \u2212 30\u00b0 to + 30\u00b0 shaft rotations with an increment of 10\u00b0 for FEM modeling. Illustrations in Fig.\u00a010b\u2013d, f\u2013h, j have shown the stresses at star-ring set for each angular position (i.e., from \u2212 30\u00b0 to + 30\u00b0 with an increment of 10\u00b0, means total seven different positions of star-ring set) of star loaded with the total load resulting from the shaft torque (T) = 1.8\u00a0N-m and the differential pressure ( \u0394pi ) of 1\u00a0MPa. The results have been presented for low pressure as the experiment has also been carried out in low pressure to avoid the bulging out of the transparent end cover of the pump used for flow visualization in photo imaging technique (vide [32]. Journal of the Brazilian Society of Mechanical Sciences and Engineering (2021) 43:429 1 3 429 Page 14 of 20 At the starting position, when star angular rotation amount \u03b8 = \u2212 30\u00b0 as in Fig.\u00a010a and b, those chambers formed between the two consecutive active contacts (chambers C1, C2 and C3) are in expansion mode. After anticlockwise direction of rotation of both star and ring, the star comes to a new position, i.e., \u03b8 = 0\u00b0 as per Fig.\u00a010e and f. Here, chamber C3 (in LPZ) is in expansion, while chamber C4 is in compression and due to the differential fluid pressure, a gap is created at the transition contact, i.e., 4\u20134\u2032 lobe contact point for a small duration of time. This expansion of the chamber C3 occurs until the star\u2019s new rotational position, i.e., \u03b8 = + 30\u00b0 (vide Fig.\u00a010j). The meshed model and the state of stresses for the initial position of the star-ring set, i.e., \u03b8 = \u2212 30\u00b0 are presented in Figs.10a and b, respectively. Mesh generation and boundary condition (means consider the outer periphery of the ring is fixed) for this starting position are shown in Fig.\u00a010a. Figure\u00a010b shows that in successive tooth pairs 1\u20131\u2032, 2\u20132\u2032, 3\u20133\u2032, 4\u20134\u2032 and 7\u20131\u2032 stresses occur, whereas in pairs 5\u20135\u2032 and 6\u20136\u2032, no stresses occur. The highest stresses can be observed at the root of star lobe 1\u2032. The magnitude of maximum stress for this star-ring position is 74\u00a0MPa which is lower than the maximum acceptable stress. The occurrence of stress causes mutual deformation of star and ring lobes at their active contact in successive tooth pairs 1\u20131\u2032, 2\u20132\u2032, 3\u20133\u2032, 4\u20134\u2032 and 7\u20131\u2032. Due to this differential fluid pressure, gaps are generated at lobe 5\u20135\u2032 and 6\u20136\u2032 as they are disengaged. At \u03b8 = \u2212 20\u00b0 (Fig.\u00a010c), maximum stress of magnitude of 81\u00a0MPa occurs at the root of lobe 1 of star and ring lobe 1\u2032, i.e., here at the contact point. At \u03b8 = \u2212 10\u00b0 (Fig.\u00a010d), gap is generated at transition active contact, i.e., at 4\u20134\u2032. In this position, stress and deformation occur at 1\u20131\u2032, 2\u20132\u2032, 3\u20133\u2032 and 7\u20131\u2032 whereas gaps are created at 4\u20134\u2032, 5\u20135\u2032 and 6\u20136\u2032. Similar to previous two positions, here also maximum stress of magnitude 98\u00a0MPa occurs at 1\u20131\u2032 active contact. When star is at the central position, which means at \u03b8 = 0\u00b0, maximum gap is generated at 4\u20134\u2032 active contact point and maximum stress is generated at 7\u20131\u2032 contact. Gaps occur at the contacts which correspond to high-pressure chamber side. In Fig.\u00a010f, it can be shown that gaps occur at 5\u20135\u2032, 6\u20136\u2032 and also in the transition active contact 4\u20134\u2032. Though maximum stress occurs at the contact of 7\u20131\u2032, the magnitude (75\u00a0MPa) is comparatively less compared to maximum stress occurred in these previous star-ring set positions. Fig.\u00a010g at \u03b8 = + 10\u00b0 shows that stresses occur at active contacts 1\u20131\u2032, 2\u20132\u2032 and 3\u20133\u2032, respectively, whereas there are no stresses in lobe contacts between 4 and 4\u2032, 5\u20135\u2032, 6\u20136\u2032, 7\u20131\u2032. Here, the maximum stresses can be observed at the root of lobe 1\u2032 of star of magnitude 84\u00a0MPa which indicates again rise in stress as star-ring set passes the central position. Figure\u00a010h shows the state of stress for the star-ring position of \u03b8 = + 20\u00b0. Here, due to fluid pressure and acted torque, deformations and stresses are generated at active contacts 1\u20131\u2032, 2\u20132\u2032 and 3\u20133\u2032, respectively. In the other active contacts, i.e., at 4\u20134\u2032, 5\u20135\u2032, 6\u20136\u2032 gaps are created. Here, lobe 7 of the ring is in contact in a portion of the star in-between lobe 1\u2032 and 6\u2032. For this star-ring set position, in this portion of the star, maximum stresses (80\u00a0MPa) occurred. In these structural analyses, the last position is at \u03b8 = + 30\u00b0 (vide Fig.\u00a010j) new active contacts are in compression mode, these are 1\u20131\u2032, 2\u20132\u2032, 3\u201332\u2032, 6\u20136\u2032, 7\u20136\u2032. Only 4\u20134\u2032and 5\u20135\u2032 active contacts are stress-free as the gaps are created due to fluid pressure. In these positions, maximum stresses occur at the active contact 6\u20137\u2032 which has a magnitude of 76\u00a0MPa. From these analyses, it can be said that for certain shaft rotational positions at the transition active contact, gaps are generated, which in terms shows zero stress in results. Along with this, it also can be said that as no mechanical seal has been used in the operating zone, leakages during operation through the transition contact are unavoidable" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003376_icuas51884.2021.9476688-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003376_icuas51884.2021.9476688-Figure7-1.png", + "caption": "Fig. 7: Obstacle Course Swap experiment results. The two Tello UAVs need to swap positions going through an obstacle course. The start point of the first Tello is point A and its goal point is B and vice versa for the second Tello. The blue curve is the planned trajectory for the first Tello with the red being the actual trajectory as it was captured by the MoCap system. For the second Tello, the planned trajectory is in green with the actual one in magenta.", + "texts": [], + "surrounding_texts": [ + "Authors in [5] determined that Kinodynamic RRT* is asymptotically optimal and the probability of finding the optimal trajectory approaches 1 as the iterations approach infinity. The tightly-coupled approach shares this quality, as no change was made to the core algorithm, the only changes made were to the CollisionFree method and the plant. The loosely-coupled variant however does not share this quality. Figure 9b shows the cost of swapping two Tellos locked to a specific height. The Tellos begin at [0.25, 0.5]m and [0.75, 0.5]m and are restricted to a 1x1 square, as shown in figure 9a. The original model is used, with the vertical controller removed. Initially, the sequential case is able to find the faster solution by optimally connecting Tello 1\u2019s trajectory and planning Tello 2\u2019s trajectory with respect to it. However, given sufficient time, the simultaneous method finds better routes than the sequential method. This is an example of a problem in which the sequential case will not converge to the optimal solution, but the simultaneous planner will. Although, for more complex problems such as the ones used 874 Authorized licensed use limited to: University of Glasgow. Downloaded on August 18,2021 at 05:21:14 UTC from IEEE Xplore. Restrictions apply. in the experiments presented here, the additional time that the simultaneous planner needs to find a better solution is too high, making it an infeasible solution. For example, for the Obstacle Course Swap experiment, using the data shown in Figure 4, we extrapolated that the cost of the simultaneous method would become equal or lower than the sequential method around the one millionth iteration. This extrapolation is probabilistic and is not guaranteed to be lower at this point. Additionally, there are potential deadlock conditions in the sequential case. The possibility of deadlocks alone can remove the optimality guarantee, since the problem has a solution, but the sequential method may never be able to find such a solution given infinite iterations. Deadlocks may occur when the trajectory of higher priority robot blocks the trajectory of lower priority robot at some critical time until and including the final point in the trajectory. Figure 10a shows one such scenario where a deadlock may occur. If the states of the team were planned simultaneously, Tello 1 would have the chance to move down, allowing Tello 2 past. Using sequential planning however, a deadlock will occur. Figure 10b shows the sequential planner in the deadlock described above. Tello 1 starting at [0.6, 0.8]m is able to find a valid trajectory to [-0.8, 0.8]m (shown as a green line). Tello 2 starting at [-0.6, 0.8]m and ending at [0.8, 0.8]m is able to trajectory plan in the region that Tello 1 has not yet arrived. The farthest that Tello 2 is able to travel is the approximate middle of Tello 1\u2019s trajectory, as they will both arrive around the same time. Figure 10c shows the simultaneous solver with a solution to this deadlock situation. If there is a valid solution, the simultaneous solver will find it, given sufficient iterations." + ] + }, + { + "image_filename": "designv11_83_0002878_0954406221999077-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002878_0954406221999077-Figure5-1.png", + "caption": "Figure 5. Test setup of static characteristics.", + "texts": [ + " The maximum tensile displacements are 17.84mm and 14.04mm, and the fiber deformation rates are 5.23% and 4.13%, respectively. The deniers of UHMWPE and Aramid 1414 are 1600 and 1500D respectively. According to the denier formula Da\u00bc 9000 100 q pd2/4, where the crosssectional area of a single fiber is calculated for pd2/ 4, the diameters of the two fibers are 0.48 and 0.38mm at natural state, respectively. The material properties of the fibers is shown in Table 3. The test device for the WHAMs is shown in Figure 5. This device consists of a mounting part, power source part and controller part. One end is connected to the upper mounting plate a connector and turnbuckle for adjusting the contraction, and the other end is connected to the tension sensor and the joint bearing to prevent the axis from being eccentric. The displacement sensor is mounted on the middle part to record the contraction of the WHAM in the test. A water hydraulic test platform manufactured by the Finnish HYTAR Company provides power source" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002924_iemdc47953.2021.9449510-Figure14-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002924_iemdc47953.2021.9449510-Figure14-1.png", + "caption": "Fig. 14. The three DP-STVM models for the comparison of cogging torques and no-load back EMFs.", + "texts": [ + " 13, after independently applying rotor modification or stator parameter optimization, the cogging torque can be reduced to 10 Nm and 14 Nm. Moreover, when both methods are applied, the final cogging torque can be further reduced to less than 0.5 Nm, which is only 2% of the original value. In conclusion, the cogging torque in irregular DP-STVMs can be well limited through the reduction of target field harmonics. V. PERFORMANCE COMPARISON OF PROPOSED METHOD AND ROTOR SKEWING METHOD The final model of DP-STVM with proposed cogging torque reduction methods is shown in Fig. 14 (c), which has unevenly distributed and unequal width stator flux modulation teeth, and modified outlines of rotor magnets & teeth. As a comparison, rotor skewing method is also applied in proposed machine to reduce the cogging torque. Since the period of cogging torque is 48th, 3.75 deg is selected as the skewing angle in 2-stage skewing method, which corresponds to 30 deg of rotor elec. angle. The 6-slot-18-Ps16-Pr DP-STVM with regular evenly distributed stator teeth is shown in Fig. 14 (a), the model with ktp = 1.1 and skewed rotor is shown in Fig. 14 (b). The cogging torque comparison of DP-STVMs with these two methods and regular model are presented in Fig. 15, which shows that the proposed method has much better cogging torque reduction effect than the skewing method, the cogging torque is close to the model with evenly distributed stator teeth. Fig. 16 shows the rms value of phase no-load back EMF of three models, it shows that the proposed topology can have higher no-load back EMF than the regular DP-STVM model as well as the irregular model with skewed rotor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002772_978-3-030-67750-3_14-Figure20.2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002772_978-3-030-67750-3_14-Figure20.2-1.png", + "caption": "Fig. 20.2 Electrical braking system (EBS)", + "texts": [ + " A disadvantage age of the current technology is the captured image signal that is influenced by environmental brightness. In 2004 a commercial version of the automatic parking assistance was introduced by Toyota Motor Corporation in the Toyota Prius. Lexus 2007 LS also has an advanced parking guidance system. In automatic parking, there are two requiredmovements, which are the vehiclemovement (in reverse mode) and the steering angle. These movements are not discussed in this paper. The electrohydraulic brake shown in Fig. 20.2 is important for automatic reverse parking because the braking effect helps to decelerate the vehicle. The system has an electric motor and when operated can replace the action of manual brake by using the feet. The speed of car in the reverse mode automatic parking is determined by the propulsion motor and brake system. The brake system is used to decelerate, and stop at a fixed parking before the final hand brake action is applied. The brake force needs to be applied to the disc brake while controlling the motor speed so as to create the required deceleration effect" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002859_s12046-020-01541-9-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002859_s12046-020-01541-9-Figure6-1.png", + "caption": "Figure 6. Effect of section thickness on the volume fraction of phases: (a) pearlite and (b) ferrite.", + "texts": [ + " Too less graphite formation can lead to the formation of carbides in a thin section of castings and shrinkage defect formation; therefore it is essential to use additive alloys such as potent inoculants in sufficient quantity, which will promote heterogeneous nucleation of graphite and ultimately increase the formation of graphite. 3.2 Effect of section thickness on microstructure The microstructure constituent fractions of different sections were evaluated by simulation and image analysis of optical micrographs of samples. It is observed that at higher cooling rates the pearlite formation is more, whereas at a lower cooling rate the formation of ferrite is facilitated. The distribution of pearlite and ferrite phases across the stepbar casting is shown in figure 6a and b; it can be observed that as the section thickness increases the pearlite percentage increases, whereas the ferrite percentage decreases. The average pearlite fraction in 5-mm section thickness is 65\u201370% and in 10 and 15 mm it is 55\u201360% and 50\u201355%, respectively, as given in figure 6a. Similarly the average ferrite fraction in 5-mm section thickness is 20\u201325%, and in 10 and 15 mm it is 25\u201330% and 33\u201335%, respectively, as given in figure 6b. The variation of phase in different sections of stepbar castings is attributed to the cooling rate; at faster cooling rates the extent of diffusion of carbon from austenite is less, which promotes the transformation of austenite to carbon-rich phase pearlite. Therefore, it is evident that the thin section will have more pearlite fractions. However, in thick sections, the carbon is easily diffused towards the graphitization centres, leaving behind carbon-depleted regions, which transform to ferrite" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002618_012042-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002618_012042-Figure1-1.png", + "caption": "Figure 1. Double mathematical pendulum.", + "texts": [ + "1088/1757-899X/1129/1/012042 In this regard, the main attention in this article is focused precisely on the analytical study of nonlinear oscillations of double pendulum using specially developed asymptotic methods. The main purpose of this article is obtaining and interpreting formulas that make it possible to analyze in detail the nonlinear modes of oscillations of a double pendulum and establish their main qualitative features, and to accompany solutions with visual graphic illustrations as well. We consider the flat double mathematical pendulum with the same links of length l and identical end loads of mass m (figure 1). Such a system can be interpreted as the simplest nonlinear mechanical system with two degrees of freedom. It is most convenient to choose the absolute angles 1 and 2 of deviations of the pendulum links from the vertical as generalized coordinates. We note that sometimes relative angles 1 and 2 of rotation in the joints of the pendulum are taken as generalized coordinates [6]. They are related to absolute angles by obvious relationships: 1 1 and 2 2 1 . The angles 1 and 2 are often used in tasks of control motions because they are measurable" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000016_iscas.2005.1465271-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000016_iscas.2005.1465271-Figure3-1.png", + "caption": "Fig. 3. Structure of the NN.", + "texts": [ + " The multilayer backpropagation networks is especially useful for this purpose, because of its inherent nonlinear mapping capabilities, which can deal effectively for real-time online computer control [6]. However, problem of time consuming because of learning hours will be arise when use NN in actual plant, it is necessary to reduce learning hours by online tuning NN weights. The NN of proposed method has 3 layers: an input layer with 4 neurons, a hidden layer with 4 neurons and an output layer with one neuron as shown in Fig.3. Let xi(k) be the input to the i-th node in the input layer, pj(k) be the input to the j-th node in the hidden layer, q1(k) be the input to the node in the output layer. Furthermore, wji be the weight between the input layer and the hidden layer, w1j be the weight between the hidden layer and the output layer. The NN is used to compensate the nonlinearity of plant dynamics that is not taken into consideration in the conventional MRAC. The input xi(k) to NN is given by x(k) \u2261 {v(k \u2212 d), ym(k + d), y(k \u2212 1), u(k \u2212 d)} (15) The relations between inputs and output of NN is expressed as follows pj(k) = \u2211 i wji(k)xi(k) (16) q1(k) = \u2211 j w1j(k)f(pj(k)) (17) v\u0304(k) = f(q1(k)) (18) where f(", + " \u2202y(k) \u2202u(k \u2212 d) . \u2202u(k \u2212 d) \u2202f(q1(k)) . \u2202f(q1(k)) \u2202q1(k) . \u2202q1(k) \u2202f(pj(k)) . \u2202f(pj(k)) \u2202pj(k) . \u2202pj(k) \u2202wji(k) (24) where \u2202E(k) \u2202y(k) = \u2212(ym(k) \u2212 y(k)), \u2202u(k \u2212 d) \u2202f(q1(k)) = 1, \u2202q1(k) \u2202w1j(k) = f(pj(k)), \u2202q1(k) \u2202f(pj(k)) = w1j(k), \u2202pj(k) \u2202wji(k) = xi(k), \u2202f(q1(k)) \u2202q1(k) = \u00b5 2a (a \u2212 f(q1(k)))(a + f(q1(k))), \u2202f(pj(k)) \u2202pj(k) = \u00b5 2a (a \u2212 f(pj(k)))(a + f(pj(k))) \u2202y(k)/\u2202u(k \u2212 d) is given by \u2202y(k) \u2202u(k \u2212 d) = \u2202y(k) \u2202v(k \u2212 d) . \u2202v(k \u2212 d) \u2202u(k \u2212 d) = \u2202y(k) \u2202v(k \u2212 d) / \u2202u(k \u2212 d) \u2202v(k \u2212 d) (25) From Fig.3, we obtain u(k \u2212 d) = v(k \u2212 d) + v\u0304(k \u2212 d) (26) \u2202u(k \u2212 d) \u2202v(k \u2212 d) = 1 + \u2202v\u0304(k \u2212 d) \u2202v(k \u2212 d) (27) According to (15) and (18), \u2202v\u0304(k \u2212 d)/\u2202v(k \u2212 d) can be obtained as \u2202v\u0304(k \u2212 d) \u2202v(k \u2212 d) = \u2202f(q1(k)) \u2202x1(k) = \u2202f(q1(k)) \u2202q1(k) . \u2202q1(k) \u2202f(pj(k)) . \u2202f(pj(k)) \u2202pj(k) . \u2202pj(k) \u2202x1(k) (28) Here \u2202pj(k)/\u2202x1(k) = wj1(k) Furthermore, the approximate value of \u2202y(k)/\u2202u(k\u2212d) can be calculated from the linear model as the following equation y\u0304(k) = \u2212a\u03041(k)y(k \u2212 1) \u2212 ... \u2212 a\u0304n(k)y(k \u2212 n) +b\u03040(k)v(k \u2212 d) + .." + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000332_50003-4-Figure3.7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000332_50003-4-Figure3.7-1.png", + "caption": "Fig. 3.7. Side-view of the brush tyre model at braking (no sliding considered).", + "texts": [ + " Further, as expected, the actual tyre generates an aligning torque larger than according to the model. For the brush-type tyre model with tread elements flexible in longitudinal direction, the theory for longitudinal (braking or driving) force generation develops along similar lines as those set out in Section 3.2.1 where the side force and aligning torque response to slip angle has been derived. To simplify the discussion, we restrict ourselves here to non-negative values of the forward speed Vx and of the spe~ of revolution ~2. In Fig.3.7 a side-view of the brush model has been shown. As was indicated before, the so-called slip point S is introduced. This is an imaginary point attached to the wheel rim and is located, at the instant considered, a distance equal to the effective rolling radius r e (defined at free rolling) below the wheel centre. At free rolling, by definition, the slip point S has a velocity equal to zero. 102 THEORY OF STEADY-STATE SLIP FORCE AND MOMENT GENERATION Then, it forms the instantaneous centre of rotation of the wheel rim", + " We may think of a slip circle with radius r e that in case of free rolling rolls perfectly, that is: without sliding, over an imaginary road surface that touches the slip circle in point S. When the wheel is being braked, point S moves forward with the longitudinal slip velocity V~x. When driven, the slip point moves backwards with consequently a negative slip speed. In the model, a point S' is defined that is attached to the base line at its centre (that is, at the base point of the tread element below the wheel centre, cf. Fig.3.7). By definition, the velocity of this point is the same as that of point S. That means that S' also moves with the same slip speed V~x. It is assumed that the tread elements attached at their base points to the circumferentially rigid carcass, enter the contact area in vertical position. At free rolling with slip speed V~x (of both points S and S') equal to zero, the orientation of the elements remains vertical while moving from front to rear through the contact zone. Consequently, no longitudinal force is being transmitted and we have a wheel speed of revolution: V ~ _ g2o_ x (3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002794_5.0049922-Figure10-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002794_5.0049922-Figure10-1.png", + "caption": "FIGURE 10. Bearing elements and 4000 rpm, speed applied on glass-epoxy composite driveshaft.", + "texts": [ + " In the rotodynamic analysis, the finite element model developed in ANSYS ACP (Pre) is shared with the modal analysis, and the bearing elements COMBIN14 are added at the ends of the shaft as stationary parts. The Coriolis Effect from the rotodynamic controls is added which applies the gyroscopic and rotating damping effect to a rotating shaft. In the analysis setting the six modes to extract and five steps of speed for each mode are selected. The translation in all directions and rotation about the lateral directions are fixed at bearing locations and the rotational velocity is applied at one end in the five steps as shown in Table 3. Figure 10, shows the 4000 rpm speed applied at one end and the combin14 bearing element is added to the other end of the drveshaft. Figure 11 shows the six modes of vibration at each speed which has plotted the graph of frequency at each calculated mode to the corresponding speed. In the first six frequency plots at a speed of 4000 rpm, the first natural frequency for this speed is 282.74 Hz and the frequency for the sixth mode is 1345 Hz. The next six modes are at 8000 rpm speed and the natural frequency for the first mode at 8000 rpm is 279" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002862_09544062211016076-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002862_09544062211016076-Figure3-1.png", + "caption": "Figure 3. The prototype of the alignment mechanism: (a) 3D model; (b) the entity.", + "texts": [ + " However, the vibration characteristics of the combined system could be estimated using the inertia of the payload and stiffness of the support according to the preliminary tests. As a result, holistic stiffness of the alignment is of great importance and this subsection would introduce the application of the proposed method to obtain an analytical stiffness model for optimizations. VJM model of the parallel alignment mechanism The specific structure of the alignment mechanism in China\u2019s large space telescope is shown in Figure 3(a) and (b). There are three main parts, namely the base, six legs and payload platform. The base is fixed on the truss structure of the telescope while the payload platform is installed with the optical payload and the legs support the platform. The preliminary tests showed that the first six order mode frequencies that computed using the inertia parameters of the payload and the holistic static stiffness of the alignment mechanism are well consistent with the FEM results, which is considered as the result of the high-stiffness of payload as well as that the movable mass of combined system mainly concentrates on the payload" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000084_icar.2005.1507478-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000084_icar.2005.1507478-Figure2-1.png", + "caption": "Fig. 2. Repulsive forces of various angles.", + "texts": [ + " A platform function is adopted to describe the repulsive forces produced by an obstacle point. The repulsive force at angle by the obstacle at angle is defined as follows: ) )( arcsin()( )(, )( 1 )()(, )()( 1 )(, )( ,0 )()(),( ),( d D Dd vDD DdvD vDd Dd k Otherwise k k sx s m sym msy sy sy p ssp RF where )(d is the distance of obstacle at angle Dm is the maximal estimate distance For a certain angle , the total repulsive force is set as the maximum produced by the obstacles at all angles. The repulsive force function is illustrated in Figure 2 and formulated as follows: )(max)( ],0[ RFRF kK C. Attractive Forces Field The repulsive forces field represents the resisting effects of obstacles to the mobile robot. Otherwise, for leading the robot move to the object location, attractive force produced by the object point should be considered. A cosine function is used to define the attractive forces at various angles: )cos()( objGFK where obj is the direction angle of the object point A cosine function is used because of follow reason: While the angle has a large departure from the object angle, the force varies significantly in order to lead the robot to object direction rapidly; On the other side, while the departure is small, the force varies insignificantly, to avoid that the robot is over attracted by the object point and ignore obstacles" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002403_s40430-021-02907-8-Figure15-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002403_s40430-021-02907-8-Figure15-1.png", + "caption": "Fig. 15 Grasp configurations evaluated based on the validated FE model", + "texts": [ + " Figure\u00a014 shows how a FE model can be exploited to populate a database; is used to discriminate stable and unstable grasps in the database. Given the validated FE model, apart from the external perturbation to the grasp, the grasp configuration, friction coefficient, contact force and mass of the object can be changed to accommodate different grasp scenarios in the FE simulation. In all the cases, can be computed and stored for each combination of the parameters under consideration. As an illustration, consider the grasp scenarios depicted in Fig.\u00a015. Although these configurations are significantly different in real-world, in FE framework, changing the gravity and lift direction in the validated FE model (Fig.\u00a015a) leads to the configuration depicted in Fig.\u00a015b. Table\u00a03 shows the values of obtained for these grasp configurations of HDPE bottle. The mass and friction at the contact interface are changed, and the values show the influence of them on different grasp configurations. It can be seen that at low friction ( =0.2), 1000 g botttle slips, thereby resulting in = 0 . Nevertheless, with increase in friction ( =1), the grasp attains a stable state and values are greater than 0 for both 27 g and 1000 g bottle. Figure\u00a011 shows the evolution of contact area during lifting for the grasp configurations considered" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003385_ur52253.2021.9494662-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003385_ur52253.2021.9494662-Figure6-1.png", + "caption": "Fig. 6. Frame of mobile manipulator", + "texts": [ + " V = [AdX\u22121Xd ]Vd(t) +KpXe(t) +Ki \u222b t 0 Xe(t)dt (9) X is the homogeneous transformation matrix from the reference frame to the end effector according to time. Xd is the homogenous transformation matrix representing the desired motion of the end-effector, where [Vd] = X\u22121 d X\u0307d. The desired motion is input through the feedforward, and compensation for Xerr is given through the PI controller. Xerr is not simply X \u2212 Xd, and is represented as [Xe] = log(X\u22121Xd) instead. Mobile manipulation refers to when the mobile platform and the manipulator are controlled simultaneously [6]. As shown in Fig. 6, the fixed coordinates {s}, the coordinates of the mobile platform {b}, the base coordinates of the manipulator {0}, and the coordinates of the end-effector {e} are designated. The homogeneous transformation matrix from the fixed coordinates {s} to the end-effector is as shown in (10). X = Tse(q, \u03b8) = Tsb(q)Tb0T0e(\u03b8) (10) Je is required for controlling the end-effector using the velocities of the wheels and joints. Je consists of [ Jbase Jarm ] and the velocity of the end-effector is independent of the composition of the mobile platform, thus, it is not dependent on q" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001777_epepemc.2008.4635622-Figure12-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001777_epepemc.2008.4635622-Figure12-1.png", + "caption": "Figure 12. Midvalue of flux density at n = 1500 rpm with an excitation of If = 1760 Aturns.", + "texts": [ + " All simulations deliver an output current larger than the measured Table I OVERVIEW OF THE NUMERICAL APPROACHES CHOSEN TO SIMULATE THE CLAW POLE ALTERNATOR. Simulation approach Comment a) iMOOSE \u2202\u0398\u03d5r The induced voltages are calculated by finite differences similar to postprocessing. b) iMOOSE \u2202Ir M The induced voltages are calculated by the variational approach as proposed in II-C. c) JMAG Commercial software package. Implementa- tion and exact approach unknown. The phase and the output current waveforms simulated by implementation b) are depicted in Fig. 11. The flux density of the modeled claws is illustrated in Fig. 12. The proposed field-circuit coupling method is applicable to simulate complex problems, 2D or 3D, with or without motion. Two approaches to calculate the motion induced voltages have been implemented. Despite the good agreement obtained between simulated and measured currents, the energy based approach is numerically stable. Though the calculation of the torque is crucial for this approach, it appears to be more reliable than the calculation of the electromotive forces by finite differences. Further investigations on the calculation of the torque are expected to give more information about the offset in the simulated mean output current" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002838_012021-Figure17-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002838_012021-Figure17-1.png", + "caption": "Figure 17. Strain Distribution in S-Glass", + "texts": [ + " Analysing Testing Result of S-Glass 3.4.1. Total Deformation The Max. and Min. Total Deformation in S-Glass is 0.18974 mm and 0 mm respectively shown in Figure 15. The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. 3.4.2. Stress distribution The Max. And Min. Stress distribution in S-Glass is 59.887 MPa and 0.3646 MPa respectively shown in Figure 16. 3.4.3. Strain Distribution The Max. and Min. Strain distribution in S-Glass is 0.00069941 and 0.000005478 respectively shown in Figure 17. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.5. Analysing Testing Result of E-Glass 3.5.1. Total Deformation The Max. and Min. Total Deformation in E-Glass is 0.23872 mm and 0 mm respectively shown in Figure 18. The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. 3.5.2. Stress Distribution The Max. and Min. Stress Distribution in E-Glass is 62.053 MPa and 0.37676 MPa respectively shown in Figure 19" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002581_icees51510.2021.9383730-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002581_icees51510.2021.9383730-Figure6-1.png", + "caption": "Fig 6. 5 mm Fin Geometry at the Housing o f the Motor", + "texts": [ + " Thereby increases the rate of heat transfer from surface to adjacent fluid. As a heat sink, the finned casing is been used and the varying impact of the geometry of fin on the functioning of cooling has been examined using three various designs. By introducing rectangular fins, the high temperature on the end windings will be reducing. The temperature of the windings is reduced by increasing the depth of the fins. Geometry of finned casing was generated for two different dimensions as shown in fig 6 and fig 7, But, in this motor application it is not feasible to increase the difference in the temperature, because the temperature in the system is fixed by other constraints. The only option is to increase the area. This is done using the extended surface or fins. Fin dimensions and its efficiency is given in table IV. TABLE IV. Fin Dimensions and its Efficiency. S.No Fin Dimension mL Qfin/Qlongfin = tank mL 1 0.1 0.100 2 0.2 0.197 S.No Fin Dimension mL Qfin/Qlongfin = tank mL 3 0.5 0.462 4 1 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000438_haptic.2006.1627118-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000438_haptic.2006.1627118-Figure4-1.png", + "caption": "Figure 4: (a) Cross-section of a rigid convex object and a face of an elastic object before contact (b) Deformed shape of the elastic object after contact", + "texts": [ + " the distance from the point to the contact point and proportional to the pressing force magnitude at the contact point. The deformed shape inside the local neighborhood is convex when viewed from the rigid object. Proof: Consider deformation of the elastic object under a single point contact caused by a compressive load F applied normal to a face S of the elastic object. Let W be a a cross-section of the elastic object normal to the face of the elastic object in contact and contains the contact point as shown in Fig. 4a, where the curve of the contacting face of the elastic object inside the local neighborhood of contact with radius d can be viewed as a straight-line segment. Let FW be the portion of force F on W . Let r \u2264 d be a distance variable inside the local neighborhood. Let \u03b8 be the angle between r and the extension of the line of the load FW , \u03c3rr and \u03c3r\u03b8 be stress components. Thus, \u222b +\u03c0/2 \u2212\u03c0/2 (\u2212\u03c3rr cos\u03b8 +\u03c3r\u03b8 sin\u03b8 )rd\u03b8 = FW (1) When \u03b8 = \u00b1\u03c0/2, we can get the following according to [5] ur = \u2212 (1\u2212\u03c5\u2217)FW 2E\u2217 (2) u\u03b8 = \u2213( 2FW E\u2217\u03c0 ln( d r )\u2212 (1+\u03c5\u2217)FW E\u2217\u03c0 ) (3) Here ur and u\u03b8 are the r- and \u03b8 - components of the deformation displacement inside the local neighborhood of contact on the crosssection", + " 3 denotes the logarithmic function shaped deformation, and the deformation displacement is proportional to the force magnitude. Such results apply to any crosssection W because the material of the elastic object is homogeneous isotropic. Therefore, the deformed shape inside the local neighborhood along the face S is also logarithmic and proportional to the applied force. The logarithmic shape of deformation forms convex boundary features inside the local neighborhood region of the elastic object. The total deformation effect is shown in Fig. 4b. Now we consider properties related to global deformation (as defined in Section 2.3). Theorem 3.5 (Smoothness) - For any non-vertex boundary feature of an elastic object, when a point contact happens, its segment outside the local neighborhood of the contact point will keep its smoothness. Proof: Since no external force directly acts on the segment not in contact, its deformation displacement is only caused by stresses spread to it. From elastic mechanics, we have the relationship between the strain \u0393 and the stress T as \u0393 = 1 E [(1+\u03bd)T\u2212\u03bd\u0398I] (4) Here \u0398 = \u03c311 +\u03c322 +\u03c333, \u03c311 , \u03c322 and \u03c333 are stress components where the subscripts 1, 2, and 3 represent three coordinate axes respectively, and I is a unit matrix" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002804_j.matpr.2021.04.568-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002804_j.matpr.2021.04.568-Figure7-1.png", + "caption": "Fig. 7. Fatigue analysis of Carbon/Epoxy.", + "texts": [], + "surrounding_texts": [ + "CAD models are designed in CATIA which contains special tools for the generation of traditional surfaces which will then be transformed into solid models, using traditional and composite monoleaf spring materials. The measurements of a spring leaf of a TATA SUMO vehicle are used for the design of the mono-leaf spring. Table1 displays the mono leaf spring configuration parameters and the planned mono leaf spring CAD model, as shown in Fig. 1. Finite Element Analysis (FEA) is a computational method for deconstruction into very small elements of a complex structure. In the simulated world, ANSYS offers an affordable way to examine the success of goods or processes. Digital prototyping is called this method of product growth. Users will iterate different scenarios using simulated prototyping techniques to refine the software even before the production is launched. This allows the probability and expense of failed designs to be reduced. In this study, a model was developed that was imported into the ANSYS workstation and FEA analysis. Meshing is the mechanism where the entity is divided into very small pieces called components. Elements. It is sometimes called a piece by piece. The leaf spring model is here meshed with a 10 mm brick mesh part scale. The front end is restricted and the rear end only in Y and Z is restricted; translational movement is permitted in X direction. Conditions of loading include applying a force on the middle of the spring of the leaf upward on the base of the leaf spring. The loading range is between 1000 N and 5000 N. The meshed model and limit and conditions of loading of a leaf spring are seen in Figs. 2 and 3." + ] + }, + { + "image_filename": "designv11_83_0000412_j.ijrmms.2003.12.052-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000412_j.ijrmms.2003.12.052-Figure1-1.png", + "caption": "Fig. 1. Diagram of stress\u2013s", + "texts": [ + " Ultimate axial strain is in turn dominated by the flaws contained in the specimen. The failure process is caused by fracture development initiated at flaws. 2. Strain rate-dependent ultimate strength reduction is accounted for by ultimate axial strain reduction and secant modulus reduction. The secant modulus reduction reflects a strain component due to the strain rate effect. The strain component caused by the strain rate effect is statistically and mathematically determinable. It can be approximately determined using the new method suggested in this study. Fig. 1 is an illustration of the method. For the tested welded tuff, the strain rate effect strain component increases from 5.1% to 16.5% of the total axial strain when the strain rate decreases from 10 4 to 10 8 s 1. ng author. ss: lumin@unr.nevada.edu (L. Ma). th paper see CD-ROM attached. ms.2003.12.052 train decomposition." + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000790_s1560354708040072-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000790_s1560354708040072-Figure3-1.png", + "caption": "Fig. 3. When a hexagonal pencil is at rest on a plane whose inclination is \u03b8 < \u03c0/6, the center of mass of the pencil is at height a cos[\u03c0/2 \u2212 (\u03b8 + \u03c0/3)] = a cos(\u03c0/6 \u2212 \u03b8) above its lowest point.", + "texts": [ + " However, we will find a slightly more restrictive limit on \u03b8min in Section 2.2. Combining Eqs. (8), (9) and (11) we estimate the asymptotic linear velocity to be \u3008v\u221e\u3009 = 3 \u03c0 a \u3008\u03c9\u221e\u3009 \u2248 3 \u03c0 \u221a ag k \u221a 1 + \u03b5 1 \u2212 \u03b5 sin \u03b8 \u2212 0.178 cos \u03b8, (13) which vanishes for \u03b8 = \u03b8min. 2.2. \u03b8 < \u03c0/6 In this case the pencil will not roll unless it is given an initial kinetic energy E1 > mga[1 \u2212 cos(\u03c0/6 \u2212 \u03b8)], (14) such that the center of mass of the pencil can rise to the vertical during the first 1/6 turn, as shown in Fig. 3. The pencil will not continue to roll through a second 1/6 turn unless sufficient energy remains after its collision with the plane. That is, we need the energy at the beginning of the second 1/6 turn to satisfy E2 = \u03b5(E1 + \u0394E) > mga[1 \u2212 cos(\u03c0/6 \u2212 \u03b8)]. (15) Similarly, the energy at the beginning of the nth 1/6 turn must satisfy En = \u03b5n\u22121E1 + \u03b5\u0394E 1 \u2212 \u03b5n\u22121 1 \u2212 \u03b5 > mga[1 \u2212 cos(\u03c0/6 \u2212 \u03b8)], (16) recalling the argument that led to Eq. (5). The asymptotic condition is that \u03b5 1 \u2212 \u03b5 \u0394E = \u03b5 1 \u2212 \u03b5 mga sin \u03b8 > mga[1 \u2212 cos(\u03c0/6 \u2212 \u03b8)], (17) REGULAR AND CHAOTIC DYNAMICS Vol" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000851_imece2007-41027-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000851_imece2007-41027-Figure4-1.png", + "caption": "Fig. 4: Visualization of test circuit [5]", + "texts": [ + " By reason of test time shortening it is necessary to select larger torque in closedloop circuit than in industrial operation. In our case, we are limited by torque sensor in circuit up to 5000 Nm. In the course of 1500 RPM the circuit is dimensioned for maximal virtual power 785 kW. Testing is mostly running on one load level because of better possibility of result comparing. The whole test-rig with PLC, control panel, converter and hydraulic devices is shown in figure 3. Developed back-to-back circuit is visualized in figure 4, its schema is in figure 5 and descriptions are in following paragraphs. Loading equipment must ensure easy creation of torque in the circuit. Creation of torque is realized by axial movement of the gearwheel (Fig. 5, Pos. 3) with a helical gear in mesh with the pinion in the additional gearbox (Fig. 5, Pos. 14). This system is similar to NASA Glenn Research Center Spiral Bevel Gear Facility [4]. The tensioning screw gives rise to axial force that causes reaction (tangential force) in the gearing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000092_14689360600734112-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000092_14689360600734112-Figure2-1.png", + "caption": "Figure 2. The velocity vector behaves like a nonlinear pendulum.", + "texts": [ + "0 0 and such that RC2 \u00bc !2 0 cos and RC1 \u00bc !2 0 sin . Then \u20ac\u2019 \u00bc !2 0 cos sin\u2019 sin cos\u2019\u00f0 \u00de \u00bc !2 0 sin \u2019 \u00f0 \u00de, and defining \u00bc \u2019 , we obtain \u20ac \u00bc !2 0 sin , \u00f032\u00de D ow nl oa de d by [ U ni ve rs ity o f C on ne ct ic ut ] at 0 8: 00 0 9 O ct ob er 2 01 4 which is the equation for the pendulum, with !2 0 representing the ratio g/l of the acceleration of gravity to the length of the wire of the pendulum. The relationship between the optimal control for the plate\u2013ball system and the dynamics of a pendulum is illustrated in figure 2, where we have set \u00bc 0 for simplicity. The solutions to (32) have been widely studied and involve Jacobi elliptic functions. Having thus shown that the angle \u2019 oscillates like a pendulum with respect to some fixed orientation on the table, it is easy to see that every solution to (32) gives rise to a family of solutions to the plate\u2013ball isoparallel problem. If (t) is a solution to (32) for a given !0, then choose any R>0 and any , and define _x \u00bc R cos\u00f0 \u00fe \u00de, sin\u00f0 \u00fe \u00de, 0\u00f0 \u00de v \u00bc R sin\u00f0 \u00fe \u00de !2 0 sin R , R cos\u00f0 \u00fe \u00de \u00fe " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002324_012099-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002324_012099-Figure5-1.png", + "caption": "Figure 5. Schematic diagram of the experimental device.", + "texts": [ + " The purpose of character recognition is achieved through character comparison between the segmented characters and the classifiers learned and trained from a large number of images. In order to complete the on-line detection of wire bundles, the detection experiment is designed according to the principle and method described above. The identification system includes an industrial CCD camera, a light source, a wire beam clamp, a full-angle reflector, a wire beam, a lifting platform and an optical platform. The experimental device is shown in figure. 5. Design and selection of the line beam recognition device: the center of the camera, lens, light source and total reflector is located on the same line to form a coaxial optical system. Because the full-view mirror and cable are black and white, the working wavelength of the system can be defined as visible light region, so the camera can choose black and white camera. Since the cross-section of the total reflector is a 92*40mm rectangular surface, and the target surface of the sensor is usually a 4:3 rectangle, ISPECE 2020 Journal of Physics: Conference Series 1754 (2021) 012099 IOP Publishing doi:10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000143_6.2006-6241-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000143_6.2006-6241-Figure1-1.png", + "caption": "Figure 1. The vertical take off and landing aircraft (VTOL) model.", + "texts": [ + " Furthermore, simulation results are also presented to demonstrate that the preview-caused error in the inverse input exponentially decays with the increase of preview time\u2014with the decay rate related to the unstable poles of the linearized internal dynamics of the aircraft. II. Preview-Based Stable-Inversion In this section, we illustrate the implementation of the preview-based stable-inversion technique to achieve precision trajectory tracking, by using a simplified nonlinear model of VTOL aircraft shown in Fig. 1.7, 8, 23 We start by describing the VTOL aircraft dynamics and the inversion-based control scheme. Model of VTOL Aircraft We consider the following simplified VTOL aircraft model (see, e.g., Ref. 7,8,23) [ x\u0308 z\u0308 ] = [ \u2212 sin \u03b8 \u03b5 cos \u03b8 cos \u03b8 \u03b5 sin \u03b8 ][ u1 u2 ] + [ 0 g ] , M\u03b8 [ u1 u2 ] + [ 0 g ] , \u03b8\u0308 = \u03bbu2 (1) y = [y1 y2] T = [x z]T (2) where x and z are the horizontal and vertical position of the aircraft mass center, respectively, \u03b8 is the roll angle and g is the gravitational acceleration. The inputs u1 and u2 are the thrust and the rolling moment, respectively, and \u03b5 describes the coupling effect between the rolling moment and the dynamics in the lateral and vertical directions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003341_s11665-021-05551-4-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003341_s11665-021-05551-4-Figure2-1.png", + "caption": "Fig. 2 The dynamic model of Ti-6Al-4V alloy sample", + "texts": [ + " In order to improve calculation efficiency, a 1/4 dynamic simulation model was established based on the symmetry of laser action, and a symmetric boundary condition was applied to the corresponding symmetry plane. That is, axial constraints were applied to the two symmetric planes, and full constraints were applied to all nodes of the bottom surface in this model. Where, the size of the model was 5 mm 9 5 mm 9 3 mm, a size of 0.05 mm was used to the impact zones, and a size of 0.2 mm was adopted for the part outside the impact zones. Figure 2 shows the dynamic model of Ti-6Al-4V alloy sample. Different laser parameters have different influences on the degree of plastic deformation of the material s surface. However, technical limitations of current laser generating devices limit the adjustable range of various parameters of pulsed laser. To elucidate the impact effects with different conditions for Ti-6Al-4V alloy sample during LSP, the Journal of Materials Engineering and Performance Volume 30(8) August 2021\u20145517 following will investigate the effects of nanosecond pulsed laser-generated processing Ti-6Al-4V alloy sample based on different conditions by numerical simulations and experiments" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000950_eej.20456-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000950_eej.20456-Figure2-1.png", + "caption": "Fig. 2. Two-link arm.", + "texts": [ + " But there are no simulations that take account of elements of a real robot\u2019s dynamics such as friction, gravity, or system delays. In this study, prior to implementing FPL on a real robot, we performed learning simulation experiments to confirm appropriate learning of the inverse dynamics of the controlled object including the above dynamical elements. 3.1 Controlled object and learning conditions Here we consider the forelegs of AIBO as the controlled object, and thus we deal with the two-link arm shown in Fig. 2. The coordinates express the sagittal plane, with the x-axis and y-axis pertaining, respectively, to the back-forth and up-down directions. The joints and links are denoted as Joint 1, Joint 2, Link 1, Link 2, starting from the origin. The dynamics of this two-link arm is described as follows: Here t = [\u03c41, \u03c42] T (Nm) is the joint driving torque, q = [\u03b81, \u03b82] T (rad) is the joint angle, M(q) is an inertial term, h(q, q . ) is a term expressing the centrifugal and Coriolis forces, Bq . is a viscous friction term, and g(q) is a gravity term" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000739_j.precisioneng.2007.01.001-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000739_j.precisioneng.2007.01.001-Figure1-1.png", + "caption": "Fig. 1. Elgiloy \u2018clock\u2019 spri", + "texts": [ + " Literature s lacking on studies of creep of elgiloy springs under load in he spiral configuration. The primary concern here for applicaion of elgiloy springs is characterization of creep rates so that ffective service lives can be estimated during which the springs rovide the required amount of load to the given mechanisms or proper operation. The configuration of elgiloy springs conidered here is that of a wind-up torque spring that must provide iven amounts of torque for specified degrees of rotation, as seen n Fig. 1. The springs are made of flat wire that is 0.1 mm thick \u2217 Corresponding author. Tel.: +1 7046878228. E-mail address: kclynn@uncc.edu (K.C. Lynn). s d a w M t m 141-6359/$ \u2013 see front matter \u00a9 2007 Elsevier Inc. All rights reserved. oi:10.1016/j.precisioneng.2007.01.001 nd 1.0 mm in width, and has 85% cold-work reduction with ssociated yield strength of approximately 230 ksi [6]. Approxmately eight coils of the spring are wound into a housing that as an inner pocket with a diameter of approximately 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002801_s00006-021-01119-6-Figure15-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002801_s00006-021-01119-6-Figure15-1.png", + "caption": "Figure 15. The post-transformed Delta Manipulator, where the values of \u0394H1,6,11, \u03b31,6,11 and \u03b4m5,10,15 are shown in Table 7. The blue and red dotted lines indicate the preand post- transformed joint axis vectors si and s \u2032 i, respectively", + "texts": [ + "901 radians, respectively, about the x-axis and rotated by 0.165, 1.047 and -1.047 radians, respectively, about the z-axis (Eq. 29). Finally, \u03b4m5,10,15 = 0 in Table 11 indicates that there should be no change in the end-effector mass, to produce the desired {1,6,11}(R\u03b4i(\u03b1i, \u03b3i)) and {1,6,11}(R\u03b41,6,11 (\u03b11 = \u22120.756, \u03b16 = 1.047, \u03b111 = 0.901, \u03b31 = 0.165, \u03b36 = 1.047, P 1c {5,10,15}(\u03b4m5,10,15) = 0 end-effector external wrench (Q\u0302 \u2032 j) while maintaining \u03c4 \u2032 l \u2212 \u03c4l = 0. The posttransformed Delta manipulator is shown in Fig. 15. Using the solutions of the Difference Lagrangian system of equations, as shown in Table 11, we arrive at the specific reassembling transformation: P 1a {1,6,11}(\u0394H1 = 0.047,\u0394H6 = \u22120.015,\u0394H11 = 0.098), P 1b1 {1,6,11}(R\u03b41,6,11(\u03b11 = \u22120.756, \u03b16 = 1.047, \u03b111 = 0.901, \u03b31 = 0.165, \u03b36 = 1.047, \u03b311 = \u22121.047)), P 1c {5,10,15} (\u03b4m5,10,15) = 0. Having applied this, we produce the post-transformed torques presented in Fig. 29, while achieving the same end-effector trajectory, velocity, acceleration and end-effector external wrench, as shown in Figs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003169_s00202-021-01352-z-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003169_s00202-021-01352-z-Figure4-1.png", + "caption": "Fig. 4 A 2D schematic of the bridgeless STPMIG", + "texts": [ + " Jor is the PM outer rotor inertia, and Jir is the squirrel-cage inner rotor inertia. T ,or and T ,ir are the mechanical torque of the outer rotor and the mechanical torque of the inner rotor, respectively. As mentioned, the proposed generator adopts 34 poles, and the stator winding is regarded based on the mentioned combination of poles. The characteristics of the proposed generator are presented in Table\u00a01 for both the bridged and bridgeless structures. The 2D structure of the bridgeless STPMIG is depicted in Fig.\u00a04. In this paper, finite element analysis (2D-FEA) is employed to analyze the steady-state and transient performance of the proposed generator. The eddy current effect of the cores and permanent magnets is neglected to omit the detrimental effect on the back-EMF. Furthermore, nonlinear B-H characteristics of the cores are taken into account. Therefore, the design characteristics and principal (10) T ,or + Te,ir \u2212 Te,s = T ,or + 3p 4 ( miiqr \u2212 moiqs) = Jor d or dt (11)T ,ir \u2212 Te,ir = T ,ir \u2212 3p 4 ( driqr \u2212 qridr) = Jir d ir dt parameters of the proposed generator are realized by exploiting the simulation software of finite element analysis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003042_s11668-021-01191-x-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003042_s11668-021-01191-x-Figure6-1.png", + "caption": "Fig. 6 aDisplacement in Modification 1, b displacement in Modification 2, c displacement in Modification 3 and d displacement in Modification 4", + "texts": [], + "surrounding_texts": [ + "The 3D modeling of EGR cooler system is meshed using the Hypermesh software using the solid 3D, four node and tetrahedron elements with three degrees of freedom. The tetrahedral element meshing of the EGR cooler housing is shown in Fig. 3. Tetrahedral elements are used here because they fit arbitrary shaped geometries very well with their simple computations. In comparison, the hexahedral meshes are more accurate with the number of elements, since one hexahedral equal to six tetrahedral elements. However, the tetrahedral elements are best to model complex geometry domain with little distortion of mesh. While meshing the model the conditions given are linear elastic, isotropic and temperature independent. The material properties inputted are listed in Table 1 for aluminum for housing, cast iron for bracket and steel for the bolt. The chemical composition of the aluminum used is listed in Table 2. The bolts are pre-tensioned at the load of 25735N. After giving the DOF and Load, Nastran solver is used to solve the problem." + ] + }, + { + "image_filename": "designv11_83_0002838_012021-Figure29-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002838_012021-Figure29-1.png", + "caption": "Figure 29. Strain Distribution in HSLA Steel", + "texts": [ + " Analysing Testing Result of HSLA Steel 3.8.1. Total Deformation The Max. And Min. Total Deformation in HSLA Steel is 0.19371 mm and 0 mm respectively shown in Figure 27. The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. 3.8.2. Stress Distribution The Max. And Min. Stress Distribution in HSLA Steel is 181.96 MPa and 1.2159 MPa respectively shown in Figure 28. 3.8.3. Strain Distribution The Max. And Min. Strain Distribution in HSLA Steel is 0.0011085 and 0.00000815 respectively shown in Figure 29. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 The Analysis report of maximum total deformation, maximum equivalent stress and maximum equivalent strain is presented below in bar graphs: 4.1. Maximum Total Deformation: Maximum Total Deformation is shown in carbon fiber i.e. 0.68523 mm and it is shown at the upper flange part of the rim. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003132_978-981-16-0119-4_20-Figure78.2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003132_978-981-16-0119-4_20-Figure78.2-1.png", + "caption": "Fig. 78.2 Conceptual design of propeller with saw tooth cut at both LEs", + "texts": [], + "surrounding_texts": [ + "enhancement technique and its positions on the propeller for this current work [9]. While coming to the methodology section, the two engineering approaches have beenused to test the aerodynamic behavior on the relevant components. The methodologies are experimental test andComputational FluidDynamics (CFD)based simulation, in which most of the aerodynamic analyses were executed with the help of CFD simulations. Because of these huge implementations, the pre-processor steps are available to everyone [11]. Especially, the boundary conditions such as type of solver, type of turbulence model, quantity and quality of turbulence model are easily available data to the new researchers. The high amounts of analyses were used velocity inlet to their problems. Thus, with these inputs, the CFD simulations are executed in these comprehensive aerodynamic analyses on various propellers [13]. b. Solution Techniques \u2013 CFD Analysis In this work, analyses the comparative aerodynamic performance on a UAV\u2019s propeller by using CFD tool, i.e., Ansys fluent. The fundamental aim of this work is to select the suitable lift enhancement technique for UAV\u2019s propeller. In this regard, six different designs are modeled, in which five conceptual designs are comprised of a propeller with edge modifications, and the other one is the conceptual design of base propeller. The techniques implemented in the 5-inch diameter propeller are curvy cut saw tooth cut, aero cut, etc., and in general, the profile modifications in the UAV\u2019s propeller is executed for noise reduction. From the literature survey, it was clearly understood that the noise induced due to the abnormal environment is reduced [1, 2]. Nowadays, UAV industry needs a quite propeller so the design modifications-based propellers are suggested a lot for the construction of quite UAV. But the problem along with these types of profile modified propellers may have a chance to generate low-aerodynamic forces. Hence, the conduction of an integrated study is very important in the updated propellers to increase the implementation of the UAVs in real-time applications. The steady- and pressure-based turbulent flow is used as fundamental behavior to the working fluid for all these analyses. The aerodynamic parameters such as lift, drag, CL, CD are used as selection parameters for this comparative analysis [15, 16and17]. c. Conceptual Design of various propellers UAV\u2019s propeller and its designs are the key role of this aerodynamic performance investigation. From the previous work, it was understood that propeller with saw tooth cut has been provided the low turbulence noise than base propellers. Thus, in this work, the aerodynamic performance of low acoustic profiled propellers is computed, in which the conceptual design of all the propellers are used from the literature survey [18, 19,and20]. Figures 78.1, 78.2, 78.3, 78.4, and 78.5 are revealed the conceptual design of propellers with saw tooth cuts, in which the locations of the saw tooth cuts are formed at the various edges of the propellers such as both leading edges, both trailing edges, one leading edge cum one trailing edge, only one trailing edge. Apart from these saw tooth cuts, the one more relevant cut is located at the leading edges of the propeller, which is v-cut. 78 Comparative Aerodynamic Performance Analysis on Modified \u2026 973" + ] + }, + { + "image_filename": "designv11_83_0003132_978-981-16-0119-4_20-Figure76.5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003132_978-981-16-0119-4_20-Figure76.5-1.png", + "caption": "Fig. 76.5 Finite element model of the operator platform", + "texts": [ + "The swivel seat is placed above theoperator platform, and the steering wheel mounted on the front console is also fixed to the operator platform. The vibrations transferred from the front chassis to the operator platform and further to the seat and the steering wheel depend upon the stiffness of AVM along three directions. The stiffness of AVMs decides the natural frequency of the operator platform which needs to be away from the machine operating frequencies and engine low idle and high idle speeds. The finite element model of the operator platform is shown in Fig. 76.5, and FEA is carried out using commercial analysis software, ANSYS. Details of the FE model like element, real constant and material properties are shown in Table 76.2. Modal analysis is carried out to find out the natural frequencies and corresponding mode shapes of the operator platform. FE model is first solved using the existing AVM stiffness values. The natural frequencies of the operator platform are also listed in Table 76.4. It is evident from Table 76.4 that a few of the natural frequencies (highlighted) are closer to the machine operating frequencies as well as engine idle speeds (shown in Table 76" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001413_j.jsv.2008.02.056-Figure9-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001413_j.jsv.2008.02.056-Figure9-1.png", + "caption": "Fig. 9. Axial forces on ring.", + "texts": [ + " (57b) Hence, TT p \u00bc V T p \u00f0R2 R3\u00de HT p \u00f0R3 R\u00decot a \u00bc pp\u00bd\u00f0R2 2 R2 1\u00de\u00f0R2 R3\u00de h2 \u00f0R1 \u00fe R2\u00de\u00f0R3 R\u00de=\u00f0R2 R1\u00de and Tp \u00bc TT p =2pR; resulting in Tp \u00bc p\u00bd\u00f0R2 2 R2 1\u00de\u00f0R2 R3\u00de 2R ph2 \u00f0R1 \u00fe R2\u00de\u00f0R3 R\u00de 2R\u00f0R2 R1\u00de . (58) In addition to pressure loadings on the shell structure, which have been considered previously, we make provision for the application of loads in the direction of the axis of symmetry. Thus, we require an analysis of the deformation of each individual ring element under these axial loadings. The loadings to be considered for any ring element are shown in Fig. 9. These loadings induce a torque TL on the ring, where TL is the torque per unit length of circumference measured at the centroidal axis of the ring element. It may be seen from Fig. 9 that the total circumferential load applied at the location of R1 is LT \u00bc 2pR1L. The total circumferential load applied at the location of R2 is also LT \u00bc 2pR2\u00f0LR1=R2\u00de. These two opposing and equal longitudinal forces develop the torque TT L \u00bc 2pR1L\u00f0R2 R1\u00de and a value of TL \u00bc LR1\u00f0R2 R1\u00de=R. (59) With axial restraint provided at the edge of the ring element associated with the radius R2, we obtain the torque due to combined pressure loadings and axial loadings by combining the results from Eqs. (58) and ARTICLE IN PRESS T" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000209_2005-01-2405-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000209_2005-01-2405-Figure1-1.png", + "caption": "Fig. 1 The Basic Scheme of the automatic transmission", + "texts": [], + "surrounding_texts": [ + "For the purpose of enhancing the fuel economy in compact cars, it has been tried to extend the lock-up rpm of damper clutch in automatic transmission. But this way is likely to cause the interior booming noise accompanied by the body vibration. This is why lowering the engine rpm, where the damper clutch starts to be locked up, makes the powertrain vibration increase due to the shortage of the engine power compared to the vehicle load. Therefore, this study will examine the causes and paths of the booming noise and body vibration caused by extending the damper clutch lock-up rpm to lower region. In other words, when the damper clutch runs in lockedup state, this study will examine the increment of the powertrain vibration, the transfer paths of the powertrain vibration and the resonating body system. Specially, this study will use RMA(Running Mode Analysis) method. Through investigation, this study will propose a few kinds of structural modifications and among these modifications the optimized modification will be adopted as the refinement of the booming noise and the body vibration." + ] + }, + { + "image_filename": "designv11_83_0002540_reepe51337.2021.9388090-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002540_reepe51337.2021.9388090-Figure3-1.png", + "caption": "Fig. 3. Section of wire with m conductors removed from the surface.", + "texts": [], + "surrounding_texts": [ + "In laboratory conditions, experiments to determine the fatigue life of materials are carried out at a fixed value of the stress amplitude and a constant frequency of a periodic external action. Real operating conditions differ from laboratory ones in that the impacts are non-periodic. The spectrum of vibrations consists of various frequencies from fractions of a Hertz to several hundred Hertz. The time dependence of fluctuations has the form of a random function. A significant difference between real conditions and laboratory conditions is the presence of tensile stresses. These tensions are also non-periodic. The magnitude of the stresses is unevenly distributed between the central steel part of the cross-section and the aluminum conductors. Thus, in reality, there is a complex stress state of the overhead line wire. Therefore, the destruction of the wire and its breakage can occur in a certain area of it near the attachment to the insulator. In figure 1 shows a fragment of a wire with a break in individual cores [12]. Figure 2 shows a photograph of the state of the wire before its complete break.\nIII. CHANGING THE MICROSTRUCTURE OF THE MATERIAL\nUnder the influence of such irregular power loads, each wire is subject to bending and torsional oscillations. The greatest amount of deformation reaches at the place of suspension of the wire. Torsion deformation occurs due to the asymmetry of ice deposits, which, under the influence of gravity, are able to rotate the wire relative to its axis. The long duration of this kind of exposure allows us to talk about the multi-cycle fatigue of the wire material. The stress state of the wire is represented by the superposition of stresses in the clamps or suspensions, variable stresses caused by the oscillations of the wire, as well as static tension due to the total tension of the wire during its installation and the gravity of the span. Thus, during the operation of overhead line wires, a complex stress-strain state occurs in them. To it is added interaction with the ice shell in winter. Deformations of the shell can cause cracking and destruction processes in it. The rigid adhesion of the ice layer to the surface of the conductor activates similar processes in the surface layers of the conductor.\nMulti-cycle effects on the wire material cause changes in its microstructure, affecting its strength characteristics. In the process of cracking, a surface layer of polycrystalline metal plays a special role. There are many defects in the crystal structure, such as vacancies, dislocations, micropores and their clusters. By combining, they lead to the appearance of microcracks [13]. This occurs in the cyclic loading diagram area up to the start line of irreversible damage (French line). Experimentally observed thickness of surface layer is either grain size or value of order of ten microns in CCC polycrystals [14]. The thickness of such a layer in aluminum is about 30 \u03bcm [15]. Earlier plastic flow of thin surface layers is explained by the following factors [14]. 1. The peculiarity of fixation of surface sources of dislocations with the critical stress of the beginning of work is significantly lower than in volume. 2. The presence in the surface layer is coarser than in the volume of the Frank dislocation grid with a lower dislocation generation voltage. 3. Presence of surface stress concentrators. 4. Different speed of dislocation movement at the surface and inside the crystal. The area of the diagram between the French line and the fatigue curve corresponds to the stage of crack propagation.\nThe actual wires used for overhead lines are a twist of several separate cylindrical conductors, reinforced by a central steel core. Conductors located on the twist surface experience the greatest deformation. The probability of microcracks is also greatest. It should be noted the peculiarity of cyclic\nAuthorized licensed use limited to: California State University Fresno. Downloaded on July 01,2021 at 18:19:08 UTC from IEEE Xplore. Restrictions apply.", + "loading in real operating conditions of overhead lines, which is sharply different from laboratory conditions. Periods of operation of wind loads of different intensity are replaced by periods of their complete absence. Therefore, the accumulation of the number of cycles is uneven, and the deformations of the wire have different values. Another feature of the aluminum wire fatigue curve is the absence of a physical endurance limit. This means that there is no stress boundary value below which fatigue failure does not occur.\nWe present the following scheme of formation and development of fatigue crack. Pores located at the grain boundaries in the conductor metal are microscopically small in size during the initial period of operation and do not pose any danger. Under the influence of variable stresses of tension and compression, periodic sources of vacancies act on sections of borders adjacent to the pore. Such vacancies diffuse from boundaries to pores and in the opposite direction. In general, such processes do not lead to pore growth. If the total tensile stress applied to the line wire is taken into account, the number of vacancies generated will exceed the number absorbed. For this reason, their flows to the pores will exceed the flows in the reverse direction. This leads to an increase in pore volume. Upon reaching a sufficient size, the grain boundary pore becomes a fatigue crack.\nIV. A METHOD OF MONITORING FATIGUE DAMAGE\nThe amount of electrical resistance is very sensitive to the concentration of point defects, in particular vacancies, as well as micro-densities, micropores and microcracks. This makes it possible to monitor the state of metal damage by measuring the electrical resistance of the damaged layers.\nSince the occurrence of cracks occurs on or near the surface of the wire in a thin layer, it is necessary to measure the electrical resistance of this layer. This can be done if such measurements are made by passing a high frequency current through a wire. It is known that high frequency currents flow in the near-surface layer of the conductor due to the surface effect. The thickness of the layer can be varied by varying the frequency of the current. The electrical resistance of the layer is sensitive to the presence and concentration of defects in the crystal structure in German. This method of monitoring the condition of the wire can be even more effective when individual conductors break in the twist during operation. Conductors break in the zone of greatest bending deformations, i.e. on the surface. In this case, the crosssectional area of the high-frequency current conducting tube is reduced.\nThe effective thickness of the conductive surface layer is defined by\n\u03c9\u00b5\n\u03c1 =c . (1)\nHere \u03c9 is the cyclic frequency, \u03bc is the absolute magnetic permeability of the wire material, \u03c1 is its resistivity. By taking the resistivity of the aluminum wire to be \u03c1 = 0.028 \u03bc\u03a9\u2022m, and the above surface layer thickness c = 30 mc m, an estimate of the current frequency for measuring the resistance \u03c9 \u2248 2.5\u2022107 Hz can be obtained from (1). Value of electrical resistance of surface layer\nDc\nl R\n\u03c0\n\u03c1 = taking into account (1) is defined by\n\u03c9\u03c1\u00b5 \u03c0 = D\nl R , (2)\nwhere D is the diameter of the overhead line wire, l is the length of the section on which the resistance is measured.\nIn case of breaking of m conductors of thickness d in twist located on surface, conductivity of all wire decreases. To estimate the effect, assume that as a result, the perimeter of the section P. If we take as n the total number of conductors in the surface layer of the original wire, then the total length of the arcs of large and small circles according to Figure 1 is\n( ) \u2212\u03c0=\u2212\u03c0+ \u2212\u03c0= n md D n m dD n m DP 21 . (3)\nTaking into account (3), the expression (2) takes the form\n( ) \u03c9\u03c1\u00b5 \u2212\u03c0\n\u03b2 =\nndmD\nl R . (4)\nThe geometric coefficient 0 < \u03b2 < 1 takes into account the loose abutment of the broken strands to the remaining part, their folding. This factor cannot be accurately calculated. In (4), two side bridges between arcs of large and small circles of height d are not taken into account due to the fact that in the presence of a surface effect, electric current passes through the layers as far away from the center as possible.\nTheoretical calculation of specific electrical resistance depending on the concentration of defects cannot currently provide sufficient accuracy. Therefore, it is necessary to be able to compare this value with the resistance of a wire not subject to fatigue loading of the same length. For this purpose, you can use sections of wire that are far from the suspension place, and use the bridge method for comparison.\nAuthorized licensed use limited to: California State University Fresno. Downloaded on July 01,2021 at 18:19:08 UTC from IEEE Xplore. Restrictions apply.", + "V. CONCLUSIONS\nThe described physical phenomena can be taken into account when developing methods for monitoring and diagnosing the technical condition of overhead lines. The effects underlying the operation of monitoring devices can be taken into account when creating stand-alone systems that allow monitoring the degree of fatigue damage to a wire. Each device can give a timely signal to the dispatcher about the possible upcoming breakage of individual cores or the entire wire. Such systems will significantly improve the reliability of the overhead line.\nThe research is funded by Russian Federation public contract FSWF-2020-0025 \u00abTechnique development and method analysis for ensuring power system object security and competitiveness based on the digital technologies\u00bb.\n[1] D. P. Poli, Pelacchi, G. Lutzemberger,T. Scirocco, F. Bassi, and G. Bruno, \u201cThe possible impact of weather uncertainty on the Dynamic Thermal Rating of transmission power lines: A Monte Carlo error-based approach\u201d, Electric Power Systems Research, vol. 170, pp. 338-347 2019.\n[2] D.A. Yaroslavsky, M.F. Sadykov, A.B. Konov, D.A. Ivanov, M.P. Goryachev, and T.G. Yambaeva, \u201cMethods for monitoring ice deposits on overhead lines taking into account misalignment of line reinforcement\u201d, Izvestiya vysshikh educational institutions. Energy problems. vol. 19, 2017, pp. 89-97\n[3] \u201cConstruction climatology SP\u201d 131.13330.2018.\n[4] L.M. Keselman, \u201cFundamentals of the mechanics of overhead power lines\u201d, Moscow, Energoatomizdat, 1992, pp. 352.\n[5] P. Hung, H. Yamaguchi, M. Isozaki, and J. Gull, \u201cLarge amplitude vibrations of long-span transmission lines with bundled conductors in gusty wind\u201d, Journal of Wind Engineering and Industrial Aerodynamics. vol. 126. 2014, pp. 48-59.\n[6] W.E. Lin, E. Savory, R.P. McIntyre, C.S. Vandelaar, and J.P.C. King, \u201cThe response of an overhead electrical power transmission line to two types of wind forcing\u201d, Journal of Wind Engineering and Industrial Aerodynamics. vol. 100, 2012, pp. 58-69.\n[7] P.S. Landa, \u201cStall flutter as one of the mechanisms of excitation of selfoscillations of power lines\u201d, Izvestiya vysshikh educational institutions, Applied nonlinear dynamics,\u201d vol. 17. 2009, pp. 3-15.\n[8] G.F. Kutsenko, and O. Yu., \u201cPukhalskaya Main reliability indicators of 6-10 kV transmission lines\u201d, Energetika. News of higher educational institutions and energy associations of the CIS. 2006, pp. 20-23.\n[9] S.I. Malafeev, \u201cPower supply reliability,\u201d St. Petersburg, 2018. pp.368.\n[10] R.K. Aggarwal, A.T. Johns, J.A.S. Jayasinghe and W. Su, \u201cAn overview of the condition monitoring of overhead lines\u201d, Electric Power Systems Research, vol. 53. 2000, pp. 15-22.\n[11] A. Alhassan, X. Zhang, H. Shen, and H. Xu, \u201cPower transmission line inspection robots: A review, trends and challenges for future research\u201d, International Journal of Electrical Power & Energy Systems, vol. 118, June 2020.\n[12] Composite wire for transmission lines HVCRC Smart Conductor. (http://svetofory.sea.com.ua/oborudovanie-dlya-energetiki/kompozitnyjprovod-dla-lep-hvcrc-smart-conductor)\n[13] B. Jurkiewicz, B. Smyrak, M. Zasadzi\u0144ska, K. Franczak, and P. Strz\u0119pek, \u201cThe researches of influence of strengthening on fatigue strength of aluminium wires for OHL conductors\u201d, Archives of Civil and Mechanical Engineering, vol. 19 2019, pp. 862-870. (https://doi.org/10.1016/j.acme.2018.10.006)\n[14] V.F. Terentiev, \u201cFatigue of metal materials\u201d, Science. Moscow, 2003, pp. 254.\n[15] V.P. Alekhine, \u201cPhysics of strength and plasticity of surface layers of materials\u201d, Moscow, Science, 1983, pp. 280.\nAuthorized licensed use limited to: California State University Fresno. Downloaded on July 01,2021 at 18:19:08 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_83_0003433_978-3-030-40667-7_6-Figure4.3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003433_978-3-030-40667-7_6-Figure4.3-1.png", + "caption": "Fig. 4.3 (a) The red curve shows the \u201cengineering\u201d stress\u2013strain behavior of the material, while the blue curve is the \u201ctrue\u201d one because it considers the change of the cross-sectional area A, shown in (b)", + "texts": [ + " The necking deformation is heterogeneous and will reinforce itself as the stress concentrates more at small section. Such positive feedback leads to quick development of necking and leads to fracture. If instead of considering the original crosssectional area A0 to define the stress, the real-time area A (that changes with the strain) is used, then the curve has a different behavior before the fracture. In this way, the real stress to which a point inside a material is subjected when it is strained can be calculated. In Fig.\u00a04.3a is shown the difference between these two methods. While the difference is small in the elastic region (no big difference between A and A0), it gets larger with the permanent deformation of the material. It is also visible in the example in Fig.\u00a04.3b, where the cross-section area in the point P (in the middle), where the stress is applied, decreases as the strain increases. In physics, the magnetic induction field (also improperly called magnetic field) at a point in a medium is identified by the vector B composed of a first component indicated with \u03bc0H and a second component indicated with \u03bc0M, due to microscopic phenomena that occur in the medium as a determined alignment of atomic spins. B is measured in tesla (T) or in weber on square meter (Wb/m2) and is also called magnetic flux density or magnetic induction; H is called \u201cmagnetizing field\u201d and is measured in ampere per meter (A/m) 39 or even in oersted (Oe); M is the \u201cmagnetization vector,\u201d also in A/m; \u03bc0 is the magnetic permeability of vacuum equal to 4\u03c0\u00b710\u22127\u00a0T\u00a0m/A" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002772_978-3-030-67750-3_14-Figure9.14-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002772_978-3-030-67750-3_14-Figure9.14-1.png", + "caption": "Fig. 9.14 a Mesh of original component, b mesh of optimized component", + "texts": [ + " NASA is exploring the opportunity of 3D printing in space for tool manufacturing as well to manufacture objects which previously could not be launched to space (Harbaugh 2015). TWI employed the LMD process to manufacture a helicopter engine combustion chamber as shown in Fig. 9.13. The component consists of overhanging geometries but it was built without support structures by utilizing the 5-axes of the LMD printer. The thin walled part showed a density of more than 99.5%. The part was built in 7.5 h with 70% powder efficiency (Hauser 2014). In academic literature Seabra et al. (2016) optimized the topology of an aircraft bracket (Fig. 9.14) to bemanufactured using SLM. Compared to the original part, the new part had 54% reduced material volume and weigh 28% less though the material was changed from aluminum to titanium which resulted in increased factor of safety by 2. In medical industry customization is really favored as the products must be tailored fitted for each patient, available on demand and at a reasonable price. Additive manufacturing fulfills all these demand, thus it is highly preferred by the medical industry. Additivemanufacturing offers patient-specific parts which are strong and lightweight consisting of lattice structures as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003039_icit46573.2021.9453629-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003039_icit46573.2021.9453629-Figure6-1.png", + "caption": "Fig. 6. Leaf-spring-based compression-force-reduction mechanism", + "texts": [ + " The actuator is combined with a highly sealable, rectangular, and nonstretchable fabric. It expands when air pressure is applied. The expansion force is transmitted to the human body via the aluminum plate and increases the assistive force to the upper body. The amplification mechanism can help close the plate during evacuation via a tension spring connected to the plate. This mechanism is closed when no assistance is performed, meaning that the device does not hinder the movement of the wearer. 3) Compression-force-reduction mechanism using a leaf spring: Fig. 6 shows the compression-force-reduction mechanism. A 550-mm-long and 1.00-mm-thick leaf spring (hardened SK-5 carbon steel) was used in this mechanism. The mechanism reduces the compressive force on the spine by receiving the reaction force of the assistive force when the actuator is driven. As the leaf spring used is more flexible than that used in AB-Wear II, its shape can be altered according to the spine posture. Therefore, wearers can perform complex motions that were not possible using AB-Wear II, such as bending and rotation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001107_00207160701477476-Figure9-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001107_00207160701477476-Figure9-1.png", + "caption": "Figure 9. The cylinder\u2013plane blending cyclide for conversion to rational form.", + "texts": [ + " Positive weight bi-cubic conversion of the 1/4-patch using the algorithm (left); resulting four-patch NURB representation of cyclide (right). Figure 8. Views of the positive weight four-patch 1/4-cyclide NURB conversion. the 1/4-cyclide shown in Figure 8. Hence, a 16-patch positive weight NURB parametrization of the full cyclide may be induced by the algorithm. 9.3. Blending cyclide conversions The remaining examples are concerned with rational parametrizations of cylinder\u2013plane blending cyclides, as considered by other authors (). The blending cyclide used is shown in Figure 9. D ow nl oa de d by [ U ni ve rs id ad d e Se vi lla ] at 0 1: 18 2 1 O ct ob er 2 01 4 Alternative conversions to bi-quadratic and bi-cubic rational forms \u2013 all determined from the inductive process presented in the paper \u2013 are then discussed. Many other conversions are possible. The \u03c6 displacement in all cases is defined by \u03c60 = 0, \u03c61 = \u03c0/2. Example 6 Two-patch blending cyclides. (i) The patch bi-quadratic patch on the left of Figure 10 is defined by \u03b80 = 0, \u03b81 = \u03c0 . The conversion has weights w00 = 4, w01 = 2 \u221a 2, w02 = 6, w10 = 0, w11 = 0, w12, = 0, w20 = 8, w21 = 4 \u221a 2, w22 = 6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000534_s0065-2458(08)60221-1-Figure56-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000534_s0065-2458(08)60221-1-Figure56-1.png", + "caption": "FIG. 56. Another interpretation of thc vortex chambcr amplifier.", + "texts": [ + "5 The Vortex Chamber Amplifier The vortex chamber amplifier is not only a good example of how easily momentum and boundary layer control may be combined (it may become difficult to attribute the proper importance to each of the two principles!), i t also illustrates how large a field hydrodynamics opens to the inventive mind looking around for active and passive, linear and preferably nonlinear elements. spirals, conserving their angular momentum. Thus they gain considerable kinetic energy a t the expense of static pressure, and by preventing static pressure recovery in the axial exhaust tube as much as possible, a considerable net pressure drop results. So far the device has been considered as a whole. Figure 56 introduces a new aspect in interpreting the vortex chamber amplifier by dividing it up into a switching device and a nonlinear passive element. A wall interaction type amplifier suited for boundary layer control is shown in connection with a cylindrical chamber. This cylindrical chamber represents a very interesting nonlinear device : the pressure drop is not only a nonlinear function of the flow-rate; it depends also on the direction of the flow at the inlet (omission of the wedge is possible and would make it easier to speak of one inlet), and it is this second phenomenon which is made use of in the vortex chamber amplifier", + " The fact that there is no proportionality between changes of the pressure drop through the amplifier and changes of the control pressure needs no further discussion in a survey dealing with elements for digital control. Expression (65), however, contains also the possibility of negative resistance, Analogously to procedures followed by electronic engineers the pressure drop as a function of flow-rate would then be described by curves as shown in Fig. 58. 227 H. H. GLAETTLI No conclusions as to the future use of this effect are possible at the moment. The effect has been verified experimentally in ref. 66. It is, however, expected that other possibilities in connection with the interpretation suggested by Fig. 56 may be used more successfully to effect memory. The great merit of the vortex chamber amplifier over other nonmoving part elements is the fact that it allows a more effective control of the flow rate. Contrarily to the efficiency controlled diffusor no change of the static pressure occurs if the vortex chamber amplifier is in the low resistance condition. This avoids the troubles usually encountered in recovering the static pressure. As to the use of the vortex chamber amplifier for doing logic it must be remembered that only one output signal is obtained out of one element. There is, however, more freedom than for instance in the case of the turbulence amplifier or efficiency controlled diffusor: referring to Fig. 56 it can easily be seen that due to the symmetry in the active part, a positive or negated output signal may be obtained. Although the vortex chamber amplifier projects into the third diniension, it can be considered as being essentially a two-dimensional device, which fits it to production techniques mentioned earlier. As far as means for improving or preventing pressure recovery are considered (for instance vanes) there is an interesting possibility making use of the two different senses of rotation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002945_978-3-030-69178-3_15-Figure4.11-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002945_978-3-030-69178-3_15-Figure4.11-1.png", + "caption": "Fig. 4.11 Visualisation of sensors field of view and placement in the collaborative cell", + "texts": [ + " The status of the CAS message consists of the following bits: \u2022 Active: bit value set indicates that the CAS detected a potential collision and is sending requested joint values of the robot, otherwise the joint values sent should be discarded; \u2022 Heartbeat: bit toggled in every message to indicate that the CAS works properly, \u2022 Stop: request to stop the robot in place and ignore transmitted joint values; \u2022 Miscellaneous: bits for future extension, e.g., extended collaboration scenarios, error codes; \u2022 Error: bit indicates that an internal CAS error, e.g., sensor failure, occurred. The collaborative assembly cell consists of the robot surrounded by three assembly stations and a conveyor belt. Three depth sensors are placed so that each covers all three workstations from a different angle, as presented in Fig. 4.11. To verify the placement of sensors, snapshots from IR cameras are captured while the robot is positioned in different locations. i.e., in front of each assembly station. As can be seen from snapshots presented in Fig. 4.12, all stations are covered well by the sensors. The procedure for the robot response time measurement is performed by the CAS server and covers communication delays among CAS, RCM and the robot controller in both directions, and dynamic response of the robot. During the test, logging of all messages between CAS and RCM is turned on" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002874_s10846-021-01410-5-Figure11-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002874_s10846-021-01410-5-Figure11-1.png", + "caption": "Fig. 11 Illustration of misalignment due to inaccuracies in construction. Note this offset has been exaggerated for clarity", + "texts": [], + "surrounding_texts": [ + "robot workspace G may be treated as though they occupy the same space by translating the object a set amount between frames. Effectively, conveyor space is ignored and treated as an intermediary consolidation step prior to an interaction. Figure 8b illustrates the function of the conveyor system. Note that the intermediate conveyor translate step introduces a fixed error, shown in Table 5.\nEach load-cell unit was calibrated individually. Note that a load-cell unit consists of an HX711 amplifier and TAL220 load-cell pairing. Calibration was specific to each unit and varied slightly between units. 50 initial measurements were taken with no weight present to tare the unit. A calibration tool was bolted to the load-cell and used for calibration. The calibration tool and bolts weighed 1079.60 g. Measurement accuracy was found to be consistent within \u00b1 0.1 g of the calibration weight for individual load-cell units. A 133.63 g calibration disc was used to measure the combined load-cell error. Errors associated with this subsystem are tabulated in Table 6.\nBecause an object is supported by a fixed number of loadcells\u2014and the total weight is known\u2014the ratio of weight distribution between axes can be calculated. This ratio may be used in conjunction with known distances between loadcells to compute a relative COG position within vision. The load-cell coordinate frame LC is quantified by 4 sensitive components\u2014the conveyor platform rests on these components. Load-cell space LC and the top-view camera frame It share two axes\u2014Fig. 8a and Fig. 8.\nTo formalise load-cell coordinate space in terms of vision x- and y-axes, the output of each component is measured. Ratio coefficients are used to estimate the COG position of an object xtCOG\u00bd ; ytCOG in image space It directly:\nxtCOG \u00bc acoeffx\u2212b \u00f05\u00de ytCOG \u00bc ccoeffy\u2212d \u00f06\u00de\nwhere a, b, c and d are found through experimentation. Ratio coefficients are defined by:\ncoeffx \u00bc weightRF \u00fe weightRR weighttotal\n\u00f07\u00de\ncoeffy \u00bc weightRF \u00fe weightLF weighttotal\n\u00f08\u00de\nwhere weighttotal is computed as the sum of all weight measurement inputs. A 133.63 g calibration disc was used to", + "measure the response of coeffx and coeffy as the disc was iterate across It[x, y]. Disc centre location was found through vision. The resulting relationships between It[x] and coeffx and It[y] and coeffy are graphed in Figs. 9 and 10, respectively.\nThe error associated with the COG measurement subsystem is shown in Table 7.\nDue to inaccuracies in construction, the conveyor platform surface was not perfectly perpendicular with the robotic manipulator z-plane. This offset is illustrated in Fig. 10.\nTo compensate for this error, bed level was modelled in terms of robot z-axis values. This was achieved by sampling endeffector z-axis positions of the conveyor surface along the xaxis with y-axis coordinates set to 0. A second-degree polynomial relationshipwas found to relate the true conveyor bed height with the robot x-axis\u2014illustrated in Figs. 11 and 12. This model was used to mitigate the slight error between the robot z-plane and conveyor level surface, improving manipulation accuracy.\n5 calibration discs with rubber bottoms were used to tune the vision system, find threshold values, and consolidate coordinate systems\u2014shown in Fig. 5a. Discs were matt black with 20 mm diameters. An approximate pixel to mm conversion was found by calculating the disc area in mm:\nAdiscmm \u00bc \u03c0D2\n4 \u00f09\u00de\nwhere Adiscmm is the calculated area of a calibration disc from a top-down perspective in squared millimetres. A conversion may then be established through relationship:\nmm per pixel \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi Adiscmm\nAdiscpix\ns \u00f010\u00de\nwhere Adiscpix is the area of the disc measured by the vision system, in pixels. This calculation provides an approximate conversion frommm to pixel. Note that this conversion is only an estimate and assumes a linear relationship between pixel and mm. Due to minor lens distortion, this value may vary slightly between x-axis, y-axis, and distance from the camera. 5 calibration discs were used to consolidate the G x ; y\u00bd components of the robot coordinate frame with the top-view camera frame It[x, y]. The calibration pattern is shown in Fig. 13.\nDisc centre positions were found through vision. Figs. 14 and 15 graphically illustrate the relationship between It[x] and G y\u00bd and It[y] and G x\u00bd , respectively.", + "Linear models were fit to the resulting data. These relationships are used to estimate robot coordinatesG x ; y\u00bd , given topview camera coordinates It[x, y]. Therefore, grasps generated in vision space by a grasp synthesis methodology may be converted to robot space for implementation via the following transform:\nx \u00bc a I t y\u00bd \u00fe b \u00f011\u00de y \u00bc cI t x\u00bd \u2212d \u00f012\u00de \u03b8 \u00bc \u03b8t \u00fe e \u00f013\u00de\nwhere a, b, c, d and e are found through experimentation. Robot end-effector angle \u03b8 and top-view grasping rectangle angle \u03b8t are related by a fixed offset. No significant rotational offset was found between coordinate frames. As such, simple linear relationships could be applied accurately, as opposed to computationally expensive rotation matrices that account for rotational offsets between coordinate systems. The error associated with the transformation between vision and robot frames is shown in Table 8.\nPhotoelectric diffuse-reflective sensors situated at the entrance to the vision enclosure were used to translate objects to It x-axis centre. As the assessed object is stepped toward the" + ] + }, + { + "image_filename": "designv11_83_0000914_20080706-5-kr-1001.01259-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000914_20080706-5-kr-1001.01259-Figure4-1.png", + "caption": "Fig. 4. Diagram of the wheelset and suspended mass simplified railway vehicle model.", + "texts": [ + " These calculate a lateral and longitudinal creep force at the contact patch, plus a spin creep moment about the normal axis of the contact patch. At low creepages, the creep force behaviour can be treated as linear (see figure 3). This is an acceptable assumption for dry conditions where the contact forces are sufficient to prevent easy saturation. For a given contact patch size and shape, the gradient of the creep-force curve (known as the creep coefficient) is constant. A linearised quarter vehicle model, consisting of a single wheelset and a suspended mass is used to simulate the vehicle (see figure 4). The plan view dynamics (yaw and lateral displacement) are sufficient to describe the stability and guidance response to lateral track irregularities. This model includes nonlinear wheel and rail geometries in the form of a nonlinear conicity term and is given by: y\u0308 = 1 m { \u22122f22 v ( y\u0307 + r0\u03bb l y\u0307 \u2212 v\u03c8 ) \u22122f23 v \u03c8\u0307 \u2212 W\u03bb l (y \u2212 d) + Fsy } (27) \u03c8\u0308 = 1 I { \u22122f11 ( l\u03bb r0 (y \u2212 d) + l2 v \u03c8\u0307 ) \u2212 2f33 v \u03c8\u0307 \u2212Iwyv\u03bb lr0 y\u0307 + f23 v ( y\u0307 + r0\u03bb l y\u0307 \u2212 v\u03c8 ) +W\u03bbl\u03c8 +Ms\u03c8} (28) y\u0308m = 1 mm {\u2212Fsy} (29) where y is the lateral position of the wheelset, ym is the lateral position of the suspended mass, \u03c8 is the yaw angle, W is the wheel load and d is the lateral track irregularity input" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002783_09544062211012724-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002783_09544062211012724-Figure7-1.png", + "caption": "Figure 7. Singular case where axes of three limbs intersect at one point.", + "texts": [ + " For the prismatrically actuated 3-RPR mechanism, it has three singular types, as shown in Figures 6 to 9. The singularity depicted in Figure 6 occurs in the limit situation where at least one limb of the 3-RPR PM owns zero limb length. This theoretical situation is easy to avoid in practice. Then the effect of employed load on this situation is not further discussed here. The planar 3-RPR manipulator lose its rotational DoF when three axes of limbs intersect at the common point O1, as shown in Figure 7. Two necessary conditions need to be met to have this kind of singularity when employed loads are considered, that is, 1. Axes of the three limbs can intersect at the common point geometrically within the workspace of the manipulator. 2. No torsional force is applied to the manipulator. It is clear that employed torque can eliminate this singularity. Assuming that the two necessaries conditions are met and a linear force F is employed. The singular configuration of the 3-RPR mechanism can be obtained by the error compensation method discussed in subsection 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001120_speedham.2008.4581096-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001120_speedham.2008.4581096-Figure1-1.png", + "caption": "Fig. 1. Two-phase model of an induction motor", + "texts": [ + " It shows that this estimator is less sensitive to the variation of the rotor resistor, but more sensitive to the variation of the saturation level. To overcome this problem, an adaptive estimator is proposed, based on a previous saturation phenomenon study [4]. In order to show the accuracy of the presented flux estimator, simulation results are presented using A_MOS program [5]. The classical stator flux estimator uses the stator electric equation written in a fixed two-phase reference: \u03a8=0, Fig. 1. Fig. 2 shows simulation results of a 45(KW) induction machine controlled by the DTC with the previous estimator considering that there is a difference of 15% between the motor stator resistor and its value implemented in the control estimator. Adaptive stator flux estimator for the induction machine Direct Torque Control T. Kasmieh* tkasmieh@netcourrier.com * Higher Institute for Applied Sciences and Technology, P.O.Box 60318 Damascus, (Syria) SPEEDAM 2008 International Symposium on Power Electronics, Electrical Drives, Automation and Motion 1239 978-1-4244-1664-6/08/$25" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001718_13506501jet295-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001718_13506501jet295-Figure2-1.png", + "caption": "Fig. 2 Purely squeeze-film geometry between two cylinders with different radii R1 and R2", + "texts": [ + " For h1 = 0 and h2 = h2(t), equation (13) reduces to the squeeze-film problem in parallel rectangular plates using couple stress fluid by Ramanaiah and Sarkar [17] and by Ramanaiah [18]. For h1 = 0, h2 = h2(t), and l \u2192 0, the classical Newtonian-lubricant case of the squeeze-film problem in parallel rectangular plates by Archibald [1] and Hays [2] is recovered. To illustrate the application of the non-Newtonian squeeze-film Reynolds-type equation between two convex cylinders, the squeeze-film mechanism between two different cylinders of infinite width with non-Newtonian couple stress fluids is depicted in Fig. 2, in which the upper cylinder of radius R2 is approaching the lower fixed cylinder of radius R1 with a squeezing velocity Vsq = \u2212\u2202h/\u2202t . The total film thickness h, provided R2 x, is approximated by Hamrock [6] h = h1 + h2 = hm + (R1 + R2)x2 2R1R2 (15) JET295 \u00a9 IMechE 2007 Proc. IMechE Vol. 221 Part J: J. Engineering Tribology at WEST VIRGINA UNIV on April 17, 2015pij.sagepub.comDownloaded from where hm denotes the minimum film thickness along the vertical line of centres, and h1 = hm 2 + x2 2R1 (16) h2 = hm 2 + x2 2R2 (17) In the present case, the non-Newtonian couple stress squeeze-film Reynolds-type equation (13) reduces to the following \u2202 \u2202x [ f (h1, h2, h, l) \u2202p \u2202x ] = 12\u03bc \u2202h \u2202t (18) After introducing the dimensionless quantities, the non-Newtonian couple stress squeeze-film Reynoldstype equation is written in a non-dimensional form \u2202 \u2202x\u2217 [ f \u2217(h\u2217 1, h\u2217 2, h\u2217, \u03c4) \u2202p\u2217 \u2202x\u2217 ] = \u221212 \u03b3 (19) where the functions are defined by f \u2217 ( h\u2217 1, h\u2217 2, h\u2217, \u03c4 ) = 6h\u2217 1h\u2217 2h\u2217 + 3(h\u2217 1 \u2212 h\u2217 2) 2h\u2217 \u2212 2(h\u22173 1 + h\u22173 2 ) \u2212 12\u03c4 2h\u2217 + 24\u03c4 3 tanh ( h\u2217 2\u03c4 ) (20) h\u2217 = h\u2217 1 + h\u2217 2 = h\u2217 m + (\u03b1 + 1)x\u22172 2\u03b1\u03b3 (21) h\u2217 1 = h\u2217 m 2 + x\u22172 2\u03b1\u03b3 , h\u2217 2 = h\u2217 m 2 + x\u22172 2\u03b1 (22) In these expressions \u03b3 denotes the initial film-radius parameter, \u03b1 is the radius ratio, and \u03c4 is the couple stress parameter \u03b3 = hm0 R2 , \u03b1 = R1 R2 , \u03c4 = l hm0 (23) where hm0 represents the initial minimum film thickness" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001937_6.2007-1911-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001937_6.2007-1911-Figure1-1.png", + "caption": "Figure 1. Basic Architecture of the Compliant Constant Force Electrical Contact.", + "texts": [ + "2 Exercise equipment that uses strain energy for mechanical resistance would better mimic a free-weight exercise experience if a constant resistance force were present over the input displacement range.3 Devices that provide a near-constant output force over a large prescribed deflection range are known as constant force systems. In this paper we present the robust design optimization of constant-force mechanisms with simulated pin joints (described shortly). We specifically focus on the robust design optimization of small scale (roughly 6 mm tall), mass produced, electrical contacts, the basic architecture of which is shown in Fig. 1. The purpose of the electrical contact is to provide a reliable and separable electrical connection between two electrical devices. The constant force contact is designed to carry out that purpose with minimal variation in output force over a large range of input deflection. The constant force contact is a promising solution to common signal-integrity-difficulties encountered in separable electrical interconnections; especially those used in devices or vehicles exposed to mechanical vibrations.4 As illustrated in Fig. 1, the constant force contact comprises (i) a plastic housing, and (ii) a metallic beam that is fixed at one point on the housing and makes contact with a plastic cam surface on another. A linearly increasing large input deflection is applied at a point on the beam. The reaction force at the point of applied displacement is the contact normal force and is designed to be relatively constant over a large displacement range. As will be shown, this architecture can be viewed as a flexible-beam mechanism that can be modeled using traditional kinematic analysis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001709_apec.2008.4523001-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001709_apec.2008.4523001-Figure2-1.png", + "caption": "Fig. 2 Alternator components, diode bridge and voltage regulator", + "texts": [ + " The FETs are switched synchronous to the alternator frequency, leading to a current ripple on the dc link which is determined by dc current level and load power factor. The variation of this ripple as a function of current level and load power factor will be presented. II. STANDARD ALTERNATOR The Lundell alternator is used in most automotive applications worldwide, because it is low cost and extremely rugged in the environmentally unfriendly engine compartment. This machine is separately excited with slip rings on the rotor but it contains two sets of claw poles to shape the airgap flux, colored in white in Fig. 2. The stator is a wound in a threephase configuration, with each phase displaced by 120\u00b0 [10] The stator slots are not skewed, and it is wound full pitch. The three-phase voltage waveforms are rectified to dc by 6 diodes which are mounted on the end cap of the alternator so that the alternator airflow cools the diodes and stator windings. A voltage regulator is mounted on the same end cap as the diode bridge to control the field voltage to the rotor. Field current or field voltage are not measured directly but they are controlled indirectly through the feedback loop which controls output voltage through a separate sense line" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001034_ijvd.2008.021154-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001034_ijvd.2008.021154-Figure1-1.png", + "caption": "Figure 1 Schematic of a multidisk clutch with a \u2018finger\u2019 piston: (1) finger piston; (2) apply plate; (3) friction disk; (4) steel disk and (5) end plate", + "texts": [ + " The critical speed depends on the design of the frictional system, material and dimensional parameters, and the friction coefficient. Although TEI is well defined and understood now, analysis of a practical frictional system is usually challenging because of complexity of the phenomena involved and possible complexity of the system\u2019s design. In this paper we study an important practical case: a multidisk wet clutch in an automotive transmission with a so-called \u2018finger\u2019 piston. This type of piston has non-cylindrical geometry with several axial rods (Figure 1). It is used in shifting clutches in automatic transmissions where one clutch is nested inside the actuator of another clutch and the inner clutch has to be mechanically connected with an external retainer. The gaps between the fingers of the outer clutch are used for connecting the inner clutch to the retainer. This arrangement is not unusual in modern automatic transmissions. The shifting clutches operate in the mode of short term engagement, which lasts usually less than 1s, with high initial sliding speed which decreases at high rate to reach zero at the end of the process", + " Finite element results are verified experimentally on SAE#2 testing machine. In these tests a clutch pack equipped with a special plate that imitates the finger piston was used. Based on the results of theoretical and experimental studies, countermeasures against hot spotting for a clutch with the finger piston are sought. They result in two alternative pack designs that are then evaluated using finite element simulations and experimental tests. The thermoelastic behaviour of the clutch with a finger piston shown in Figure 1 was modelled. The following components were included in the model: finger piston (1); metal apply plate (2) located between the piston and the pack of disks; pack comprising friction disks (3) interleaved with steel disks (4); metal end plate (5). Each friction disk (3) has a steel core with friction material layers bonded to both sides of it. Friction disks are rotationally connected with an inner hub (not shown), while the steel disks, the apply plate, and the end plate are connected with an external retainer (not shown)", + " As a consequence, a three dimensional model of the clutch is computationally unmanageable. However, as demonstrated in studies by Lee and Barber (1993), Zagrodzki et al. (2001), Decuzzi et al. (2001) and Zhao and Zagrodzki (2001), a two-dimensional model, which includes circumferential and axial directions, is capable of representing most important effects related to focal hot spotting. In this study we employ a 2D model shown in Figure 2 that is similar to that presented by Zhao and Zagrodzki (2001). Note that the axial cross section in Figure 1 is only included to show the arrangement of the pack and it does not reflect the relevant coordinates. The mechanical part of the problem is treated as a quasi-static multi-surface elastic contact problem. In the thermal part, a multi-region transient heat transfer problem is considered. The coordinate system used in the model is fixed to the finger piston and, consequently, to the steel disks. The thermal problem in these parts is purely conductive. In contrast, the friction disks move circumferentially relative to the coordinates and therefore the heat transfer includes both the conductive term and the term reflecting mass convection", + " Development of hot spots is initiated by pressure variation which can have different sources: geometric imperfections, structural features like the finger piston, or other factors. In this pressure variation, only the components which are stationary relative to the good thermal conductor play an important role in initiating the instability as only they can provide excitation consistent in phase with potential hot spots and therefore can effectively support their development (Zagrodzki and Truncone, 2003). In the clutch design studied in this paper, shown in Figure 1, the finger piston is stationary with respect to the steel plates. Hence, the pressure variation caused by this piston design is stationary relative to the good thermal conductor and therefore strongly contributes to the hot spotting problem. Based on discussion in Section 4.1, we expect that the hot spotting can be mitigated by a modification of the clutch pack design which will cause the pressure variation produced by the finger piston to move (rotate) relative to potential unstable modes (Zagrodzki and Zhao, 2008)", + " Since the hot spots arise in the steel disks, the stationary pressure variation supports their development in the original pack design; in contrast, the pressure variation from finger piston constantly changing phase relative to the potential hot spots in the inverted double-sided design does not support their development. In the other one of the two alternative clutch designs considered, the single-sided pack, one set of steel cores, which are exposed to frictional heating, rotates relative to the finger piston while the other is stationary. There are also other structural differences between the three packs that have important influence on their thermoelastic behaviour. In the original double-sided pack shown in Figure 1, the apply plate 2 has a sliding interface, hence is subjected to frictional heating, which is not the case with the two alternative packs. The apply plate is relatively thick; its thickness is over two times greater than that of steel plates inside the pack. It is known that the wavelength of a mode of thermoelastic instability depends on the thickness of the plate and generally increases with the thickness (Lee and Barber, 1993, Zagrodzki and Truncone, 2003). Consequently, the thick apply plate involved in the thermoelastic interaction will contribute to a mode characterised by relatively long wavelength (low number of hot spots around the circumference)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001413_j.jsv.2008.02.056-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001413_j.jsv.2008.02.056-Figure7-1.png", + "caption": "Fig. 7. Meridional stress diagram for F 01.", + "texts": [ + " (32) By substitution of Eqs. (20) and (31) into Eq. (32) and performing the integrations, one finds the equation for d4H1 to be d4H1 \u00bc H1RR1 EIz0 c\u00fe n\u20182sin a cos a 12R nc\u2018 sin a 2R . (33) To determine the change in length of the ring element in the direction of the axis of symmetry due to H1, we apply parallel to the axis of symmetry of the ring (before application of the real loading H1) the virtual loading F 01 \u00bc 1.0 lb per unit length in the circumferential direction of the ring as shown in Fig. 7. The virtual stresses sx are found by use of Eqs. (1) and (8) to be sx\u00f0F 0 1\u00de \u00bc F 01R1\u00f0Rz R1\u00de Iz0 \u00bdy cos a z sin a . (34a) In accordance with the development and discussion of Eqs. (13)\u2013(16), the virtual stresses sy will be assumed to be given by sy\u00f0F 0 1\u00de \u00bc F 01R1 cos a Rt . (34b) ARTICLE IN PRESS T.A. Smith / Journal of Sound and Vibration 318 (2008) 428\u2013460442 By invoking and applying the principle of virtual work, we find our equation for determining dVH1 to be 2pR1F 01dVH1 \u00bc Z v 2 \u00f0H1\u00des\u00f0F 01\u00dedv" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001073_s0025654407040073-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001073_s0025654407040073-Figure1-1.png", + "caption": "Fig. 1.", + "texts": [ + " EQUATIONS OF MOTION Suppose that the normal pressure is distributed over the disk according to the law (1.1). This problem was studied in [5], where the principal vector and the principal moment of friction forces acting on the disk were calculated. Owing to symmetry, the principal vector of friction forces is strictly opposite to the velocity vector of the disk center of mass; i.e., the disk center of mass moves along a straight line. Suppose that v is the velocity of the disk center of mass and \u03c9 is its angular velocity (see Fig. 1). Following [5], we introduce the dimensionless parameter k = v/u, where u = \u03c9R. We write out the equations of motion of the disk (the equation of motion of the center of mass along the x-axis and the equation of moments with respect to the center of mass): mv\u0307 = \u2212Ffr(k), Ju\u0307 R = \u2212Mfr(k), (2.1) where m is the disk mass, J is its moment of inertia with respect to the center of mass, and Ffr(k) and Mfr(k) are the friction force and the moment of friction forces with respect to the center of mass, *E-mail: gr51@mail.ru 552 which are functions of the parameter k and are given by the formulas [5] Ffr(k) = fp0R 2 \u23a7 \u23aa\u23a8 \u23aa\u23a9 \u03c02 2 k, k \u2208 (0, 1), \u03c0k arcsin 1 k + \u03c0 \u221a k2 \u2212 1 k , k \u2208 (1,\u221e), (2.2) Mfr = fp0R 3 \u23a7 \u23aa\u23a8 \u23aa\u23a9 \u03c02 4 (2 \u2212 k2), k \u2208 (0, 1), \u03c0 2 (2 \u2212 k2) arcsin 1 k + \u03c0 2 \u221a k2 \u2212 1, k \u2208 (1,\u221e). (2.3) Here f is the friction coefficient. We show how to obtain formulas (2.2) and (2.3) following [4, 5]. Let P be the disk instantaneous center of velocities (see Fig. 1); then PC = v \u03c9 = Rv u = Rk, PA = \u03c1, v = x\u0307C . \u03c9 = \u03d5\u0307. Integrating the elementary friction forces and their moment with respect to point P and using formulas (1.1), we obtain the following formulas in the polar coordinates with pole P : Ffr(k) = \u03bb \u03c0\u222b 0 cos \u03b8 d\u03b8 q2\u222b q1 S dq, k \u2208 (0, 1), Ffr(k) = 2\u03bb \u03b8\u2217\u222b 0 cos \u03b8 d\u03b8 q2\u222b q1 S dq, k \u2208 (1,\u221e), Mfr(k) = \u2212kRFfr(k) + \u03bbR \u03c0\u222b 0 \u03b8 d\u03b8 q2\u222b q1 qS dq, k \u2208 (0, 1), Mfr(k) = \u2212kRFfr(k) + 2\u03bbR \u03b8\u2217\u222b 0 d\u03b8 q2\u222b q1 qS dq, k \u2208 (1,\u221e), q = \u03c1 R , S = S(q, \u03b8, k) = q \u221a 1 \u2212 (q \u2212 k cos \u03b8)2 \u2212 k2 sin2 \u03b8 , q1 = k cos \u03b8 \u2212 \u221a 1 \u2212 k2 sin2 \u03b8, MECHANICS OF SOLIDS Vol" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002945_978-3-030-69178-3_15-Figure13.2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002945_978-3-030-69178-3_15-Figure13.2-1.png", + "caption": "Fig. 13.2 MDH-based coordinate frames of an industrial robot", + "texts": [ + " In this dynamic model, the friction in the robotic system is considered, and the friction models both in the presliding and sliding regimes are developed which can contribute to good control performance of the robotic system. This section presents the geometric model of a manipulator, and then DH geometric parameters are identified. Considering the popularity of robots in manufacturing, the geometric model of an industrial robot with 6 degrees of freedom is taken as an example which can show the whole process of modelling the geometry of the industrial robots. The MDH-based coordinate frames of an industrial robot are shown in Fig. 13.2a, and the geometric model of a real industrial robot (KUKAKR6R700 sixx [95]) that is used in a case study in the following section is shown in Fig. 13.2b. The corresponding geometric parameters based on the modified DH (MDH) reference coordinate systems built are shown in Table 13.1. Based on the reference coordinate frames specified for the manipulator, the kinematic model can be developed to obtain the transfer relation between the endeffector\u2019s pose in the Cartesian space and the position in the joint space. The transfer function is formulated as follows: X = f (q), (13.1) whereX is the end-effector\u2019s pose in the workspace, and q is the robot joint position" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000882_s1061920808040109-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000882_s1061920808040109-Figure5-1.png", + "caption": "Fig. 5. Two nongeneric linkages.", + "texts": [ + " 4(b), we show a small perturbation (obtained by decreasing l2) of the linkage under consideration: the two \u201chalves\u201d of the curve have fallen apart, and we obtain a generic linkage for which C consists of two disjoint closed curves. Note that if we increase l2 instead of decreasing it, we will obtain a generic quadrangle with a connected configuration space looking like the one in Fig. 1. 5.2. The second quadrangle that we consider here is specified by the data \u30081; 5/8, 7/8, 5/2\u3009. The position of this linkage is entirely determined by the position of the midpoint M of the bar BC (Fig. 5(a)). The figure shows two positions of the linkage, ABCD and AB1C1D, determined by the points M and M1. The second position (A,B1C1D, in which B1 = D) is in a sense degenerate: the two bars B1C1 and C1D have merged, and together they rotate about their common extremity B1 = D as M1 spins around the circle of radius 3/8 centered at D. The geometric locus of the point M consists of this circle and a concave smooth curve; the two curves have two points of common tangency (1/8, 0) and (7/8, 0). These two curves constitute the image under the projection from R 4 to R 2 given by (2) of the canonical configuration space of the mechanical linkage under consideration. The canonical configuration space itself is an algebraic curve consisting of two smooth ovals tangent at two (singular) points. 5.3. The third linkage is in a sense the most degenerate one among those considered here; all three of the mobile hinges are of length 1. As before, the position of this linkage is entirely determined by the position of the midpoint M of the bar BC (Fig. 5(b)). The mechanical system has, besides the obvious motion under which the point M rotates around the circle of radius 1 centered at the origin, two \u201cdegenerate\u201d circular motions occurring when two bars have merged, and together rotate about their common extremity; the two smaller circles along which M then rotates are shown in the figure. They are tangent to each other at the origin, and are tangent to the big circle at the points (\u22121, 0) and (1, 0). RUSSIAN JOURNAL OF MATHEMATICAL PHYSICS Vol" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001462_isam.2007.4288476-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001462_isam.2007.4288476-Figure7-1.png", + "caption": "Fig. 7. Meniscus view, front (a) and side (b)", + "texts": [ + "8mNm\u22121). The two tested components were steel (St) and silicon (Si) blades. Grippers were steel cylinders (GCSt-0x) of 2 and 3 mm of diameter with a length of 11.9 mm. The measured angles have been averaged (\u03b8A and \u03b8R) and their standard deviations computed (\u03c3\u03b8A and \u03c3\u03b8R ). 2) Volume of liquid: The liquid amount is calibrated with a manual dispensing device (from 0.1 to 2.5\u00b5L with steps equal to 2nL) or it can be estimated by surimposing the simulated meniscus on the experimental meniscus shape (see Fig. 7(a)). The volume V , depending on \u03c6 and z, can be calculated with an expression similar to 16 as briefly described in section IV-B. Fig. 7(b) presents a side view of a 0.5\u00b5L silicon oil meniscus obtained between a steel cylinder and a Si blade. For this configuration, the meniscus profile is along the cylinder axis as expected in the model. C. Force at contact This section briefly presents the force at contact curve, obtained when the gap between the cylinder and the component is equal to zero (z = 0). Fig.8 shows the force F versus the volume V , the model is in good agreement for small diameter and large volume of liquid. D. Force-distance curve This section presents simulations and experimental results of force-distance curves realised with a gap z different from zero" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001460_imece2007-41055-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001460_imece2007-41055-Figure4-1.png", + "caption": "Figure 4: Plan View Cross Section of the Pump Assembly", + "texts": [], + "surrounding_texts": [ + "The initial problem was noted to be an increase in noise level by operations personnel. Normal conversations could only be heard at a distance of at least forty feet from the pump. The sound frequency was not measured, but was noticed to be considerably lower than the 60 Hertz frequency typically emanated from electrical transformers. Investigation noted that the entire platform on which the pump rested was vibrating. The vibration could be felt when standing on the platform. This vibration led to the increased noise level near the pump. Other than the increased noise level, apparent operating problems were unnoticed." + ] + }, + { + "image_filename": "designv11_83_0002878_0954406221999077-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002878_0954406221999077-Figure2-1.png", + "caption": "Figure 2. Forced structure diagram of an artificial muscle: (a) transverse section and (b) longitudinal section.", + "texts": [ + " With equations (1) and (2), the theoretical contraction force expression of an ideal cylindrical artificial muscle can be obtained Fideal\u00f0e; p\u00de \u00bc \u00f0pD2 0=4\u00dep\u00bda\u00f01 e\u00de2 b ; 0 e emax (3) where the constants a, b and the maximum contraction ratio emax are defined a \u00bc 3 tan2h0 ; b \u00bc 1 sin2h0 emax \u00bc 1 1ffiffiffi 3 p cosh0 (4) Meller et al.22 introduced tuning parameters in the ideal model, and the empirical model is as follows Fempirical\u00f0e; p\u00de \u00bc kF\u00f0pD2 0=4\u00dep\u00bda\u00f01 kee\u00de2 b (5) where the contraction tuning parameter ke and the static force tuning parameter kF are defined ke \u00bc 1 ffiffiffiffiffiffiffi b=a p emax ; kF \u00bc Fmax \u00f0pD2 0=4\u00dep\u00bda b (6) Establishing an ideal model based on the force balance method, this model does not consider any nonlinear physical factors. As shown in Figure 2, the horizontal and vertical force balance equations are derived F\u00fe pD2 4 p \u00bc Fa \u00bc Nfa (7) pDL \u00bc 2Fc \u00bc 2nNfc (8) where Fa, Fc, fa, fc, n and N are the axial resultant force of the fiber strands, the longitudinal resultant force of the fiber strands, the axial force of a single fiber, the longitudinal force of a single fiber, the number of turns of a single fiber, and the total number of fiber strands, respectively. The relationship between the lateral and longitudinal forces of a single fiber strand is tanh\u00bc fc/fa" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002873_s00170-021-07615-0-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002873_s00170-021-07615-0-Figure3-1.png", + "caption": "Fig. 3 Stress distribution on bolted connection", + "texts": [ + " In this study, the friction coefficient of the contact surface is set to 0.15. The finite element mesh of each component and assembly is shown in Fig. 2. Taking into account the calculation accuracy and efficiency, the threaded connection and the overlapping area of the connecting plate where slippage occurs are finely meshed. The hexagonal shape of the bolt head and nut is simplified into a circle, the same as in previous work [36]. The material characteristics of the bolt connection are listed in Table 1. Figure 3 illustrates the interface pressure distribution diagram when a tightening torque of 50 nm is applied to the nut. In the bolt connection model after pre-tensioning, the stress concentration area appears at the first few turns of the thread, indicating that the screw and the nut are completely matched. The boundary conditions are applied to the model, where the left end of the connection is completely fixed, while the right end only retains the tangential displacement degrees of freedom. As shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002794_5.0049922-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002794_5.0049922-Figure1-1.png", + "caption": "FIGURE 1. The surface mesh of the glass epoxy composite driveshaft.", + "texts": [ + "00 020007-1 The glass-epoxy composite driveshaft with stacking sequence [45\u00ba/0\u00ba/90\u00ba/-45\u00ba/60\u00ba/-45\u00ba]15 is analyzed for torsional strength, torsional buckling strength, natural frequency, and critical speed in ANSYS Composite PrepPost. Material properties of glass epoxy with 60% fiber volume fraction and ply thickness is 0.125 mm have been used for analysis (Table 1). Internal and external diameters are 40 mm and 62 mm, respectively, the length of the driveshaft is 660 mm, so the number of layers is 90 [18]. The meshed model using shell 181 elements (Figure. 1) developed in ANSYS mechanical workbench having 756 elements and 768 nodes with average quality index 0.99 is used for analysis. The torsional strength required for a specific vehicle having sell growth of 9% every year is 3557 Nm [18]. This torque is applied on the glass-epoxy composite driveshaft. Figure 2 (a) shows the variation of normal stress in the nineteenth ply with maximum normal stress is at the outer fiber, which is compressive and equal to 204.02 N/mm2. The maximum shear stress is in the eighty-ninth ply as shown in Figure 2(b), equal to 45" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001151_oceans.2007.4449163-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001151_oceans.2007.4449163-Figure2-1.png", + "caption": "Fig. 2. 2-link underwater robot model", + "texts": [ + " First, through computer simulations we show that both control methods have similar control performance in continuoustime domain. Next, experiments using the underwater robot shown in Fig. 1 are done. Since digital computers are utilized for robot controllers in practical situation, the control performance is affected by discritization error of controller and sensor signals. Experimental results show that the control performance of our proposed method is better than the computed torque method. II. MODELING The UVMS model used in this paper is shown in Fig. 2. It has a robot base (vehicle) and a 2-DOF manipulator which can move in a vertical plane. First, we derive kinematic and dynamic equations of n-DOF manipulator model shown in Fig. 3 to obtain the 2-DOF mathematical model. Next, the 2-DOF mathematical model is described from the n-DOF manipulator model. Symbols are defined as follows: n: number of joint X, inertial coordinate frame Xi: link i coordinate frame (i = 0, 1, 2, , n; link 0 means vehicle) 0-933957-35-1 \u00a92007 MTS Ia,, added inertia tensor of link i with respect to Zi xo: position and attitude vector of 70(= IVOT0T]T) xe: position and attitude vector of end-effector(= [pT, OT]TT) Xo linear and angular vector of o0(= [v T, wTIT) linear and angular vector of end-effector(= [V[T, w L (3) where the linear functions \u03c1\u0303(\u03be) and \u03be\u0303(\u03be) have the following expressions \u03c1\u0303(\u03be) = \u03be \u03c1 f , \u03be\u0303(\u03be) = \u03bb f \u2212 \u03be \u2212 L R1 (4) With reference to the machine set-up shown in Fig. 2 and adopting the same notation employed in Section 4 of [6], we define the unit vector a that marks the direction of the fixed machine cradle axis a and the unit vector b(\u03c6) that marks the axis b of the pinion blank. The parameter of motion \u03c6 introduced represents the angle of rotation of the machine cradle. All the machine motions depend on \u03c6. We refer to the Gleason\u2019s implementation of the face-milling process [12], where vertical motion \u2206EM(\u03c6) and helical motion \u2206XB(\u03c6) are polynomial functions of \u03c6 \u2206EM(\u03c6) = \u2206EM0 + V1 \u03c6 + 1 2 V2 \u03c6 2 + 1 3 V3\u03c6 3 (5) \u2206XB(\u03c6) = \u2206XB0 + H1 \u03c6 + 1 2 H2 \u03c6 2 + 1 3 H3 \u03c6 3 (6) Functions \u03c8(\u03c6) and \u03d5(\u03c6), controlling the rotation of the cradle and the blank, respectively, (see definition (39) in [7]) specialize in \u03c8(\u03c6) = \u03c6; \u03d5(\u03c6) = m0 [ \u03c6 \u2212 2C 2 \u03c62 \u2212 6D 6 \u03c63 \u2212 24E 24 \u03c64 \u2212 120F 120 \u03c65 ] , (7) Copyright c\u00a9 2007 by ASME url=/data/conferences/idetc/cie2007/71790/ on 04/10/2017 Terms of Use: http://www", + " In the most general framework of UMC machines, also the machine root angle \u03b3m(\u03c6), the radial setting S r(\u03c6) and the machine center to back \u2206XD(\u03c6) can vary during generation as functions of the motion parameter \u03c6, but in the mathematical model presented in this paper they will remain fixed. Also the tilt and swivel angles, \u03c3(\u03c6) and \u03b6(\u03c6) respectively, are represented here by constant terms \u03c30 and \u03b60. The cutter head surface, expressed in its reference configuration by the position vectors pe(\u03be, \u03b8), can change its orientation according to two subsequent rotations \u03c30 and \u03b60 about unit vectors jc and kc, which mark the tilt and swivel axes (Fig. 2). In the proposed invariant approach, this is expressed by the following relation p\u0303e(\u03be, \u03b8) = R ( R(pe(\u03be, \u03b8), jc, \u03c30),kc, \u03b60 ) (8) We recall (e.g., [6], [10]) that the rotation operator R in (8) applied, e.g., to the vector pe, the axis unit vector jc and the angle \u03c30, performs the following operation on pure vectors R ( pe, jc, \u03c30 ) = (pe \u00b7jc)jc+[pe\u2212(pe \u00b7jc)jc] cos\u03c30+jc\u00d7pe sin\u03c30 (9) In S the components jc and kc of jc and kc have the following simple expressions jc = (cos q, sin q, 0), kc = (0, 0,\u22121) (10) 3 Downloaded From: http://proceedings", + " The expressions for f and \u03a6 are now pretty straightforward to obtain by employing definitions (14) and the results in (13), (15), and (16). It is worth observing that in the invariant approach all geometrical quantities have a unique representation. Let us define x \u2208 R n to be the vector of the n machinetool settings selected as optimization variables. We do not need to distinguish between parameters related to the geometry of the tool surface like, e.g., Rp and \u03b1p in Fig. 1, and those related to the machine kinematics like, e.g., \u2206XD and 2C in Fig. 2, to mention just a few. The vector of the basic (initial) machine-tool settings is defined as x(0) and its current value at iteration k is x(k). To compact the notation, let us introduce the vector of the Gaussian and motion parameters \u03b6 = (\u03be, \u03b8, \u03c6) and its discrete version \u03b6 i = (\u03bei, \u03b8i, \u03c6i) representing their values at the i\u2212th point Pi of an m-point grid sampled on the tooth surface with settings x. Similarly, we define the i\u2212th triplet on the basic tooth surface as \u03b6(0) i (i.e., for x(0)) and \u03b6(k) i at iteration k (i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002563_012068-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002563_012068-Figure2-1.png", + "caption": "Figure 2. ROV frame of reference and Earth reference frame", + "texts": [ + " Note that by convention for underwater vehicles, the positive x-direction is taken as forward, the positive y-direction is taken to the right, z- the positive direction is taken as down, and the right-hand rule applies to angles. Table 1 describes the standard notation of motion of objects. International Conference on Technology and Vocational Teachers (ICTVT) 2020 Journal of Physics: Conference Series 1833 (2021) 012068 IOP Publishing doi:10.1088/1742-6596/1833/1/012068 frame that will be used as a standard in determining the motion of objects as in Table 1. From Figure 2 there are 3 axes (x, y, z), each of which has the following functions: Axis x Surge The ROV moves forward / backward in the direction of the x axis Roll The ROV rotates about the x axis Y axis Sway The ROV moves sideways along the y axis Pitch ROV rotates about the y axis Sumbu z Heave OV moves up / down following the z axis Yaw ROV rotates about the z axis To be able to drive the ROV and convert energy from electricity to force, a BLDC motor equipped with a propeller is used. Basically, the BLDC motor works by using the principle of the attractive force between two magnets with different poles or the repulsion between two magnets with the same poles [4]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003438_j.indcrop.2021.113890-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003438_j.indcrop.2021.113890-Figure2-1.png", + "caption": "Fig. 2. Response surface for the effects of a) ultrasound power versus buffer solution/oil mass ratio(B/O), b) ultrasound power versus catalyst/substrate mass ratio (C/S), and c) catalyst/substrate mass ratio (C/S) versus buffer solution/oil mass ratio (B/O) on the yield of enzymatic hydrolysis of crambe oil using fresh seeds.", + "texts": [ + " = experiment; Po = ultrasound power; C/S = catalyst/substrate ratio; B/ O = buffer solution/oil ratio; PC = point central. F. Tavares et al. Industrial Crops & Products 171 (2021) 113890 room temperature, with a yield of 87.98 %. In the same conditions, one experiment was carried out yielding 85.0 \u00b1 2.6 %, which confirms that within the studied range, the model can predict the yield satisfactorily. The response surface plot for the enzymatic hydrolysis yield of the crambe oil with fresh seeds is presented in Fig. 2. Each of the variables can affect the hydrolysis reaction by giving more or less energy to the system, causing enzymatic denaturation, modifying the medium viscosity, and altering the availability of sites for lipase catalysis. Nevertheless, the ultrasound power does not have a significant effect on the response surfaces in the investigated range, unlike the B/O and C/S variables. Although the amount of buffer solution affects the yield of the enzymatic hydrolysis reaction, this effect will have less influence on the reaction yield above 1.0 as seen in Fig. 2 (a). A similar effect is observed in Fig. 2 (b), where for a ratio of catalyst/substrate above 0.15, the impact of the variable is less significant; conversely, an increase in the yield can still be observed when increasing the ultrasound power. Notably, a high amount of catalyst (fresh seeds) makes the reaction medium more viscous, demanding more ultrasound power to emulsify the medium, resulting in the correlation of the two variables. In Fig. 2 (c), the effect of the most significant variables are evaluated simultaneously. The increase of B/O by keeping C/S constant and vice versa results in a limited increase of yield, however, because higher C/S means more catalyst to the reaction and higher B/O means lower medium viscosity resulting in better mixing and dispersion of catalyst, both variables must be increased for higher yields. The reactions were performed under the conditions of the optimal values obtained from Eq. 4 (Power 70 % for 3 min at initial room temperature, 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001657_imece2008-66830-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001657_imece2008-66830-Figure5-1.png", + "caption": "Fig. 5 FE analysis data of frictional shear forces on a pressure flank of bolt\u2019s thread", + "texts": [ + " Figure 4 shows the surface pressure distribution on the thread surface and bearing surface when radius Rp = 10 mm is used. Immediately after the initial bolt axial tension is generated, the surface pressure is distributed evenly on the thread surface and bearing surface. However as the washer is moved and external force is increased, the surface pressure increases in the forward direction of the external force. This trend is the same for both the thread surface and bearing surface. The distribution of friction shear force is indicated by vectors. The results when radius Rp = 10 mm is used are shown in Fig. 5 (a) and Fig. 5 (b). The results when radius Rp = \u00a5 is used are shown in Fig. 5 (c). In all cases, the figures show a view from the bolt shank end side, and the bold solid line arrow at the top of each figure indicates the direction of loosening rotation. Figure 5 (a) shows step 2 from Fig. 3 (a), and is composed of loosening-side vector components with instantaneous center at point Cp. Figure 5 (b) shows the vector components in the loosening direction on the right half, and vector components in the tightening direction on the left half; the instantaneous center is separated from the thread surface and located on the left side of the figure. This process corresponds to step 3 in Fig. 3 (a), with bolt tightening rotation 3 Copyright \u00a9 2008 by ASME rl=/data/conferences/imece2008/70944/ on 02/17/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Down loaded 4 Copyright \u00a9 2008 by ASME From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/imece2008/70944/ on 02/17/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Dow followed by loosening rotation. For this rotation, as shown in Fig. 5 (b), a vector representing a large loosening rotational force occurs at a lower right position on the thread surface. While the value of this vector is small initially, as the external force gradually increases and slip on the thread surface progresses, the surface pressure at the bottom right position, which is a rising surface, changes and increases as shown in Fig. 4. As a result, it is assumed that the vector component for loosening rotation increases and exceeds the rotational force in the tightening direction. Figure 5 (c) shows a case in which the washer was displaced linearly in the x-axis direction. Even when the washer is displaced linearly with radius Rp = \u00a5, the instantaneous center Cp, derived from the vectors of friction shear force on the thread surface, is located close to the thread surface, and is not apart from the bolt body portion. During vibration, when the washer movement is reversed and displaced in the x-axis minus direction, this rotation center reverses to the minus direction of the y-axis, with the x-axis as the axis of symmetry. The results are the same as the vectors which would be obtained by rotating all of Fig. 5 (c) by 180\u00b0. In addition, at radius Rp = 0, vectors occur around the center of the thread axis. Because these vectors can be easily visualized, the vector diagram has been omitted. Figure 6 shows the slide direction of the thread surface and bearing surface. On the bearing surface, the washer displacement has a strong effect on the slide direction. However on the thread surface, the rotational component is stronger than the translational component of the washer displacement. The reason is that the slide component in the bolt axis radial direction on the thread surface is small because the thread half-angle is large and the sliding during climbing on the thread surface is prevented. In the case of radius Rp = \u00a5, there is almost no difference from the instantaneous center that is intuitively derived from the vector directions in Fig. 5 (c). In the case of radius Rp = 10 mm, in Fig. 5 (a) the rotational component around the bolt axis Fig.7 Directions of external load Pitch diameter d2 of threads \u03b8i \u03bei -\u03b8i Rf Rfri Center of torsion Pi x y r u OC Rp Rfui r2 nloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?ur center becomes larger, and as a result it can be determined that the instantaneous center of rotation moves to a position closer to the bolt axis center as shown in Fig. 6 (b) (Rp = 10 mm). Therefore when the bolt rotation direction components in the friction shear force vectors is examined, it is assumed that the instantaneous center of rotation is eccentric in the same direction as the slide vector instantaneous center of rotation", + " Next, the friction shear force will be derived when an external force is applied from the balance of forces on the thread surface, and compare the force with the results of FEM. Under conditions with only translational force applied (equivalent to radius Rp = \u00a5), Yamamoto, Kasei, et al assumed that the external force load points were located at only two locations on a cross sectional surface that passes through the bolt center, and calculated the slide direction [3]. At these two locations, the slide direction calculated in this study using FEM agrees with these results. Based on the formula for the balance of forces used by Yamamoto et al, the distribution shown in Fig. 5 was derived. First the external forces operating on the thread surface were expressed as shown in Fig. 7. In Fig. 7, point OC is on the bolt axis line, and r2 is a length equal to 1/2 of the pitch diameter. The point Rp on the abscissa is found, and external force is applied to the thread surface in a direction perpendicular to the bolt axis. The force is exerted as a torque around point Rp. This external force is distributed over the pitch diameter, and its vector is Rfi, using local Cartesian coordinates (r-u coordinates)", + " If the number of divisions is the same for the thread surface and bearing surface, the divided bearing surface friction force \u03bc\uff57 \u00d7 Ff i is hypothesized to be equal to the resultant value of the translational component and the rotational component of the thread surface friction force. However, when this value is calculated for the thread surface, because there is a difference between the sizes of the bearing surface friction equivalent diameter dw and the thread surface effective diameter d2, the friction force in the circumferential direction (which becomes the rotational component) must be corrected using the value of the ratio of the two diameters. Because the instantaneous center of rotation is not at the bolt axis center, as shown in Fig. 5 and elsewhere, the required correction value is not the constant value of dw \u00f7 d2, and varies according to the position on the thread surface. Therefore, based on the provisional assumption of dw / d2 \u00bb 1.0, the calculations were proceeded. In addition, after the specifications of the joined structure have been determined, coefficient C is applied in consideration for the distance between the bearing surface and thread surface in the bolt axial direction, and CRfi becomes the external force exerting on the thread surface", + " 2/)1()sin(sec cos}2/)1()sin({ 1 1 DAFf DAFdu ngis ngSi +\u00b4\u00b4am+ b\u00b4-\u00b4\u00b4= (7) When this is converted to the global coordinate system (x-y coordinate system), the vector is obtained. iiiii dudrdx xx sincos -= iiiii dudrdy xx cossin += (8) 7 Copyright \u00a9 2008 by ASME rl=/data/conferences/imece2008/70944/ on 02/17/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Dow Figure 9 shows the results obtained from the equations explained above, expressed as vectors. The conditions of the radius Rp and load direction shown in Fig. 9 each correspond in sequence to each of the conditions in Fig. 5. The directions of the vectors shown in Fig. 9 closely match the directions of the vectors shown in Fig. 5. However, there are differences in the absolute values of vectors. In Fig. 5, the values at positions close to Cp are large. One possible reason for this is that the calculations for Fig. 9 do not consider the uneven surface pressure distribution on the thread surface which changes as the load increases due to factors such as the thread surface inclination. Another possible reason is that this calculation assumes the value of dw/d2 is constant, which means the instantaneous center is located at the bolt axis center. In the FEM results, the average of Rfri is approximately 97% of mw \u00d7 Ffi" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001982_20080706-5-kr-1001.01655-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001982_20080706-5-kr-1001.01655-Figure2-1.png", + "caption": "Fig. 2. Mechanical composition of the robot", + "texts": [ + " The robot arm provides a basis for the coupling of auxiliary components in order to develop a wide range of applications. The system configuration of the laboratory robot \u2013 mechanical and from a information processing point of view \u2013 is described in the following section. In Figure 1 the arrangement of the equipment is displayed. The central element is the robot portal. Beside the portal the control cabinet is arranged. The cabinet contains the complete control hardware including axis processors and robot control computer. The mechanical resp. kinematical composition is displayed in Figure 2. The axle drives constitute the link between robot control and mechanical components. They are connected via analog or digital interfaces resp. the implemented Controller Area Network (CAN) bus to the axis processors of the lower control level. The axis processors communicate among themselves or with the robot control processor via CAN bus. The robot used in the laboratory platform is a folding arm robot with five axes (serial kinematics). The folding arm is attached suspending to a carrier moving on two parallel longitudinal axes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003407_j.optlastec.2021.107425-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003407_j.optlastec.2021.107425-Figure2-1.png", + "caption": "Fig. 2. Double ellipsoid heat source model and heat source movement path (a) Double ellipsoid heat source model (b) Heat source movement path.", + "texts": [ + " In the SLS process of WSPC powder, the laser radiation energy acts on the powder bed in the form of heat flow. However, the laser radiation energy will be transmitted along the depth direction of the powder bed through interspace between the powder particles, so the effect of laser radiation energy in the depth direction of the powder bed should also be considered. Therefore, volume heat source should be used for analysis. The most representative heat source in the volume heat source model is the double ellipsoid heat source model in Fig. 2(a), and its movement trajectory is shown in Fig. 2(b). The double ellipsoid heat source model is composed of two 1/4 ellipsoids. The power density gradient of the front ellipsoid is large, while that of the rear ellipsoid is small. The power density distribution functions of the front and rear ends are as follows: q(x, y, z, t) = 6 \u0305\u0305\u0305 3 \u221a \u03b7ff Q abc1\u03c0 \u0305\u0305\u0305 \u03c0 \u221a e \u2212 3x2 a2 e \u2212 3[y+v(\u03c4\u2212 t)2 ]2 c2 1 e \u2212 3z2 b2 (7) q(x, y, z, t) = 6 \u0305\u0305\u0305 3 \u221a \u03b7frQ abc2\u03c0 \u0305\u0305\u0305 \u03c0 \u221a e \u2212 3x2 a2 e \u2212 3[y+v(\u03c4\u2212 t)2 ]2 c2 2 e \u2212 3z2 b2 (8) Whereby, ff is the proportional distribution coefficient of the total energy occupied by the ellipsoid at the front end of the model; Q represents the laser power (W); \u03b7 is the absorption rate of powder; v is the moving speed of the heat source (m/s); \u03c4 is the time delay factor; t is the acting time of the heat source (s); fr is the proportional distribution coefficient of the total energy occupied by the ellipsoid at the rear end of the model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002138_s11668-020-01090-7-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002138_s11668-020-01090-7-Figure6-1.png", + "caption": "Fig. 6 a Schematic of opening and closing behavior of crack, b Model of crack closure line", + "texts": [ + " F1 and F2 are the stress intensity factor correction factors of a mode I crack. FII and FIII are the stress intensity factor correction factors of mode II and mode III crack.h is the angle between the axis and the crack. In this paper, the shaft crack is a transverse crack with a specified angle, h \u00bc p=2. Write the flexibility coefficients in matrix form. According to the static equilibrium theory, the stiffness matrix of the crack element can be obtained by using the transformation matrix T. Kc\u00bcTF 1TT \u00f0Eq 29\u00de where F is the flexibility coefficients matrix. As shown in Fig. 6, the crack opens and closes periodically as the shaft rotates. Darpe [13] et al. make the point where the SIF of mode I crack is zero as the crack closure line position (CCLP). According to this theory, the range of integration of Eqs 7\u201328 can be determined. The first two elements in the leading diagonal of the flexibility matrix F as a function of rotation angle are shown in Fig. 7. In this case, the radius and the length of the cracked element are 15 mm and 60 mm, respectively. When the crack depth (a = 15 mm) is the radius of the rotating shaft, all the flexibility coefficients increase as the crack opens, and decrease as the crack closes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002873_s00170-021-07615-0-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002873_s00170-021-07615-0-Figure1-1.png", + "caption": "Fig. 1 Bolted test sample", + "texts": [ + " Therefore, the mathematical model of the pre-tightening torque Tf can be obtained as [32]: T f \u00bc P 2\u03c0 \u00fe 1:155\u03bctrt \u00fe \u03bcbrb Fb \u00f02\u00de where, P is the pitch of threads, \u03bct is the friction coefficient of the contact surface between the external thread of the bolt and the internal thread of the nut, rt is the pitch radius of the bolt, \u03bcb is the friction coefficient between the lower surface of the nut and the bearing surface of the joint member, while rb is the effective radius of the nut surface. In order to investigate the effect of vibration on bolt preload attenuation of the bolted connection in machine tools, a connection structure that is commonly used in machine tools was designed to perform bolt looseness tests. As shown in Fig. 1, the bolt loosening experiment was carried out on the designed specimen. The design of the sample meets the requirements of the INSTRON 8801 testing machine, which is used for the loosening experiments. Hexagon bolts of strength grade 8.8 with M16\u00d780 nuts were used in the loosening tests. During the nut tightening process, the nut is lowered 1 mm above the contact surface, while then a digital torque wrench is used, to achieve the torque load required for the test [33]. The same tightening operation is performed on each test sample, to ensure that the same torque load is applied on the nut. The research object shown in Fig. 1 is used to explore the influence of deterministic load on the characteristics of the bolted connection in machine tools. First, the finite element model of the bolted connection is established, while the deterministic load is applied to the bolt connection after pretightening for simulation calculation. Croccolo [34, 35] assumed that the friction coefficient of the threaded joint friction interface would change during pre-tightening. The threaded connection is used in the finite element model because the friction coefficient of the contact surface needs to be considered, when the nut is pre-tightened" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002874_s10846-021-01410-5-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002874_s10846-021-01410-5-Figure1-1.png", + "caption": "Fig. 1 Annotated isometric view of the proposed benchmarking system with covers removed from the vision enclosure", + "texts": [ + " By maintaining consistent test conditions and employing comprehensive evaluation criteria, a grasping pipeline may be placed with respect to other approaches and the advantages and disadvantages of a specific methodmay be recognised. The benchmarking template offered in this paper is industrially focused and aimed at 2-fingered, force-closure grasp * Jacques Janse van Vuuren j.jansevanvuuren@massey.ac.nz 1 Department of Mechanical and Electrical Engineering, SF&AT, Massey University, Palmerston North, New Zealand synthesis methodologies that utilise machine learning (ML). The platform is largely comprised of a conveyor system, vision enclosure and robotic manipulator\u2014Fig. 1. Two RGB cameras are employed, offset by 180\u00b0. The conveyor rests on a load-cell subsystem, which is capable of weighing objects and centre of gravity (COG) acquisition. A beam sensor is used to automatically situate assessed objects near the centre of the vision system. Construction of the proposed system is relatively simple and is mainly comprised of extruded aluminium. The robotic manipulator is part of an education package costing approximately $1600 USD. The calibration and integration of various subsystems is also considered in this paper", + " Key aspects of design, calibration and other details related to the proposed benchmarking platform are covered in this section. Comprehensive documentation for construction is availa b l e a t : h t t p s : / / d r i v e . g o o g l e . c om / o p e n ? i d = 1VsEjCl6hrX3FeL9VRF-J9CzVL7JHXO15. The proposed benchmarking platform illustrated in Fig. 2a is portable and self-contained. The layout consists of a conveyor system, vision enclosure and robotic manipulator, framed by 45 \u00d7 45 mm slotted aluminium extrusion. Figure 1 illustrates an annotated diagram of the proposed system. The 460 \u00d7 430 \u00d7 420 mm (L \u00d7W \u00d7 H) vision enclosure employs HD webcams, diffuse LED lighting and a photoelectric through-beam sensor. The inside panels were matt white. 2 LED strips were mounted at the top of the enclosure, facing downward. A plastic diffusor was added at the top of the enclosure, below the lighting. Two identical Microsoft Livecam Studio webcams were used for vision\u2014offset by 180\u00b0. Specifications related to the components present in the vision enclosure are tabulated in Table 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001609_indcon.2008.4768808-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001609_indcon.2008.4768808-Figure4-1.png", + "caption": "Fig. 4. Possible orientations for a fixed tip position.", + "texts": [ + " This is indeed very difficult task even for a random test motion generation. The human hand-eye coordination is very good in judging relative sizes and distances whereas a computer requires precise quantity for delivering the result. Effect ofend-effector's position on its orientation: It is well known that we can vary the position of the wrist using the first three major links in an articulated robotic arm. But it is often overlooked that the end-effector's orientation would also continuously vary with the movement. A reverse phenomenon is also possible. Consider Fig. 4. In this, the end-effector's tip can be maintained at a specific location Pd, still any arbitrary orientation can be generated by moving the wrist all around on an imaginary sphere centered on the tip, using all the six joints. This, of course, subjected to the available ranges of the various joints. The issue, however, is that how to arrive at the precise data (p,R) to tell the robot what it must do? We, therefore, have to use the GUI, described in the next section, to move the robot as per our wish in an online interactive manner" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000629_9780470612286.ch1-Figure1.6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000629_9780470612286.ch1-Figure1.6-1.png", + "caption": "Figure 1.6. Example of a planar robot with three degrees of freedom", + "texts": [ + " Calculation of the Jacobian matrix by derivation of the DGM The calculation of the Jacobian matrix can be done by differentiating the DGM, X = f(q), using the following relation: i ij j f ( ) J i l, , m; j l, , n q \u2202 = = = \u2202 q \u2026 \u2026 [1.21] where Jij is the (i, j) element of the Jacobian matrix J. This method is easy to apply for robots with two or three degrees of freedom, as shown in the following example. The calculation of the kinematic Jacobian matrix presented in section 1.3.1.2 is more practical for robots with more than three degrees of freedom. EXAMPLE 1.4.\u2013 let us consider the planar robot with three degrees of freedom of parallel revolute axes represented in Figure 1.6. We use L1, L2 and L3 to denote the lengths of the links. We choose as operational coordinates the Cartesian coordinates (Px, Py) of point E in the plane (x0, y0) and the angle \u03b1 between x0 and x3. Px = C1 L1 + C12 L2 + C123 L3 Py = S1 L1 + S12 L2 + S123 L3 \u03b1 = \u03b81 + \u03b82 + \u03b83 with C12 = cos(\u03b81 + \u03b82), S12 = sin(\u03b81 + \u03b82), C123 = cos(\u03b81 +\u03b82 + \u03b83) and S123 = sin(\u03b81 +\u03b82 + \u03b83) The Jacobian matrix is calculated by differentiating these relations with respect to \u03b81, \u03b82 and \u03b83: S1L1 S12L2 S123L3 S12L2 S123L3 S123L3 = C1L1 C12L2 C123L3 C12L2 C123L3 C123L3 1 1 1 \u2212 \u2212 \u2212 \u2212 \u2212 \u2212\u23a1 \u23a4 \u23a2 \u23a5+ + +\u23a2 \u23a5 \u23a2 \u23a5\u23a3 \u23a6 J 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001190_09544062jmes539-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001190_09544062jmes539-Figure1-1.png", + "caption": "Fig. 1 Configuration of the lubricated bodies and coordinates", + "texts": [ + " The main purpose of this study was to extend the transient EHL analysis under a load impulse considered in reference [15] to a complete time-dependent numerical solution associated with a normal vibration under thermal and nonNewtonian conditions with acceptable time of computation. Because harmonic oscillation is the basic form of the vibration phenomena, as a preliminary work, in the current paper the elliptical point contact is idealized as composed of a spherical roller and an infinite plane, and, it is further assumed that the roller is oscillating harmonically in the normal direction against the infinite plane, when both surfaces are running with a certain slide\u2013roll ratio. Figure 1 shows the configuration of the spherical roller and the infinite plane studied in this work. When a constant load w0 is applied normally to the roller without lubricant, both surfaces are deformed elastically due to the contacting pressure, and a contact ellipse is therefore generated. For the given radii of curvatures Rx and Ry , the minor and major radii of the ellipse, denoted by a and b, respectively, together with the maximum Hertzian contact pressure pH, can be calculated easily with the empirical formula proposed by Markho [17]. In Fig. 1, the entrainment velocity is aligned with the minor radius of the contact ellipse, therefore, the normal gap between solids \u2018a\u2019 and \u2018b\u2019, i.e. the timedependent EHL film thickness, is given by h(x, y, t) = h00(t) + x2 2Rx + y2 2Ry + 2 \u03c0E \u2032 \u222b \u222b p(x\u2032, y \u2032, t)dx\u2032 dy \u2032\u221a (x \u2212 x\u2032)2 + (y \u2212 y \u2032)2 (1) in which the cyclic function h00(t) stands for the vibration of the roller h00(t) = hs \u2212 A sin \u03c9t (2) where \u03c9 = 2\u03c0f , f is the frequency of the vibration, A the amplitude of the vibration, and hs a constant in correspondence with the reference load w0 when vibration is absent" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002772_978-3-030-67750-3_14-Figure30.10-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002772_978-3-030-67750-3_14-Figure30.10-1.png", + "caption": "Fig. 30.10 aMotion study of wheelbarrow loader bucket. b Load applied on the loader bucket and wheelbarrow tray", + "texts": [ + " These simulations are carried out to ensure the structure design is sufficient to withstand the load applied. There are a few analysis obtained by carrying out an analysis and simulation such as stress, strain, displacement, and motor force. Firstly, the motion study was done to the design. It is very crucial to study the movement of each part and to observe where parts may collide each other so that the parts can be modified to move free as intended without any problems when the fabricating process is done. The motion study is shown in Fig. 30.10a and the results show no collision happened and the design is ready for the next level of analysis which is the structural analysis. During the structural analysis, the load of 500 N is applied to the loader bucket as in Fig. 30.10b and the deformation at certain parts is observed and analyzed to find the affected area of the developed model so that the model can be modified to have more strength and durability during the real-time use. The stress\u2013strain analysis results show that the highest value of stress is 6.297e + 07 N/m2.. In contrast, 6.247e-04 for a strain that occurs at the linkage between the loader bucket and body bracket that is attached to the wheelbarrow body tray and this value is still considered acceptable and can be seen in Figs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002581_icees51510.2021.9383730-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002581_icees51510.2021.9383730-Figure4-1.png", + "caption": "Fig 4. Stator Geometry with 36 slots", + "texts": [ + " Stator Geometry Stator is made up of dynamo grade laminations which are of thickness 0.35 mm or 0.5 mm. For motors of larger size cores are made of segmented laminations. To give most economical balance between the costs of dies, the peripheral length usually between 0.3 m to 0.6 m is chosen for one segment, the left over amount of scrap from the lamination cuttings from steel strips and the assembly cost. In the flux paths of alternating poles, to provide an equal number of turns, the total number of segments is chosen. Fig 4 shows the stator geometry of 36 slots. F. Rotor Geometry In BLDC motor two types of rotor design are there, \u2022 Outer Rotor \u2022 Inner Rotor The windings are located in the motor core for an outer rotor design. Fig 5 shows the stator windings surrounded by the rotor magnets. 2021 7th International Conference on Electrical Energy Systems (ICEES 2021) 119 Authorized licensed use limited to: Carleton University. Downloaded on June 06,2021 at 01:21:49 UTC from IEEE Xplore. Restrictions apply. G. Introduction o f fins The fins are the extended surface that is being used to decrease the thermal resistance at the motor\u2019s solid parts such as stator, magnets and rotor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000676_978-1-4020-8829-2_9-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000676_978-1-4020-8829-2_9-Figure1-1.png", + "caption": "Fig. 1. Particular cases when solving assembly constraints during optimization", + "texts": [ + " In some cases, according to the parameters given by the optimizer (e.g. dimensions of the mechanism, values of the generalized coordinates, . . . ), we may face convergence problems of the Newton-Raphson algorithm to solve the constraints (2). Especially, such problems may frequently occur when performing a geometrical optimization process, i.e. whose optimization variables are the dimensions of the MBS. To illustrate that, let us consider a four-bar mechanism in relative coordinates with a point-to-point constraint as assembly constraint (see Fig. 1) with a given partition of the variables. On the basis of this example, three particular cases can be identified when evaluating the mechanism during optimization: 1. A first case may occur when the solution in terms of v is not unique for a given vector u (or scalar u in the case of Fig. 1a). These multiple assembly solutions are obviously inconvenient since the performance are usually different for each assembly case. In most cases however, the problem requirements favor one assembly solution more than others. One possible way to reach that preferable assembly configuration is to start the Newton-Raphson algorithm with initial dependent variables close to that expected solution. But good initial variables are not always easy to guess if the dimensions vary during optimization. Another possibility is to add constraint to the assembly problem, using optimization techniques. This will be applied in the following optimization strategy of Section 3. 2. A second problem may happen if the mechanism reaches a singular configuration, the constraint Jacobian matrix Jv = \u2202h \u2202vT becoming singular (see Fig. 1b). These singularities have both a mathematical and a physical interpretation. Mathematically, they are related to the chosen partitioning and corresponds to an ill-conditioned Jacobian matrix. Physically, those singularities correspond to a loss of mobility of the MBS by locking one or more actuators associated to the independent joint variables. Let us note that purely mathematical Jacobian singularities may arise for instance: \u2022 From a wrong choice of the sequence of rotation variables in three dimensions \u2022 When redundant constraints are present \u2022 Or even if an ignorable variable is taken into account in the set of dependent variables (e", + " Practically, the proximity to the singularity should be detected thanks to the conditioning of the constraint Jacobian matrix and a new partition should be achieved accordingly. This possibility has been studied for redundantly actuated MBS by [9]. In this work, the particular case of overactuated MBS models is not considered and the partition is fixed according to the problem requirements. 3. Finally, it can be impossible to close the mechanism simply because a constraint hi has no root: the Newton-Raphson algorithm cannot converge toward a solution (see Fig. 1c). This situation occurs for instance when the optimizer evaluates the objective function with incompatible dimensional parameters (e.g. too short bars in Fig. 1c). In that case, the closed mechanism does not exist but, instead of rejecting it, we will keep it and penalize the cost function accordingly. The penalty strategy is presented in the next section. 3 Best-Assembly Penalty Optimization Strategy Suppose that we want to optimize the performance of a MBS given by the objective function f with respect to the vector of p design variables l. This objective f also depends on the n generalized coordinates q that are constrained by m independent non-linear assembly constraint equations hi(q) = 0, i = 1, ", + " The definition of these additional constraints therefore requires to know in advance the different solutions and singularities of the assembly problem. This prerequisite can sometimes be difficult to establish when complex topologies are considered. However, for many applications involving several simple loops of bodies such as parallel manipulators (see applications in Section 5), the construction of function c remains affordable. The following planar example will provide more details on the method. Assume that we want to assemble a four-bar mechanism modeled in relative coordinates (see Fig. 2). As shown in Fig. 1a, two assembly solutions exist. The additional requirement is to keep the \u201celbow-up\u201d closed configuration (i.e. when points A or B remain above the straight line between \u03b82 and \u03b83 in Fig. 2) and to avoid the singular configurations when bars are aligned. This model contains three generalized coordinates \u03b81,\u03b82, and \u03b83. The four bars have lengths l1, l2, l3 and l4. To make points A and B coincide, the two assembly equations are: h1(\u03b81, \u03b82, \u03b83) = \u2212l1 sin \u03b81 + l2 cos (\u03b81 + \u03b82) + l3 sin \u03b83 \u2212 l4 = 0 , (5a) h2(\u03b81, \u03b82, \u03b83) = l1 cos \u03b81 + l2 sin (\u03b81 + \u03b82) \u2212 l3 cos \u03b83 = 0 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001937_6.2007-1911-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001937_6.2007-1911-Figure5-1.png", + "caption": "Figure 5. Design domain and associated variables and parameters.", + "texts": [ + " Also included in Section A is a description of the method used to model geometric uncertainty. In Section B a description of how the force-displacement and stress relationships are modeled is provided. Sections C and D provide deterministic and non-deterministic optimization problem statements, respectively. This section presents the geometric model used to define the contact; a definition which is used during the optimization process. The two dimensional model depicts the contact from its side view in the design domain, which is shown as a shaded area in Fig. 5. The design domain, as defined by dimensions \u0393x and \u0393y, encloses the complete contact geometry, with the exception of the cam center, which is permitted to occupy any point in the two-dimensional space. Unlike the geometric model from our earlier work,4 the design domain model uses the Cartesian coordinates of each node (xi, yi) as variables for the design optimization. Because the nodal coordinates are permitted to be anywhere within the design domain, the optimization has nearly complete freedom to search for constant force configurations", + " Two nodes, however, must remain at fixed positions in order to preserve predetermined mating with other devices (i.e., a circuit board, another electrical device). These nodes are at the point (xo, yo) where the contact is fixed to the plastic housing, and at the point (xa, ya) at which the deflection (\u2206) is applied. The total number of nodes (nn) and elements (ne) in the design is chosen according to the desired resolution of the finite element model. Importantly, we note that the end of the cam link (xc, yc), in Fig. 5, is not restricted to be within the design domain. This is an additional freedom that is not found in our previous work.4 We simply use the cam link to simulate an actual cam that will be used in the physical implementation of our design. Therefore, the center of curvature for the cam is not required to be within the design domain. Figure 5 also illustrates variables that are used to formulate behaviorial constraints that ensure that the resulting design will be both manufacturable, and within the predictive capabilities of the simulated pin joint model. Specifically, we define the angle between two adjacent elements (\u03b1i), the length of each element (Li), the distance from each node to the cam center (Ri), the angle between the cam link and the preceding element (\u03b1c), and the angle between of the cam link as measured from the horizontal (\u03b8c)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000869_piee.1965.0259-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000869_piee.1965.0259-Figure5-1.png", + "caption": "Fig. 5 Example of determining the pole pitch of the meter's electromagnets a Determination of the length of the pole pitch b Showing disposition of the current poles with flux-density distribution in", + "texts": [ + " of rotation of the current flux O, /i = current in rotor induced by forward-travelling field 12 = current induced in rotor by backward-travelling field TQ = driving torque on the rotor, proportional to power measured by meter Tv \u2014 damping torque on rotor, proportional to square of voltage and to speed of rotation of rotor Tj = damping torque on rotor proportional to square of current and to speed of rotation of rotor V2,12, n) \u2014 torque acting on rotating rotor a = distance between outer edges of current poles {Fig. 7(i)} b = distance between outer edges of voltage electromagnet {Fig. 7(i)} c = distance between geometrical axes of the current poles {Fig. 7(i)} d \u2014 pole pitch of meter (Fig. 5a) p = parameter of parabola; i.e. coefficient of n2 in the expression for Pm Paper 4754 S, first received 13th April and in revised form 26th November 1964 Dr. Podemski is with the Technical University of Szczecin, Poland PROC. IEE, Vol. 112, No. 8, AUGUST 1965 1 q = parameter of parabola; i.e. coefficient of n in the expression for Pm n = speed of rotation of rotor, rev/min ns = synchronous speed of rotation of rotor, rev/min Introduction There are two theories for the operation of induction meters: the transformer or eddy-current theory, also known as Rogowski's theory, and the travelling-field theory", + " In the light of the preceding interpretation, the fallacy in the conclusion drawn from the travelling-field theory in References 1 and 2 becomes obvious, as it assumes that only a forward-rotating field exists in the meter, independently of the load-current/voltage ratio. The identification of the characteristic's maximum,3 as in Fig. 4, with the synchronous speed, provides additional material for the argument concerning the second point mentioned above. Indeed, in the papers quoted, with the meter pole pitch as in Fig. 5a, the synchronous speed is almost twice as small as the maximum speed of rotation of the disc rotor found in various meters. The latter fact, which essentially contradicts the travelling-field theory, would obviously prove that it is at least partly wrong, were it not for the fact that the definition of the pole pitch as in Fig. 5a is debatable. accordance with the given pole pitch 1597 If we consider that References 1 and 2 assume the correct pole pitch, the idealised space distribution of the magnetic flux density, due to the current electromagnet, in the air gap should be in accordance with Fig. 5b. It can be seen that the flux density should be greatest not only in the neighbourhood of the current poles but also between the poles. Moreover, the phase shift between the magnetic flux densities in the regions of the poles should be zero. This disagrees with the real situation existing in the air gap of the meter when the load current is switched on. This proves that the definition of the pole pitch in Fig. 5a is incorrect. As shown in the Introduction, both theories lead to the same mathematical expression for the torque acting on the revolving rotor. This prevents any possibility of deciding whether a travelling field exists by measuring the torque T(P, V2, I2, n) as a function of any one of the variables, or as a function of any other quantity determined by the torque. Clearly, the problem has to be solved by other means. It would seem to be essential to investigate whether the magnetic fields in the air gap of the meter, whose construction differs so greatly from that of the induction motor, conform to the assumptions of the travelling-field theory", + " 7(ii), considered within that interval. The latter interval differs little from the dimension b, and is equal to twice the distance between the axes of the current poles. Consequently, the correct measure of the pole pitch of a meter with a disc rotor should be taken as the distance between the geometrical axes of its current poles. From what has been said, it is once more obvious that determining the pole pitch using the distance between the geometrical axes of a current pole and a voltage pole, as in Fig. 5a, is incorrect. Assuming the longitudinal limits of the air gap to be as given by our preceding considerations, the curve of Fig. 7(ii) can be resolved into a portion lying within those limits and a portion lying beyond them {Fig. 7(iv)}. The portion of the curve which is essential, i.e. the one lying within the limits of the length of the air gap, is, in idealised form, a sinusoid, as in Fig. 7(iii). In the general case, it is a curve which can be represented by a Fourier series. Thus, within the longitudinal limits of the gap, the distribution of flux density due to the current poles conforms to the assumptions of the travelling-field theory" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002732_j.engfailanal.2021.105451-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002732_j.engfailanal.2021.105451-Figure7-1.png", + "caption": "Fig. 7. Free frequencies of the construction, the first oscillating mode, undercarriage, slewing platform and pylons of BWE SchRs 630.", + "texts": [], + "surrounding_texts": [ + "When it was determined that a model (and, in principle, laboratory) testing is possible and that the obtained results are applicable to the actual structure, it was decided to manufacture a sub-scaled (laboratory) model. Complete technical documentation and technology for model development have been developed. The material used for this structure was steel S355J2 + N. For practical reasons, the unification of thicknesses was performed. Plates whose thicknesses deviate by up to a maximum of 20% from the thicknesses obtained by a tenfold reduction of dimensions were used. The undercarriage and slewing platform with the lower part of the pylons in different phases of construction are shown in Fig. 11 and Fig. 12. The manufactured model has a weight of approximately 230 kg and a frame size of 1550x1650x950 mm. In order to facilitate the load input in the cut-off zone of the columns, a ribbed plate was introduced, which connects the structures of the cut-off pylons and enables further better load input. A. Petrovic\u0301 et al. Engineering Failure Analysis 126 (2021) 105451 A. Petrovic\u0301 et al. Engineering Failure Analysis 126 (2021) 105451 In the calculation of the laboratory model, a load correction in relation to sub-scaled structure was made. The load is reduced to forces in the axial directions (two horizontal in the plane of intersection in the X and Y axes and vertical in the Z direction). The results of the numerical analysis are given in Fig. 13. Comparing the corresponding places on the sub-scaled model and on the laboratory model, at first glance, the absolute coincidence of the stress field of these two models is noticeable, except that the problem of large stresses at the intersection line of pylons is solved by adding a plate that connects the pylons. As far as the displacement of the bearing points is concerned, the sub-scaled structure and A. Petrovic\u0301 et al. Engineering Failure Analysis 126 (2021) 105451 the laboratory model have almost the same displacements (see Figs. 14 and 15). A dynamic calculation of the laboratory model was performed. The results are similar to the results obtained on sub-scaled model. In our laboratory conditions, it was not possible to induce excitation of such a high frequency, so dynamic tests were abandoned. The dynamic calculation was used to once again identify the weak points of the structure. In limited laboratory conditions, it was not possible to cause multi-axis loading of the model, so it was decided to load the model with equal vertical forces (symmetrically). It was decided that the model is loaded only with vertical forces of 10 kN per pylon at the points of intersection of the vertical plates of pylons. Given the linearity of finite element calculations, this does not affect the research. The model is positioned in a rigid frame, which served to load the model. The vertical forces are caused by the hydraulic cylinders resting on the upper horizontal beam. Each of the performed measurements involved a gradual input of force using a hydraulic cylinder at two points of the model (symmetrically), with a step of 2 kN per individual input point, i.e. in a pair of 4 kN in total. The aim was to verify the numerical calculation of the laboratory model by experimenting on a physical (previously developed) laboratory model. By examining the physical model, it was finally confirmed that such structures can realistically be examined by model, while the numerical calculation only indicated that such structures, in theory, can also be examined by model. The experimental methods used are strain gauges method and the method for non-contact measurement of stresses and strains. The symmetrical load enabled (in addition to the already existing symmetry of the structure itself) parallel measurement with strain gauges and an optical system for digital image correlation. Testing of the construction by the method of strain gauges, implies positioning of strain gauges at previously selected critical places, with the aim of mapping the model. As a problem with the method of strain gauges, which is most often used for diagnostics, the positioning of gauge itself is imposed. The gauge should be placed in zones where there is a large A. Petrovic\u0301 et al. Engineering Failure Analysis 126 (2021) 105451 stress change gradient, so just locating the strain gauge has a big impact on the results. This problem is overcome by using the method of digital image correlation, which enables the recording of stress over the entire field. The system for non-contact optical measurement of displacement and deformation consists of sets of stereo cameras and lenses. The system also consists of a stand, a power supply and a PC system with Aramis software installed. The operation of the system is based on digital image correlation, which compares the position of points before and after deformation. The stochastic pattern, applied to the surface of the object to be deformed (irregular black dots on a white background), allows the system to identify the points (pixels) that it tracks during deformation. This system can measure displacements up to 0.001 mm. Thus, method of strain gauges and method for non-contact measurement of displacement and deformation would confirm each other, and the overall information obtained is more complete than if only one method was used. The readings of strain gauges during one measurement are shown in Fig. 16. From the readings of strain gauges, it can be seen that the T5, T6, T7 and T8 results, although they follow the load of the model, are unusable, because the values of the measured elongations are within the measurement error. This was expected based on the results of the calculation, because due to the limitations of the experiment, a reduced vertical load was adopted, which is not of sufficient intensity to revive all the gauges. Strain gauge T5 is positioned on upper horizontal plate of undercarriage, where vertical plate connects the horizontal plate. Strain gauge T8 is positioned on the upper horizontal plate of slewing platform where vertical plate connects the horizontal plate. Rosette (T1-T3) is positioned on the vertical plate of undercarriage, near the outer cylinder, in the zone of stress concentration. Fig. 17 shows the position of the rosette. The results obtained using strain gauges T1-T3 are compared with results obtained using system for optical non-contact measurement, and this group of the results is named MM 1\u20133. The position of the recording equipment during measurement conducted in spot MM 1\u20133 is shown in Fig. 18. A line in stress field obtained using DIC that imitates a strain gauge is shown in Fig. 19. Namely, the software provides the ability to determine the distance between any two points at any time. Thus, line is drawn along in the vertical direction in the rosette zone, which practically imitates the T2 strain gauge. According to the table, a high matching between experimental results and numerical calculation can be noticed. It can also be noticed that everything that can be measured with strain gauges, can also be measured with a system for non-contact measurement of A. Petrovic\u0301 et al. Engineering Failure Analysis 126 (2021) 105451 displacement and deformation, faster and more efficiently. Strain gauges T6, T7 and T9 are positioned as shown in Fig. 20. The measurement results at MM 6 are given below. In Fig. 21, a field of vertical displacements that is fully aligned with the field obtained by the numerical calculation can be seen. This is the point around which the pylons bend slewing platform. The measurement results at MM 9 are given below. At this point, a considerable difference between the numerical calculation and the experimental results was observed. The reason for this are the massive welded joints near this place, because the cylinder, the upper plate of slewing platform and the knot plate of sprit join there. The welded joint, on the one hand, increases the stiffness of this place, and on the other hand, it can cause the A. Petrovic\u0301 et al. Engineering Failure Analysis 126 (2021) 105451 A. Petrovic\u0301 et al. Engineering Failure Analysis 126 (2021) 105451 concentration of stress entered by welding. The results of the experiment confirm the numerical calculation of the laboratory model. By confirming that the laboratory computational model is good, all computational models that preceded it were indirectly confirmed." + ] + }, + { + "image_filename": "designv11_83_0001082_00952443070067387-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001082_00952443070067387-Figure6-1.png", + "caption": "FIGURE 6. (a) FE model of rubber component and (b) predictions of contact pressure.", + "texts": [ + " To increase its efficiency, a computational technique of multistep frictional analysis proposed in Ref. [6] is also utilized. The only difference between analytical and FEA implementations is associated with a number of possible contact statements. Conditions (7) should be satisfied in FEA for every contact element to achieve robust convergence. Within the same analysis, parameters $ for different elements or contacts segments could be different or the same according to user\u2019s choice. As an example of the implementation, consider a plane-strain model shown in Figure 6(a). This is a rotated linearly-elastic thick ring at absolutely rigid internal sheave. The ring is compressed by a flat rigid surface. Friction between the ring and the surface is defined by very aggressive COF\u00bc 1.0. (Such high friction was selected to demonstrate the robustness of the proposed approach and developed tools.) Other characteristics of the problem are: R0\u00bc 1.0; Ri\u00bc 0.3; \u00bc 0.1; E\u00bc 1; and \u00bc 0.35. A very refined uniform mesh (Figure 6(a)) with 60 and 400 elements in radial and circumferential directions, respectively, is used to avoid accuracy issues. The predicted distributions of contact pressure are shown in Figure 6(b). The corresponding distributions of displacements U1 and von Mises stresses are shown in Figure 7 as additional representative examples. No convergence challenges are observed in spite of significant friction (COF\u00bc 1.0). They may be considered as additional indications of efficiency of the approach. 1. Convergence of contact analysis can depend on frictional properties if coupling iterative algorithms between normal and shear stresses are used. 2. A simple criterion of convergence is developed for iterative frictional analysis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000332_50003-4-Figure3.10-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000332_50003-4-Figure3.10-1.png", + "caption": "Fig. 3.10. Equivalent side slip angle producing the same pneumatic trail t.", + "texts": [ + " t F { 3 0 a - 3(00) 2 + (00\") 3 } for O\" _ O'sl (3.48) F - p F for tr___ %t and obviously follows the same course as those shown in Fig.3.8. The force vector F acts in a direction opposite to V~ or -tr. Hence F _ F ~r -~ (3.49) from which the components Fx and Fy may be obtained. The moment - M z is obtained by multiplication of Fy with the pneumatic trail t. This trail is easily found when we realise that the deflection distribution over the contact length is identical with the case of pure side slip if tanaeq - ~ (cf. Fig. 3.10). Consequently, the formula (3.13) represents the pneumatic trail at combined slip as well if Oy~ry is replaced by 0a. We have with (3.13): M z - - t ( t r ) ' F (3.50) In Figs.3.11 and 3.12 the dramatic reduction of the pure slip forces (the side force and the longitudinal force respectively) that occurs as a result of the simultaneous introduction of the other slip component (the longitudinal slip and the side slip respectively) have been indicated. We observe an (almost) symmetric shape of these interaction curves" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002560_tec.2021.3069096-Figure13-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002560_tec.2021.3069096-Figure13-1.png", + "caption": "Fig. 13. 3D FEA Model of the Prototype machine.", + "texts": [ + " structure. As shown in Fig. 12, the bubble color change measured by air-gap K demonstrate its effect on position signal quality. It can be observed from the diagram that for the best rotor end designs (design points near the Pareto front), the value of K is small ranging from 1.0-1.3 where the pole tooth top contours are similar to circular arc. This is different from that of the VR resolvers. Since the rotor pole is designed with slots and teeth, the influence of the K factor is smaller. As shown in Fig. 13, the prototype machine with the RB is modeled in FEA software for 3D analysis. It is seen that the copper coils protrude from the stator iron core and are connected to an external circuit for forming up the stator winding. An air box is used to contain the machine model. The end winding part of the machine is not included as its performance is not the scope of this paper. One big concern of the PM rotor is the axial flux distribution across the contacting surface between 2 rotor parts . The paths of potential flux conduction from one magnet pole to another have access to the rotor end" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003345_ur52253.2021.9494692-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003345_ur52253.2021.9494692-Figure2-1.png", + "caption": "Figure 2. CAD file of the whole system includes; XY crane, Robot manipulator, and the circular RVI environment.", + "texts": [], + "surrounding_texts": [ + "In this section, the proposed control scheme has been discussed in detail. Nonlinear ESO has briefly explained in the first part, 2nd part describes the ISMC and the 3rd explained the integration of ISMC with ESO." + ] + }, + { + "image_filename": "designv11_83_0001802_icnc.2007.124-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001802_icnc.2007.124-Figure4-1.png", + "caption": "Figure 4. Schematic diagram of inverted pendulum system", + "texts": [ + " From the two figures it can be seen that the proposed method is superior to the ENN method. 0 20 40 60 80 100 -0.02 0.00 0.02 0.04 0.06 0.08 k Elman error curve TDRNN error curve The inverted pendulum system is one of the classical examples used in many experiments, and it is often used to test the effectiveness of different controlling schemes [7]. So in this paper, to examine the effectiveness of the TDRNN model, we investigate the application of the TDRNN to the control of inverted pendulums. The inverted pendulum system used here is shown in Fig.4, which is formed from a cart, a pendulum and a rail for defining position of cart. The Pendulum is hinged on the center of the top surface of the cart and can rotate around the pivot in the same vertical plane with the rail. The cart can move right or left on the rail freely. The dynamic equation of the inverted pendulum system can be expressed as l Mm ml ml MmxmlFg p c \u22c5 + \u2212 \u2212+\u22c5+\u2212\u2212+ = \u2212 \u03d5 \u03d5\u03bc \u03bc\u03d5\u03d5\u03d5\u03d5 \u03d5 2 12 cos 3 4 ])()sgn(sin[cossin & && && (38) Mm xmlF x c + \u2212\u2212+ = )sgn()cossin( 2 &&&& && \u03bc\u03d5\u03d5\u03d5\u03d5 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002473_tmag.2021.3068131-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002473_tmag.2021.3068131-Figure2-1.png", + "caption": "Fig. 2. Coupled 2-D\u20131-D FEA to consider laminated cores.", + "texts": [ + " The variation in H by the generation of minor hysteresis loops can also be approximately estimated by the combination of major loops expressed by (1) whose amplitudes correspond to the local minimum and maximum of B [6]. Then, the following governing equation, which includes both the effects of hysteresis phenomenon and eddy currents in laminated cores, is solved by using coupled 2-D and 1-D FEA: \u2207 \u00d7 (\u03bd\u2207 \u00d7 A) = \u2207 \u00d7 Heddy + \u2207 \u00d7 Hhys = \u2207 \u00d7 H reac (2) where A is the magnetic vector potential, Heddy, Hhys, and Hreac are the eddy currents, hysteresis, and total reaction fields, respectively. Fig. 2 shows the method to determine Hreac [4]. First, the nonlinear time-stepping 2-D FEA (main analysis) is carried out and the calculated flux density vector B is decomposed 0018-9464 \u00a9 2021 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See https://www.ieee.org/publications/rights/index.html for more information. Authorized licensed use limited to: Robert Gordon University. Downloaded on May 28,2021 at 01:52:26 UTC from IEEE Xplore. Restrictions apply. into L- and S-axes components (L-axis is defined to be the direction of B when |B| becomes maximum and the S-axis is orthogonal to the L-axis)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003527_s00707-021-03025-1-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003527_s00707-021-03025-1-Figure3-1.png", + "caption": "Fig. 3 Translation of the GRFs to the ankles", + "texts": [ + " Here, we can define a PD controller based on the CoM and momentum states so that it realizes the desired position and velocity of the CoM:{ L\u0307d K p1(cc,d \u2212 cG) + K d1(c\u0307c,d \u2212 c\u0307G) A\u0307d K p2(Ac,d \u2212 A) + K d2( A\u0307c,d \u2212 A\u0307) (9) where cc,d and c\u0307 c,d are the desired CoM location and velocity, respectively, and Ac,d and A\u0307 c,d are desired CoM momentum states. The feet of the robot occupy a small part of the robot\u2019s overall mass, so the mass of the robot\u2019s feet can be ignored. In this view, the GRF can be regarded as the force and moment applied to the ankles as shown in Fig. 3. The advantage is that we can control the angular momentum change rate of the robot by determining the positions, forces, and torques of the feet. Also, it can be easy to ensure the CoP in the interior of the foot\u2019s support polygon by controlling the ankle torque. And the momentum change rates can be expressed as follows: L\u0307 f r + f l , A\u0307 A\u0307 f + A\u0307\u03c4 , A\u0307 f (cr \u2212 cG) \u00d7 f r + (cl \u2212 cG) \u00d7 f l , A\u0307\u03c4 \u03c4 r + \u03c4 l (10) where f r and f l are the GRFs acting on the right and left foot, respectively, cr and cl represent the positions of the ankles, and \u03c4r and \u03c4l are the torques of the right and left ankles" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002772_978-3-030-67750-3_14-Figure28.7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002772_978-3-030-67750-3_14-Figure28.7-1.png", + "caption": "Fig. 28.7 Outlet 2\u2014pressure reading of Mode 2 nozzle", + "texts": [ + " The simulation was implemented to all nine steps in order to find the impact toward the outlet 2 pressure and velocity that affected the spray angle and distance. Figure 28.9 presents the pressure curve that is affected by the length of the nozzle. It clearly can be seen that the pressure at outlet 2 is inversely proportional to the nozzle length. As the nozzle length increases, the pressure will be decreased. The graph pattern is in a gradually decreasing manner. This is because, as the powder goes through along the nozzle, the pressure will continue to decrease due to the loss of potential energy. Figure 28.7 shows the pressure reading during Mode 2 nozzle was set. While, the velocity reading of Mode 2 nozzle set as illustrated in Fig. 28.8. Presented in Fig. 28.10 is the velocity of the outlet 2 against the nozzle length. It shows that as the length of the nozzle increases the velocity happening on outlet two also increases. This is because the velocity is directly proportional to the length of the nozzle. Based on this graph, it is concluded that the higher the length and the diameter of the outlet, the faster the velocity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001291_amr.33-37.1011-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001291_amr.33-37.1011-Figure7-1.png", + "caption": "Fig. 7 shows the final bone forming under compressed loading condition with different bone keeping value as Vk=60%, Vk=40%, Vk=30% and Vk=25%. The compressed forces are set 300N in all cases. The results shown, that the bone formed largely changed with Fig.5 and Fig.6, and bone shape as a pipe with hole in all models. There is some unwanted bone stack on it too in all of the cases, and the same results of with Fig. 5 and Fig.6 are observed as indefinitely unwanted bone. The unwanted bone stacks gradually decreased by decreasing the bone keeping volume from 60% to 30% and 25%, and very smoothed pipe much close to femur.", + "texts": [ + "5 results and show solid pillar. There is much unwanted bone stack on it too in all of the cases. Fx Fy F 1014 Advances in Fracture and Materials Behavior The unwanted bone positions not defined and go round and round along the pipe wall. The unwanted bone stacks gradually decreased by decreasing the bone keeping volume from 60% to 30% and nearly close to femur, in 25% close to fibula. Vk=60% Vk=40% Vk=30% Vk=25% Fig.6. Bone Final Formation under Compressed Load (F=200N) in Different Bone Keeping Volume Fig.7. Bone Final Formation under Compressed Load (F=300N) in Different Bone Keeping Volume The bone mass and space volume are increased with increase of the initial load value in a same bone keeping value, then fibula and femur bones are obtained respectively by keeping the required bone forming value. The different bone shapes are obtained by changing the both bone keeping value and the compressing force value. When set larger bone keeping value by keeping larger constant compressing force value, bone shape as a pipe with hole just like femur and largely increasing its space volume, when set smaller bone keeping value by keeping the smaller constant compressing force value, it is close to solid pillar as like fibula and decrease its space volume" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003296_j.mechmachtheory.2021.104440-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003296_j.mechmachtheory.2021.104440-Figure6-1.png", + "caption": "Fig. 6. Connection of the spur gear pair pinion and wheel via the gear mesh", + "texts": [ + " 5 shows the principle by which this is carried out, for a smaller example shaft with two disks, two bearings and three flexible elements. Node vectors {q} and arbitrary component property matrices [A]nxn are shown (Fig. 5a). The matrices are summed by moving left-to-right through the system (Fig. 5b). The final step in the construction of the full system is to insert the gear mesh to couple the two shafts. The gear mesh is represented B. Friskney et al. Mechanism and Machine Theory 165 (2021) 104440 by a nonlinear spring in parallel with a viscous damper (as shown in Fig. 6). For the purposes of the eigenvalue problem, only the mean stiffness term is used, in a similar fashion to the rolling element bearings. The terms of the gear mesh stiffness matrix are derived from the loads acting upon the pinion and wheel as functions of the dynamic transmission error (DTE), as detailed by Rao et al. [28] The mesh damping matrix follows the same construction [34]. As the pinion and wheel masses and inertias appear in the mass and gyroscopic matrices of the disks, only the stiffness and damping of the gear mesh are required to couple the shafts. As the matrices of the pinion and gear act between the nodes where they are situated (i and j in Fig. 6), their order is 10. Upon reaching the right-hand end of the input shaft (Fig. 2), the node numbering for the full system continues sequentially at the left end of the output shaft. As such, the mass, stiffness, gyroscopic and damping matrices of the assembled shafts are joined diagonally. B. Friskney et al. Mechanism and Machine Theory 165 (2021) 104440 The 10 \u00d7 10 gear mesh stiffness and damping matrices of Eq. (2) are split into four 5 \u00d7 5 cells and placed in the system stiffness and damping matrices at the locations corresponding to the nodes i and j of the engaged gear pair\u2019s pinion and wheel (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001585_icelmach.2008.4799976-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001585_icelmach.2008.4799976-Figure5-1.png", + "caption": "Fig. 5. Field distribution at rated load.", + "texts": [ + " The resistance of the stator end winding was measured with an ohmmeter and the end-ring segment between bars was measured indirectly with a resistivity meter. Values obtained are shown in Table I. Simulations of three scenarios were carried out; they are the no-load, locked rotor and load tests. A dc generator was used as a load of the induction motor, and measurements of speed, line current, voltages and power were made. Fig. 3 shows the field distribution in the locked rotor test, where it can be seen the fields are located on the rotor surface due to the induced currents in the rotor bars. Fig. 5 shows that under rated load condition the electromagnetic fields penetrate into the rotor because of the slip, and therefore the frequency of the pulsating electromagnetic field is smaller than that generated in the locked rotor test. Table II shows the comparison of simulation results against experiments of the no-load test, where the phase current is over-predicted by 0.03A, active and reactive power show discrepancies. In the locked-rotor test, the FE model underpredicts the phase current by 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001772_pes.2008.4596250-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001772_pes.2008.4596250-Figure3-1.png", + "caption": "Fig. 3 Construction of Rotor", + "texts": [ + " Otaka and M. Fujita are with Toshiba Corporation, 2-4 Suehiro-cho, Tsurumi-ku, Yokohama 230-0045 Japan. ** K. Nagasaka is with Tokyo University of Agriculture and Technology, 2-24-16 Nakamachi Koganei Tokyo 184-8588 Japan. Considering the above situation, we studied the equivalent circuit model using x\u2019q and focused on the generator transient and dynamic characteristics. The results of the study are reported below. The rotor structure of a large capacity cylindrical type synchronous machine is shown in Fig. 3. As the figure shows, the rotor consists of a solid iron core. Therefore, the surface of the shaft and wedges carry currents during a transient state, etc. We chose two generators (500,000kVA class and 900,000kVA class with cylindrical solid rotors) that were used for measuring x\u2019q and we conducted an analysis assuming that the reactance and resistance components are constant. . In addition to the step response test and sudden load application test, the load rejection test is widely used for A Study on Quadrature Equivalent Circuit Model in Large Synchronous Machine Daisuke Hiramatsu*, Yoichi Uemura*, Jyunji Okumoto*, Shinji Uemoto*, Takehiko Imai*, Mikio Kakiuchi*, Ken Nagakura*, Toru Otaka*, Masafumi Fujita*, Ken Nagasaka** T \u00a92008 IEEE" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002641_j.robot.2021.103783-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002641_j.robot.2021.103783-Figure4-1.png", + "caption": "Fig. 4. Overall mechanical structure of the actuated limb.", + "texts": [ + " Moreover, two ends of each RPR limb are connected to the frame and the moving platform through the joint bearings, and each linear driving unit adopts a servo electric cylinder integrating the servomotor and lead screw. Furthermore, the corresponding effective stroke, rated output force and the accuracy of the output displacement are 250 mm, 800 N and 0.001 mm, respectively. c Since the elastic deformation of the cylinder and putter of the electric cylinder are too small, it is a challenge to analyze the influence of the limb stiffness on the driving forces distribution. Therefore, Fig. 4 illustrates that a flexible P joint with straight beam and parallel plate is connected in serial at the end of putter of the electric cylinder within each RPR limb. The flexible P joint has much smaller stiffness in the axial direction than those in other directions, so it is assumed here that only the elastic deformation in the axial direction is considered. Moreover, two kinds of flexible P joints with different thicknesses of t = 1 mm nd t = 2 mm are designed. In order to detect driving forces exrted to the moving platform provided by the non-redundant and edundant limbs, a tension\u2013compression force sensor is installed etween the flexible P joint and the moving platform in each ctuated limb. Furthermore, Fig. 4 shows the specific structure f the actuated limbs. .2. Control system Fig. 5 shows the overall structure of control system of the RPR + P redundantly actuated PM, the form of \u2018\u2018PC + motion ontrol card\u2019\u2019 is adopted, and the motion control card uses the MAC multi-axis motion controller from Delta Tau Company of nited States. After choosing other electrical components, make reasonable layout and design the electric control cabinet, and hen complete the wiring in accordance with the electrical wiring iagram, the final experimental platform is shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001726_isam.2007.4288440-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001726_isam.2007.4288440-Figure8-1.png", + "caption": "Fig. 8. Example 2", + "texts": [ + " The length of the object U is equal to 1, its width V is equal to 0.1, and its flexural rigidity along the spine line Rf is constantly equal to 1. In this example, both ends of the spine line are on the same line but directions of the spine line at these points are different. Fig.7 shows computational results. Fig.7-(a), -(b), and -(c) illustrate the top, front, and side view of the object, respectively. As shown in this figure, the object satisfies the given geometric constraints by twisting partially. This computation took about 1,500 seconds. Fig.8 shows the second example of L-shaped object deformation. The original shape of the object is illustrated in Fig.8-(a). It is 2 long, 0.2 wide, and it has one rectangular bend on its mid point, namely, uh = 0.5U and \u03bb = \u03c0/2. Its flexural rigidity Rf is constantly equal to 1. Positional and orientational constraints are shown in Fig.8-(b). Fig.9 shows computational results. As shown in this figure, the object is not also bent but also twisted and its shape becomes asymmetrical. This computation took about 1,500 seconds. Thus, our method can estimate bending and torsional deformation of a rectangular and L-shaped belt object using only flexural rigidity of the object along its spine line if the object is isotropic. This flexural rigidity can be measured by a simple experiment. In this section, the computation results will be experimentally verified by measuring the deformed shape of a belt object" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003039_icit46573.2021.9453629-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003039_icit46573.2021.9453629-Figure2-1.png", + "caption": "Fig. 2. Diagram of the muscle-strength-assistance method", + "texts": [ + " The device is fixed with three chest fixation belts and a newly developed lumbar fixation mechanism equipped with a leaf spring. The length and mounting position of each part of the device can be changed to suit various body sizes. The entire assist suit, including the artificial muscle, amplification mechanism, leaf spring, and lumbar fixation mechanism, weighs 3.3 kg. AB-Wear III performs effectively by simultaneously assisting both muscles and postures. 1) Muscle-strength-assistance method: Fig. 2 shows the muscle-strength-assistance method. The two artificial muscles provide assistance by exerting a force almost parallel to the spine, behaving in a way similar to the erector spinae muscles. The balloon actuator expands when air is applied, and expansion force is transmitted to a wearer\u2019s body through the artificial muscles. The angle of an artificial muscle with respect to the back varies with the level of expansion of the balloon actuator, and length of the moment arm around the lumbosacral joint can be increased" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001045_12.763341-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001045_12.763341-Figure1-1.png", + "caption": "Fig. 1. (a) Simplified illustration of the self-assembly technique for the fabrication of vertical inductors. (b) SEM image of vertical inductors fabricated as described in (a). (images reprinted from Dahlmann et al14)", + "texts": [ + " Such approaches include dielectric or metal shielding between inductor and substrate3-5, removal of the substrate at the inductor region6-8 or using alternative fabrication techniques to achieve three-dimensional structures such as solenoids9-12. Typically, such techniques offer limited suppression of the substrate effects, or introduce non-standard fabrication techniques that are difficult to integrate with the CMOS process. In previous work, a self-assembly technique for the fabrication of vertical inductors has been proposed. Surface tension forces are used to rotate a planar inductor off the substrate and bring it to a vertical orientation13. An illustration of this technique, reprinted from Dahlmann et al14 is given in Fig. 1a. An SEM image of two such vertical inductors is given in Fig. 1b14. Although this technique was successful, its applicability to commercial fabrication had some particular limitations. Firstly, the inductor material was copper which can be easily oxidized. Since the structure is not *m.kiziroglou@ic.ac.uk; phone +44 20 7594 6216; fax +44 20 7594 6308; Micromachining and Microfabrication Process Technology XIII, edited by Mary-Ann Maher, Jung-Chih Chiao, Paul J. Resnick, Proc. of SPIE Vol. 6882, 68820B, (2008) \u00b7 0277-786X/08/$18 \u00b7 doi: 10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001431_12.774989-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001431_12.774989-Figure5-1.png", + "caption": "Fig. 5 Analysis of static stability", + "texts": [ + " The stability of a wheeled rover which is stationary or moving at a constant speed on a slope is expressed in terms of the gravitational stability margin. If the rover is driving parallel to a downhill slope the gravitational stability margin is the margin of longitudinal stability, and if it drives along a cross-hill slope it is the lateral stability margin, respectively. Without loss of generality, it is assumed that the radius of lower wheel is R, and the slope incline angle is \u03b8. As shown in Fig. 5, the gravitational stability margin \u03b4 can be estimated from equation: 2 cos[ sin( )] sinC A R rX X L arc R L \u03b4 \u03b1 \u03b8 \u03b8\u2212 = \u2212 = + \u2212 \u2212 (3) Where XA is projection of contact point of the lower wheel on the ground, XC projection of the center of gravity on the ground, the other variables refer to Fig. 6, the rover configuration parameters: \u03b1, L, L2 are definite, for the rover on a cross-hill slope, the stability margin equation is similarity with \u03b1, L, L2 instead of cross-section parameters. Proc. of SPIE Vol" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003132_978-981-16-0119-4_20-Figure78.10-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003132_978-981-16-0119-4_20-Figure78.10-1.png", + "caption": "Fig. 78.10 Pressure cum drag force estimations on propeller equipped with saw tooth cut at 1-LE", + "texts": [], + "surrounding_texts": [ + "numerical results are shown in Figs. 78.12 and 78.13. The drag and exit velocities are considered as important selection parameters so the streamlines and planes are used for the successful capturization of aforementioned performance parameters. From the comparative Figs. 78.12 and 78.13, it has been understood that the saw tooth cuts loaded on the edges of the UAV\u2019s propeller are performed better than base propeller. From this comparative, CFD analyses are predicted that the saw tooth cuts-based edge modification techniques have the full capability to provide good" + ] + }, + { + "image_filename": "designv11_83_0000737_ijtc2007-44228-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000737_ijtc2007-44228-Figure1-1.png", + "caption": "Figure 1. A deep groove bearing with a sputtered AlN piezoelectric film on the outer raceway.", + "texts": [ + " The aluminium target was pre-sputtered under vacuum to remove the surface oxide. Nitrogen gas was then fed into the chamber and a film of AlN deposited on bearing outer raceway. A coating thickness of around 4 \u00b5m was achieved after several hours of sputtering. The substrate needed no preheating so there was no danger of the bearing steel tempering. SEM examinations of the coating demonstrate that the films have a highly columnar structure. This gives them a strong piezo-electric property. More details of the coating process are given in [4]. Figure 1 shows a conventional deep groove ball bearing (type 6016, shaft diameter 80 mm, ball diameter 12.7 mm). A 4 \u00b5m thick layer of aluminium nitride has been sputtered onto the outer raceway on the outer bore. A rectangle of silver paint has been applied to the coated surface as the top electrode. The bearing outer race is the bottom electrode. Wires are soldered 1 Copyright \u00a9 2007 by ASME ms of Use: http://www.asme.org/about-asme/terms-of-use Dow to these electrodes and then to a tag bonded to the bearing for durability. In the photograph (figure 1) the coating is difficult to see but the soldering tag and wired are visible. Figure 2 is a schematic of the instrumented bearing showing the sensor location and detail of the sputtered coating and upper electrode. The sensor has been coated onto the outer surface of the outer raceway. The electrode has dimensions of 0.3 mm by 3 mm. The high frequency ultrasonic wave has the property that it tends not disperse greatly. The size of the contact patch during these tests is an ellipse varying from 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000602_978-3-540-74764-2_52-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000602_978-3-540-74764-2_52-Figure3-1.png", + "caption": "Fig. 3. Micromanipulation module with rotational actuator (\u00d8 9.2 mm) mounted on a printed circuit board with a tool (gripper) integrated directly onto the rotor (left). Micromanipulation module without tool (centre). The rotational actuator (right).", + "texts": [ + " The scanner is made of 4 PZT stack actuators that permits movements with 3 DOF (x,y,z) and the cantilever contains a force sensor based on a piezoresistance (see Fig. 2). The AFM tool consists of three main components: 1.) the AFM probe with the integrated piezoresistance, 2.) an AFM holder for easy probe exchange and 3.) the XYZ scan stage. A rotational drive has been developed, evaluated and redesigned to be a suitable interface between robot and tool. Five rotational drives with specific rotors for each tool have been built (see Fig. 3). Furthermore, the rapid prototyping technique for making multilayer piezoceramics, developed in the beginning of the project, has been utilised to make the drivers for the grippers. The maximum torque of the motor is 80\u00b5N at a drive voltage of 50 V and a spring force of 1.2 N. The rotational actuator has extremely good motion resolution (0, 1\u00b5rad) for driving frequencies up to about 80 Hz (0.1 rpm). Fast transport can be achieved in the frequency range 3\u20136 kHz (4 rpm). The power consumption is about 1 mW and 80 mW respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002772_978-3-030-67750-3_14-Figure24.1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002772_978-3-030-67750-3_14-Figure24.1-1.png", + "caption": "Fig. 24.1 3D DIC (stereo vision) cameras set up at a measured distance relative to the position of specimen", + "texts": [ + " Random speckling on specimen surface are done using matte black and white spray paints. A white matte layer serves as a contrasting background and random tiny black speckles pattern created using black spray paint is uniformly distributed on the entire specimen. The punch test rig consists of die sets to secure the 90 mm circular specimen blanks. The die sets have clearance in the centre to allow specimens to deform when loaded with a 20 mm diameter conical punch. The punch test rig is set up together with DIC cameras on a 100 kN Electric Mayes machine as shown in Fig. 24.1. The general method of assembling the whole rig starts with the clamping of the blanks. The blank is secured in between the dies and tightened before being put on the punch. The die set is supported with a pair of stands from the bottom for a little clearance between the punch tip and the specimen to avoid the punch from damaging the specimen before the start of the experiment. Supporting bars are then placed on top of the die set to provide a hold down force during the test and is tightly bolted at the connection which secures the position of both top supporting bar and bottom supporting bar of the Mayes machine to secure the die set", + " The 3D DIC set up is done by mounting cameras on the boom at the top of the test rig and connecting both cameras to the computer to view the area of interest on the specimen. On the computer, the software being used for the pattern recognition is based on the principles of digital image correlation is the VIC-3D by Correlated Solution. Every pair of images continuously captured by both cameras is processed into stereo images. Cameras orientation is adjusted until the area of interest is at the centre of camera views, then the best focal length is found by adjusting the lenses. Figure 24.1 shows the measurements of cameras set up relative to the specimen. Cameras are adjusted until they are in focus at the same location on the specimen preferably at the centre region. Examples of speckle images with a good focus are shown in Fig. 24.2. Once both cameras are in focus, the calibration process is carried out by using a pre-determined dot-pattern calibration plate. At least 25 images are captured for the calibration process at which later only 15 images with good calibration score are chosen to make sure the calibration error between the two cameras is below 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000936_speedham.2008.4581270-Figure16-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000936_speedham.2008.4581270-Figure16-1.png", + "caption": "Fig. 16 \u2013 3D scheme of vehicle AnsaldoBreda Sirio Milano 25m", + "texts": [], + "surrounding_texts": [ + "The effects of proposed sampling algorithms have been evaluated using a Train System Simulator developed in Ansaldobreda. The Train System Simulator is a software simulator that allows compiling the control software (the same compiled for the real control board and used on the vehicle) in a C++ environment for PC. Each physical element (control and interface boards, propulsion system, asynchronous motor, main line, wheel-track adhesion, . . .) is associated to a software object. Such tool is used both for the software unit test and for the system test. The simulation method utilized is Euler with integration step of 1 \u03bcs, sampling rate of 10 \u03bcs and 1 s of length. The sampling algorithms have been successively tested on the vehicle AnsaldoBreda Sirio Milano. Sirio Tram is a Light Railway Vehicle provided of two motor bogies for a total of four asynchronous motors. The vehicle can be unidirectional or bi-directional and it can have one or two trailer bogies to satisfy the length requirements of customers. The tram on which the simulations and the test have been carried out (Sirio Milano), has one trailer bogie, is 25 m long and its main features are reported in Table I. For each motor bogie there are two propulsion systems, each one made of one traction inverter and one braking chopper arranged on the roof ." + ] + }, + { + "image_filename": "designv11_83_0002023_iecon.2008.4758385-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002023_iecon.2008.4758385-Figure1-1.png", + "caption": "Fig. 1. Iso-torque optimum", + "texts": [ + " Here the extended criterion to be minimized is: (6) The derivation by , , respectively gives: 2 (7) 2 (8) 2 (9) Replacing i , i and i in (5) and (4) gives: 2 3 \u2013 (10) 6 3 \u2013 (11) These expressions are introduced in (7) and lead to: 2 2 2 \u2013 2 \u2013 2 \u2013 2 (12) We find the expression: (13) The demonstration is the same for and . B. Iso-curve method In this part, we will take a look at how to find the reference current expression with homopolar component (1). A simple way is to find it with the iso torque curve method. The isotorque curves are defined by the set of points M giving a constant electromagnetic torque for an angle \u03b8 (see Fig. 1). In order to minimize the copper losses it is necessary to choose the point on the curve which gives an vector as small as possible [4]. Equation (4) has the following expression in the \u2010frame: e i e i (14) Graphic presentation to find the optimum: e sin \u03b1 (15) sin \u03b1 e e e (16) We can deduce from 15 and 16 , therefore: i sin \u03b1 e sin \u03b1 e e e (17) The expression of i\u03b2 is: i e e e (18) The relations between i , i , i , i\u03b1, i\u03b2, e , e , e , e\u03b1 and e\u03b2 are: i i i i\u03b1 i\u03b2 (19) e e e e\u03b1 e\u03b2 (20) i i i T\u03a9 e\u03b1 e\u03b2 e\u03b1 e\u03b2 (21) and e\u03b1 e\u03b2 e\u03b1 e\u03b2 e\u03b1 e\u03b2 (22) e\u03b1 e\u03b2 e e e e e e (23) e\u03b1 e\u03b2 e e e (24) Consequently the relation between the currents and back emf using homopolar component is: i i i e e e e e e (25) We consider the following case; the PMSM runs with three sinusoidal back emf: e e e \u03a8 \u03c9 cos \u03c9t \u03a8 \u03c9 cos \u03c9t 2\u03c0 3 \u03a8 \u03c9 cos \u03c9t 2\u03c0 3 \u03a8 \u03c9 2 3 T P \u03c9t 1 0 (26) e e e 3 2 \u03a8 \u03c9 (27) According to (1), the calculation of optimum currents gives: i i i 3 2 \u03a8 cos cos 2 3 cos 2 3 (28) In the sinusoidal behavior the reference currents which have been found have the classical form, in phases with the back emf" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001000_ijtc2007-44403-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001000_ijtc2007-44403-Figure1-1.png", + "caption": "Figure 1. Exploded view of the Finger seal Assembly.", + "texts": [ + "asmedigitalcollection.asme.org/ on 03/28/2018 Ter brush seal, Proctor [5]. New and improved designs were later offered by Arora [6], and Proctor [7]. The designs proposed in [6] and [7] are adding pads to the end of the finger stick. These pads are meant to create experience a hydrodynamic lifting force (due to the moving surface of the rotor) that still allows sealing, but in a non-contacting manner. The complete geometry of a two-layer, 215.9 mm (8.5in) diameter finger seal is illustrated in Figure 1. Differently from Arora\u2019s [6], only the low pressure finger seal laminate contains pads meant to provide aerodynamic lifting. The details of this geometry have been further detailed in Braun et al. [8, 9]. Figure 1 presents a three-dimensional exploded view of the full solid model. It contains the back-plate, the front plate as well as the low-(LP) and high-pressure (HP) finger assembly laminates. The backplate serves as an axial support against bending when the thin, HP and LP finger laminates are under pressure. This plate also incorporates a pressure equalization manifold that allows mitigation of the Coulomb friction between the plate and the LP laminate. Figures 2a and 2b show the details of the staggered relative position of the HP laminates 1 Copyright \u00a9 2007 by ASME ms of Use: http://www" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002752_012051-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002752_012051-Figure1-1.png", + "caption": "Figure 1. Determination of the geometric location of the points of the tooth profile, taking into account the error.", + "texts": [ + " The value of the profile error and its shape have a significant effect on the nature of the gear train and determine the dynamic component of the load in the gearing. GOST 1643-81 gives the following definition of the profile error: \"The normal distance between the two nearest to each other nominal face profiles of the tooth, between which the real face effective profile of the gear tooth is located\" [3]. The International Conference on Aviation Motors (ICAM 2020) Journal of Physics: Conference Series 1891 (2021) 012051 IOP Publishing doi:10.1088/1742-6596/1891/1/012051 Figure 1 shows the scheme for determining the value of the profile error. As can be seen from the figure, there are several zones on tooth profile. The first zone is located in the section of the true involute of the AB profile. In this zone, each point M of the profile corresponds with the point of the ideal involute profile i with an error. The coordinates of the point i are known and can be determined by the formulas in [4]. The value of the error fi at point M can be determined from the angle of \u03bdi and measured along the normal from point i to point M. The second zone is located on the VA section. The limiting point Vr, located on the tip diameter, will be at a distance fv from the involute at point V, defined by the radius OV. Therefore, when measuring the error in this section, the line of the nominal involute profile must be extended to point V. Diameter Dv is currently missing on the drawings. The position of point F must also be resolved to determine the beginning of the modification line. Nevertheless, the approach shown in Figure 1 allows us to create a mathematical model, in which for each point of tooth flank, the function of the error value from the angle vi can be determined. In some cases, the radius of curvature of the involute or the radius Ri from the center of the gear O to point i can be used as an argument to the function [5]. By setting a value of intentionally introduced profile error over the entire area of the tooth flank (including zero values), we can obtain the desired modification shape. The resulting geometric model can be used in the manufacture of a gear" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001586_iembs.2008.4649475-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001586_iembs.2008.4649475-Figure3-1.png", + "caption": "Fig. 3. Computer depiction of the battery removal tool.", + "texts": [ + " Due to the torque created, minimal force would be required, and the handle at the end would allow for good control. Additionally, the tool would easily fit into various spaces and would be easier to use than the other designs, which entail wrapping a tool around a battery. The other two designs would also require more force in an upward direction, which might be difficult to create (since the device would have to be held down during the process) or unbalance a person when the battery was released. Fig. 3 contains a computer-generated design of the final battery removal tool. Note the slightly curved tip, which is subtle enough to hug any size battery from AAA to D. Also, the texture and shape of handle fits the curves of the human palm well to aid in control and movement. Other members of the Parkinson's support group mentioned that it can be difficult to make it to the front door of their home in time to greet guests. Often they expend the effort to get to their door, only to arrive after their guest has already left" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002838_012021-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002838_012021-Figure2-1.png", + "caption": "Figure 2.(Step1) Figure 3.(Step2)", + "texts": [ + " List of materials used for the analysis are: 1. Al 6061 T6 2. Structural Steel 3. Kevlar 29 4. S-Glass 5. E- Glass 6. Basalt Fiber 7. Carbon Fiber 8. HSLA Steel Properties of selected materials for the analysis of rim as shown in Table 1. 1.2. Modeling of Rim FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 Rim was designed by using 3D modeling software solid works, 2019 version [14]- [15]. Various steps used to create the 3D model of rim modelling. As figure 2. indicates the flange part of the rim which is a design by using revolve command whereas figure 3. illustrate the flange part with the hub designed by using extrude command and Figure 4. gives the complete design of the rim and its cuts by using extrude cut command. The paper aims to analyze different properties in different materials selected for the rim. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 2. Finite Element Analysis of Rim: 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001413_j.jsv.2008.02.056-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001413_j.jsv.2008.02.056-Figure2-1.png", + "caption": "Fig. 2. Plans and sections showing forces on typical ring element. A2 and B2 are at the centroid of ring element cross section.", + "texts": [ + " 1 Di0 deflection of base structure in coordinate i due to a combined external loading and temperature change, elements of column matrix of displacements with F \u00bc 0 E Young\u2019s modulus F column matrix of redundant forces F 01 unit virtual longitudinal force per unit length applied at ring coordinate 1 h height of ring element measured parallel to ring element axis of symmetry H1 radial force per unit length applied at ring coordinate 1 H2 radial force per unit length applied at ring coordinate 2 H 01 unit virtual radial force per unit length applied at ring coordinate 1 H 02 unit virtual radial force per unit length applied at ring coordinate 2 Hp radial load per unit length at R3 due to external pressure HT P total circumferential radial loading from HP Iz 0 moment of inertia of ring cross section about z0 axis \u2018 meridional length of ring element, Fig. 1 L longitudinal load per unit length applied at ring coordinate 1 Mz 0, My 0 induced internal moments about z0 and y0 axes, respectively, due to applied loadings Mj total internally redundant meridional moment in coordinate j of the base structure my0 virtual moment applied about y0 axis at points A and B, Fig. 2 p uniform pressure on ring element surface Pi total internally redundant radial force in coordinate i of the base structure Q temperature rise of ring R radius of centroid of ring element cross section R1 radius of ring element cross section at ring coordinate 1 R2 radius of ring element cross section at ring coordinate 2 R3 radius of center of pressure for uniform pressure on outer face of ring element s variable distance measured along ring element meridian from coordinate 1 S direct force due to radially applied loads t thickness of ring element of rectangular cross section t1 thickness of ring element measured nor- mal to meridional centerline at coordinate 1 t2 thickness of ring element measured normal to meridional centerline at coordinate 2 T torque per unit length about centroid of ring cross section Tp torque per unit length about centroid of ring cross section due to external pressure TL torque per unit length about centroid of ring cross section due to L T01 torque per unit length about centroid of ring cross section due to combined pressure and longitudinal loading with pressure loading resisted at R1 T02 torque per unit length about centroid of ring cross section due to combined pressure and longitudinal loading with pressure loading resisted at R2 T3 meridional moment per unit length applied at ring coordinate 3 T4 meridional moment per unit length applied at ring coordinate 4 T 03 unit virtual meridional moment per unit length applied at ring coordinate 3 T 04 unit virtual meridional moment per unit length applied at ring coordinate 4 Vp1 longitudinal load per unit length resisted at ring coordinate 1 due to pressure loading Vp2 longitudinal load per unit length resisted at ring coordinate 2 due to pressure loading Vp longitudinal load as force per unit length at R3 due to external pressure T", + " The circumferential stresses sx will be expressed in terms of the y0\u2013z0 system of coordinates. However, for purposes of integration over the ring element cross section to determine the flexibility influence coefficients, the y\u2013z system of coordinate will be used. It is thus seen that the circumferential stresses will be as determined for a coordinate system of y0\u2013z0 which has no axis of symmetry. To aid in the development of the expressions for the forces and stresses on any ring element cross section shown in Fig. 1, plans and sections showing these forces are given in Fig. 2. Shown in Fig. 2 are a plan view of the ring element, a cross section showing the y0\u2013z0 coordinate system, and a plan view of half of a ring element showing real and virtual forces applied thereto. From the half-plan view, it is seen that, for a uniform torque T around the element, equilibrium requires that Mz0 \u00bc Z p=2 0 TR cosfdf \u00bc TR: (1) To determine My0, cut the structure on the diametral line A2\u2013B2 and apply a unit virtual moment my0 \u00bc 1 as shown in Fig. 2 before the application of the torque T. Let y11 \u00bc the rotation at points A2 and B2 about the y0 axis due to my0 \u00bc unity at A2 and B2. And let y10 \u00bc the rotation at points A2 and B2 about the y0 axis with My0 \u00bc 0. The stresses sx in the ring element due to Mz0 are by Eq. (132) of Ref. [32] sx\u00f0Mz0 \u00de \u00bc Mz0 \u00f0Iy0y 0 Iz0y0z 0\u00de Iz0Iy0 I2z0y 0 . (2) ARTICLE IN PRESS T.A. Smith / Journal of Sound and Vibration 318 (2008) 428\u2013460434 The stresses sx due to My0 are by Eq. (133) of Ref. [32] sx\u00f0My0 \u00de \u00bc My0 \u00f0Iz0y0y 0 Iz0z 0\u00de Iz0y0 I2z0y0 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002738_s0263574721000539-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002738_s0263574721000539-Figure7-1.png", + "caption": "Figure 7. Dynamics model of surface elastic body.", + "texts": [ + " In the robot model, a DC motor is assumed, and its driving torque, TM , is given as follows: TM = (JM + \u03bdM)\u03b8\u0308M + Tr , (5) where JM is the entire moment of inertia of the motor rotor with the eccentric weight, \u03bdM is the motor\u2019s damping coefficient, and Tr is the external torque resistance. Here, \u03bdM can be assumed to be ideally 0 when a gear-head is not embedded in the motor. Tr is also approximated to be zero, owing to g<< 1. Accordingly, the motor torque can be given as follows: TM \u2248 JM \u03b8\u0308M , (6) where the reaction torque, \u2212TM , acts on the robot. To formulate the reaction forces, Ni and Ffi, the deformation change of elastic body must first be introduced. In this paper, we represent the elastic body as a two-link system with rotational viscoelastic joints, as illustrated in Fig. 7. The joint angles, \u03c61i and \u03c62i of the elastic body, can be expressed as follows: \u03c61i = \u03c6i + cos\u22121 ( hi h0 ) , \u03c62i = 2 cos\u22121 ( hi h0 ) , (7) where h0 is the natural length of the elastic body, and \u03c6i and hi are, respectively, defined as the equivalent joint angle and length of the elastic body, as shown in Fig. 6. hi is given as follows: hi = \u221a( x + rix cos \u03b8 \u2212 riy sin \u03b8 \u2212 pix )2 + ( y + riy sin \u03b8 + riy cos \u03b8 )2, (8) where ri(rix, riy) is the vector from the robot CoM to the base position of elastic body" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002191_j.matpr.2020.12.112-Figure9-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002191_j.matpr.2020.12.112-Figure9-1.png", + "caption": "Fig. 9. Total deformation-1.", + "texts": [ + " This point bargains about the compelled part appearing of the posts and imitating them by contributing specific physical representations and limit situations to reenact the fascinating situation plans ended likely. The segment stayed demonstrated utilizing ANSYS 14.0 software Plan Modeler programming as a strong perfect and ANSYS 14.0 worktable was utilized aimed at the assess- ment in the static examination. The key work method of a Limited Component Examination system is tended to like in Fig. 1(A) Fig. 2 Fig. 3 Fig. 4 Fig. 5 Fig. 6 Fig. 7 Fig. 8 Fig. 9 Fig. 10. The stainless harden fabric property Bulk (q): 7950 kg/m3 Young\u2019s modulus (E):206000Mpa Poisson\u2019s ratio: 0.3 The CAD typical of the bar is fit hooked on a limited quantity of components utilizing ANSYS 14.0 software inherent lattice calculation. The pillar contact locale is fit and interlinked to empower estimation of the power collaboration between them limited component investigation or FEA representing to a genuine task as a \u2018\u2018work\u201d a progression of little, consistently formed tetrahedron associated components, as appeared in the above fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002969_j.matpr.2021.05.300-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002969_j.matpr.2021.05.300-Figure8-1.png", + "caption": "Fig. 8. Al nitride clutch plate heat flux results.", + "texts": [ + " Developing 3D model of MPCs utilizing Solid edge programming. Exporting the model of MPCs from strong edge to ANSYS 19.2 work bench for FEA. Static and dynamic examination of MPCs utilizing ANSYS work seat at various working condition. Determining the wear, most extreme disfigurement and equal worries for every material utilizing FEA Showing the connection of disfigurement and obstruction property of every material utilizing figures and tables (Fig. 1. Fig. 2. Fig. 3. Fig. 4. Fig. 5. Fig. 6. Fig. 7. Fig. 8. Fig. 9. Fig. 10. Fig. 11. Fig. 12. Table 1). Comparing and conversation of results for choice of better covering materials. Fig. 5. Cast iron clutch plate heat flux results. Fig. 6. Cast iron clutch plate heat flux results. Grip disappointment and harm because of extreme frictional warmth and warmth changes to the grasp counter mate circle frequently happens to a car grips. This circumstance adds to warm weakness to the segment which causes the grasp counter mate plate to break and twist. This later will make issues, for example, grip slip, grasp drag or disappointment of hold to separate appropriately and grip shaking just as shortening the lifecycle of the part" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001917_6.2007-7894-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001917_6.2007-7894-Figure1-1.png", + "caption": "Fig. 1 CAD rendition of the high altitude airship.", + "texts": [ + " Experimental techniques have been developed to evaluate both qualitatively (optical) and quantitatively (load cell) the performance of such devices at a range of Reynolds number. Then the CFD method and its validation are described; this RANS CFD is used to clarify some of the observations from the experiments. The results of the experiment are then presented, and finally, the most significant findings are summarized. II. Experimental Approach The hull of the airship is an ellipsoid that is equipped with two horizontal stabilizers and an aft-mounted vertical rudder, Fig. 1. The ellipsoid has an aspect ratio of 3.0. A gondola, with a delta wing configuration, can be attached below the ellipsoid through an NACA airfoil shape strut. The scaled model, which is one to two orders of magnitude smaller than the real airship depending on the application, is tested in the ETHZ-LSM towing tank. Figure 2 shows the Water Towing Tank (WTT). The towing tank has dimensions of 40m in length, 1m in width, and American Institute of Aeronautics and Astronautics 3 1m in depth. In a WTT the model is dragged through stationary water [6]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002687_j.seta.2021.101240-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002687_j.seta.2021.101240-Figure7-1.png", + "caption": "Fig. 7. Modified Pole-PMDC Motor using Infolytica MotorSolve Software.", + "texts": [ + " Flux lines of conventional PMDC motor provide high armature current at no load torque, flux lines are clearly indicating their trend in comparison with R. Ullah Khan et al. Sustainable Energy Technologies and Assessments 46 (2021) 101240 the conventional PMDC Motor is shown in Fig. 6. The modified PMDC motor has small air gap between the stator and rotor position. Moreover, this modified PMDC motor net magneto force will be reduced. The modified pole PMDC motor is designed with Infolytica MotorSolve Software is illustrated in Fig. 7. Flux lines Trends of Modified Pole PMDC Motor has been shown with the help of \u2018Infolytica MotorSolve Software. After cutting the pole magnets into pieces, it has been observed that, the concentration of flux lines are more towards the gap in magnets of main poles i.e. South Poles and North Poles of the Magnets assemblies. Flux lines of modified PMDC motor is shown in Fig. 8. An air gap of the flux density distribution is the first-rate concurrence, concurrently armature speed and RMS current is full of very high" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000185_11802372_33-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000185_11802372_33-Figure7-1.png", + "caption": "Fig. 7. Static analysis of climbing", + "texts": [ + " Then the output of RL is the cooperative strategy for the multirobot to execute the specific mission. After the mission is fulfilled successfully, all the robots are waiting for a new mission transferred from father robot. Fig.6 shows that when a single child robot4 cannot climb up a slope, other three robots connected with it end to end and help it to overcome this obstacle. In order to help robo4 climb up the slope successfully, the distance between each two child robot should be adjusted during the process of climbing, such as d32 and d43 in Fig.7. Therefore, the inverse kinematics calculation is inescapable. The reference frame of a single robot is shown in Fig.8. The D-H parameters are shown in table.1.Where L1 to L6 are 60cm, 10cm, 60cm, 40cm, 20cm and 5cm respectively. By static analyzing, we found the structure that robot2 and robot1 stay under the bottom of slope can provide a maximum climbing force for robot4. We also obtained an expression, sin 1 cos \u03b1\u03bc \u03b1 \u2265 + . (1) Where \u03bc is a friction coefficient between child robot with the slope, \u03b1 is the maximum gradient of the slope the multirobot team can climb, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003132_978-981-16-0119-4_20-Figure78.3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003132_978-981-16-0119-4_20-Figure78.3-1.png", + "caption": "Fig. 78.3 Conceptual design of propeller with saw tooth cut at both TEs", + "texts": [], + "surrounding_texts": [ + "enhancement technique and its positions on the propeller for this current work [9]. While coming to the methodology section, the two engineering approaches have beenused to test the aerodynamic behavior on the relevant components. The methodologies are experimental test andComputational FluidDynamics (CFD)based simulation, in which most of the aerodynamic analyses were executed with the help of CFD simulations. Because of these huge implementations, the pre-processor steps are available to everyone [11]. Especially, the boundary conditions such as type of solver, type of turbulence model, quantity and quality of turbulence model are easily available data to the new researchers. The high amounts of analyses were used velocity inlet to their problems. Thus, with these inputs, the CFD simulations are executed in these comprehensive aerodynamic analyses on various propellers [13]. b. Solution Techniques \u2013 CFD Analysis In this work, analyses the comparative aerodynamic performance on a UAV\u2019s propeller by using CFD tool, i.e., Ansys fluent. The fundamental aim of this work is to select the suitable lift enhancement technique for UAV\u2019s propeller. In this regard, six different designs are modeled, in which five conceptual designs are comprised of a propeller with edge modifications, and the other one is the conceptual design of base propeller. The techniques implemented in the 5-inch diameter propeller are curvy cut saw tooth cut, aero cut, etc., and in general, the profile modifications in the UAV\u2019s propeller is executed for noise reduction. From the literature survey, it was clearly understood that the noise induced due to the abnormal environment is reduced [1, 2]. Nowadays, UAV industry needs a quite propeller so the design modifications-based propellers are suggested a lot for the construction of quite UAV. But the problem along with these types of profile modified propellers may have a chance to generate low-aerodynamic forces. Hence, the conduction of an integrated study is very important in the updated propellers to increase the implementation of the UAVs in real-time applications. The steady- and pressure-based turbulent flow is used as fundamental behavior to the working fluid for all these analyses. The aerodynamic parameters such as lift, drag, CL, CD are used as selection parameters for this comparative analysis [15, 16and17]. c. Conceptual Design of various propellers UAV\u2019s propeller and its designs are the key role of this aerodynamic performance investigation. From the previous work, it was understood that propeller with saw tooth cut has been provided the low turbulence noise than base propellers. Thus, in this work, the aerodynamic performance of low acoustic profiled propellers is computed, in which the conceptual design of all the propellers are used from the literature survey [18, 19,and20]. Figures 78.1, 78.2, 78.3, 78.4, and 78.5 are revealed the conceptual design of propellers with saw tooth cuts, in which the locations of the saw tooth cuts are formed at the various edges of the propellers such as both leading edges, both trailing edges, one leading edge cum one trailing edge, only one trailing edge. Apart from these saw tooth cuts, the one more relevant cut is located at the leading edges of the propeller, which is v-cut. 78 Comparative Aerodynamic Performance Analysis on Modified \u2026 973" + ] + }, + { + "image_filename": "designv11_83_0002381_j.matpr.2021.01.637-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002381_j.matpr.2021.01.637-Figure8-1.png", + "caption": "Fig. 8. Meshing.", + "texts": [ + " SOLIDWORKS allowed only the application of forces on a surface. So small areas of a few tenths of a square millimeter were created to apply these forces. They were created as small so as not to deviate too much from the real conditions (Fig. 5). Five contact zones on the tooth profile where the forces will be applied have been selected. With regard to the imposed displacements, this involves embedding on the lower surface of the tooth as shown in Fig. 7. The mesh under SOLIDWORKS was generated using given meshing parameters and, Fig. 8 shows an example of a mesh used. Once the mesh is executed the calculation of stresses and deformations is launched. For the study of the influence of the torque, all the other parameters were fixed and the torque was varied (5 Nm, 10 Nm) referring to the torques chosen for the experimental tests carried out from 2.5 Nm to 15 Nm with a pitch of 2.5, [1,13]; all the values used for the experiments were used for the simulation, which would allow comparison to be made, but two values were chosen for the article. Experience has shown that when the wear rate reaches 12%, the tooth deteriorates rapidly, therefore wear rates of 0%, 4% and 8% were chosen to remain within allowable limits. These fixed parameters are: number of teeth on the pinion, Z1 = 17; Number of teeth on the wheel, Z2 = 36, pressure angle: 20 ; Module : 5 The Figs. 7, 8, and 9 illustrate the modelling process including boundary conditions (Fig. 7), mesh profile (Fig. 8) and the Von Mises stress in Fig. 9. Key results for contact stress are presented in Fig. 10, whereas, in Fig. 11, the bending stresses at the root of the tooth are given. It was found that, the contact stresses for the different wear rates vary with the torque applied so that as the torque increases, the stresses also increase; a logical observation since the increase in torque leads to the increase of the actual normal load taken up by a pair of teeth in contact, therefore an increase in contact stresses" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000828_3.43615-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000828_3.43615-Figure1-1.png", + "caption": "Fig. 1 Steady level flight.", + "texts": [ + " If the neutral point is aft of the e.g., the configuration is statically stable. One can take advantage of the neutral point to formulate certain problems in a manner resulting in a simplification of the moment equation. This is illustrated for the followingthree cases: 1) the determination of level-flight trim conditions, 2) the determination of maneuverability, and 3) the stability of the short period mode. An airplane or missile in steady level flight is subjected to a force system shown acting at the e.g. in Fig. 1. The angle of attack a and the control surface deflection d are considered to be small. The force system is dependent on the angle of attack and the control system deflection. The system is in equilibrium. The force system of Fig. 1 is repeated in Fig. 2, with a linear dependence of the force and moment on the angle of attack and control deflection being assumed. The forces D ow nl oa de d by U N IV E R SI T Y O F M IC H IG A N o n Fe br ua ry 9 , 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /3 .4 36 15 and moment have been divided by ^pV^S and ^pVozSl, respectively, to place them in coefficient form, in which p = density of the medium VQ = freestream velocity S = a reference area I = a reference length Since the system is in equilibrium, the angle of attack and control surface deflection must be such that the lift coefficient is the trim lift coefficient, the moment coefficient is zero, and the thrust coefficient is equal to the drag coefficient at a given velocity and density" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003042_s11668-021-01191-x-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003042_s11668-021-01191-x-Figure3-1.png", + "caption": "Fig. 3 Meshed model of EGR cooler housing", + "texts": [ + " The next key components are the EGR pipes; the flexible bellows on these pipes are useful for thermal expansion during hot conditions. In this case the bellow crack failure is observed due to vibration. Therefore, design improvements due to proper design of bellow geometry and bracket support for these pipes facilitated prevention of EGR pipe failure. Also due to vibrations, even after stresses and deformation were within the permissible values, there has been leaking in the mounting bosses has been observed for the existing model (shown in Fig. 3). From the literature review it is found that a lot of work has been done in vibration analysis for the different parts of an automobile. However, frequency response analysis of EGR to find the displacements and stresses is still not found any suitable literature. This paper also describes the methodology of conducting proper failure investigation to identify the root cause for vibration-related failures of EGR system during new product development. In view of this, the main objective of this work is to carry out and compare the various geometrical models of the EGR generated using the finite element simulation method", + " Modification 4 After analyzing the modified model 3, the stress and deflection were optimum, but the bosses and ribs were adjusted to provide for material optimization. So, following changes are made in the Modification 3, (a) bosses thickness reduces back to 15 mm. (b) Increase the rib thickness Fig. 2 Modified model from 3.8 to 6 mm. (c) A third rib was added between the mounting bosses with height 5 mm. The 3D modeling of EGR cooler system is meshed using the Hypermesh software using the solid 3D, four node and tetrahedron elements with three degrees of freedom. The tetrahedral element meshing of the EGR cooler housing is shown in Fig. 3. Tetrahedral elements are used here because they fit arbitrary shaped geometries very well with their simple computations. In comparison, the hexahedral meshes are more accurate with the number of elements, since one hexahedral equal to six tetrahedral elements. However, the tetrahedral elements are best to model complex geometry domain with little distortion of mesh. While meshing the model the conditions given are linear elastic, isotropic and temperature independent. The material properties inputted are listed in Table 1 for aluminum for housing, cast iron for bracket and steel for the bolt" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000024_2005-01-3371-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000024_2005-01-3371-Figure4-1.png", + "caption": "Fig. 4 Normalized error in stress distribution over the gear; (a) radial stress (b) tangential", + "texts": [ + " 2(a) and 3(a), it is shown that the prediction follows the true solution closely. The rms error over the entire grid is about 5% . Fig. 3(b) shows the predicted normal tangential stress distribution. Again, comparison of Fig. 3(b) with Fig. 2(b) shows that the predicted and true tangential stresses are in close agreement. Extending the learning models thus constructed to test damage data, with one gear tooth made less rigid than the rest of the teeth, predictions of radial stress are made by the thus learned model. These predictions are shown in Fig. 4(a) as the difference between the true normal radial stress and the just predicted damage radial stress distributions. Inspection of Fig. 4(a) clearly shows a discrepancy in the radial stress distribution in a given region, and this local region turns out to be correspondent to be the damaged tooth. Same observation is made on inspection of Fig. 4(b) with respect to the tangential stress distribution. Figs. 5(a-b) show the line plots of the error in radial and tangential stress distributions, respectively. The abscissa denotes the grid point index; the first 420 grid points represent the first tooth; the last 420 grid points represent the nineteenth tooth. The damage is clearly predicted around the first and the 19th teeth. Finally, Fig. 6 shows the shear stress error distribution identifying the damaged tooth. stress This study thus establishes that for structural health monitoring, physics-based modeling and simulation is required to provide nominal solutions, both normal and damage, as reference solutions, which can then be used to train machine learning algorithms to subsequently test the incoming sensor data over a distributed network of sensors to detect damage" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001641_ijvnv.2007.015175-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001641_ijvnv.2007.015175-Figure1-1.png", + "caption": "Figure 1 The pumping cycle of a PAS pump (for colours see online version)", + "texts": [ + " He worked as Senior Lecturer in The City University, London and worked in British Aerospace, UK. He is currently a professor in automotive and aerospace engineering, at RMIT University. Power Assisted Steering (PAS), more commonly referred to as power steering, functions to make steering easier. The PAS system is a hydraulic system in which a pressure is applied to support the motion favoured by the torque of turning the steering wheel. The PAS pump is like any conventional vane pump containing vanes which, when rotating, cause the flow of fluid in a certain direction. As shown in Figure 1, the fluid flows in three stages where the impeller turns clockwise. In the first stage the lower pressure generated by the rotation of the vanes away from the inlet, causes a low pressure at the inlet which induces the liquid into the pump. Continuous rotation forces the flow through as conveyed in the second stage. In this stage, the fluid is under higher pressure as there is compression from the vanes into a closed area. The third stage results in the exit of the fluid through the outlet. The PAS pump is driven by the crankshaft through the associated pulley-belt system; the first order of the PAS pump shaft rotation is determined by the ratio (\u03d5) of crankshaft pulley diameter to the PAS pump pulley diameter, which means if the crankshaft rotates once, the PAS pump will rotate \u03d5 revolutions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002530_cerma.2007.4367740-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002530_cerma.2007.4367740-Figure1-1.png", + "caption": "Figure 1. A tractor and a trailer linked for an off-centered joint", + "texts": [ + " All these maneuvers are made in applying the same control law. The configuration of an articulated mobile robot in a plan can be defined by the following state vector: q = [x, y, \u03b8, \u03c61, \u03c62, . . . , \u03c6n]T where (x, y) is the position of a fixed point on the robot, \u03b8 the orientation of tractor, and \u03c61, \u03c62, . . . , \u03c6n the orientation of the n trailers joined to the tractor, all these orientations are taken with respect to the X axis. The kinematics model of an articulated mobile robot like the robot shown in Figure 1 (with one module only) can be defined by: dq dt = f1(q)v + f2(q)\u03c9 with: f1(q) = [ cos \u03b8 sin \u03b8 0 \u2212 1 BEFORE1 sin \u03c6 ]T f2(q) = [ 0 0 1 ( \u22121 \u2212 BACK0 BEFORE1 cos \u03c6 ) ]T where v is the tractor\u2019s lineal velocity and \u03c9 is the tractor\u2019s angular velocity, BACK0 represents the measurement of the length of the back joint of the robot and BEFORE1 corresponds to the measurement of the length of the front union between the first tow and the tractor. As a nonholonomic system, this system has non integrable velocities restrictions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000469_50009-6-Figure8.18-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000469_50009-6-Figure8.18-1.png", + "caption": "Figure 8.18. Lateral views of lower jaw of representative cetaceans. (a) Mysticete, minke whale, Balaenoptera acutorostrata. (From Dem\u00e9r\u00e9, 1986.) (b) Odontocete, spotted dolphin, Stenella sp.", + "texts": [ + " Below the spermaceti organ is a region of connective tissue alternating with spaces filled with spermaceti oil. This region was called the junk by whalers because it contains an oil of poorer quality (Clarke, 1978; see Figure 8.17). The spermaceti \u201cjunk\u201d is probably homologous with the melon of other odontocetes. The function of these structures in sound production is discussed in more detail in Chapter 11. The lower jaws, or mandibles, of odontocetes appear straight when viewed dorsally. The posterior non-tooth bearing part of the jaw has thin walls that form the fat filled pan bone (Figure 8.18). Norris (1964, 1968, 1969) proposed that this region is the primary site of sound reception in odontocetes (discussed in Chapter 11). In mysticetes the mandible curves laterally and there is no pan bone. The mandibular symphysis, a fibrocartilage articulation, connects the tapered distal ends of the paired dentary bones. It is analogous in structure to the intervertebral joints; its center, filled with a gelatinous substance, is surrounded by a dense fibrocartilaginous capsule. The coronoid process (see Figure 8.18) for attachment of the temporalis muscles is reduced in most odontocetes (an exception is the susu, which retains a distinct process). Among mysticetes, the coronoid process is of moderate size in balaenopterids and developed as a slightly upraised area in the gray whale, pygmy right whale, and the balaenids (Barnes and McLeod, 1984). The function of the coronoid process in mysticete feeding is discussed in Chapter 12. The hyoid bones are well developed in all cetaceans. In odontocetes the hyoids are divisible into a basal portion (basihyal, paired thyrohyals) and a suspensory portion (paired ceratohyals, epihyals, stylohyals, and tympanohyals; Reidenberg and Laitman, 1994)", + " Another conclusion from this study was that the dolphin vertebral column has the capacity to store elastic energy and to dampen oscillations as well as to control the pattern of body deformation during swimming. The sternum of odontocetes differs from that of mysticetes in several features. In odontocetes, 5\u20137 pairs of ribs usually are attached to the sternum (sternal ribs), which consists of a long, flat bone, usually segmented and widened anteriorly, with flared cup-like recesses at the site of attachments for ribs (see Figure 8.18). In mysticetes, there is only one pair of ribs attached to the sternum and the sternum is shorter and broader than in odontocetes (Yablokov et al., 1972). Also unlike most other mammals, odontocetes have bony rather than cartilaginous sternal ribs (Rommel and Reynolds, 2001). Cetaceans are unique in having single-headed ribs (but see later) that attach to their respective vertebrae only to the transverse processes instead of having two vertebral attachments points, one to the vertebral body and the other to the transverse processes, as in most mammals" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003523_s12046-021-01692-3-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003523_s12046-021-01692-3-Figure3-1.png", + "caption": "Figure 3. Schematic artwork of the tensile samples: ASTM. standard E8/E8M-09 (All dimensions are represented in mm).", + "texts": [ + " The chosen factors and their welding ranges are given in table 3. The interpulse gas tungsten constricted arc welding arrangement which involves an IPTIG machine, argon gas cylinder, and fixture is shown in figure 1. The prepared welded joints are illustrated in figure 2. For the improvements of weld strength, the heat treatment cycle is followed as 8008C-8Hrs-gas fan quenching. Tensile test specimens were extracted from each welded coupon. The ASTM E8/E8M-09 standard was referred and its dimensions are shown in figure 3. The schematic of extraction of numbers of tensile specimens and extracted pieces are shown in figure 4. The basic scale i.e., 1:1 was chosen on basis of which the welded joints and tensile specimens were collected. After the extraction, the samples were undergone for tensile testing on a computerized model number TUE-600(C) universal testing machine. The factors which were used to prepare the 17 numbers of the welded joints along with the measured tensile strength of welded joints are given in table 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003039_icit46573.2021.9453629-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003039_icit46573.2021.9453629-Figure5-1.png", + "caption": "Fig. 5. Assistive force amplification mechanism", + "texts": [ + " This artificial muscle contracts in the axial direction and expands in the radial direction by applying air pressure. It has properties similar to those of human muscles because it exerts an axial contraction force. The straight-fiber-type artificial muscle is highly constrained in the axial direction, meaning that it has a higher output density and can generate a larger shrinkage rate and contraction force than a McKibben artificial muscle of the same shape [16]-[17]. 2) Amplification mechanism: Fig. 5 shows the amplification mechanism in which a balloon actuator is mounted between two aluminum plates connected using hinges. The actuator is combined with a highly sealable, rectangular, and nonstretchable fabric. It expands when air pressure is applied. The expansion force is transmitted to the human body via the aluminum plate and increases the assistive force to the upper body. The amplification mechanism can help close the plate during evacuation via a tension spring connected to the plate. This mechanism is closed when no assistance is performed, meaning that the device does not hinder the movement of the wearer" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001774_amc.2008.4516140-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001774_amc.2008.4516140-Figure1-1.png", + "caption": "Fig. 1. Powerball\u00ae components", + "texts": [ + " The control strategy is oriented towards underactuated system, active and passive robots degrees of freedom. I. INTRODUCTION Powerball\u00ae, Dynabee\u00ae, and Gyrotwister\u00ae are gyroscopic devices popular in the 90-ties. They are dedicated to the wrist exercising and are patent pended [1]. Rolling connection of the rotor and the device housing enables the spin-up of the rotor by the appropriate wrist rotations. This movement is accompanied with the torque vector reaction proportional to the square of the rotor spin. The rotor can reach the speeds up to 16000 rpm for the plastic version showed on Fig.1. and an astonishing 20000 rpm for the metal version. Our scenario includes manipulating robot, which is able to spin-up the Powerball\u00ae. This gyroscopic device represents a dynamic load to the robot as is discussed in [2] and represents underactuated and nonholonomic load for the manipulator kinematic chain [3]. Spin-up of the wrist exerciser is based on rolling which depends on friction between the rotor shaft and the housing. Friction is neglecting or it became significant. It depends on rotor state and on the wrist rotation reaction normal force on the shaft housing connection" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001366_s1068798x07110020-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001366_s1068798x07110020-Figure1-1.png", + "caption": "Fig. 1. Total torque of the centrifugal inertial forces due to the crankshaft counterweights and its components, in the general case, for the initial counterweight positions chosen in V8-motor design and the adjusted positions intended to compensate the total unbalanced mass and eliminate the unbalanced torque.", + "texts": [ + " The torques of the centrifugal inertial forces due to the mass and geometric parameters of the counterweights are (3) (4) where m w i , m w( i + 1) , m w j , and m w( j + 1) are the masses of the counterweights on the extensions of crankshaft side pieces i , ( i + 1), j , and ( j + 1); \u03c1 i , \u03c1 ( i + 1) , \u03c1 j , and \u03c1 ( j + 1) are the distances from the axes of rotation to the center of gravity of masses m w i , m w( i + 1) , m w j , and m w( j + 1) ; l w i is the distance between the counterweights on the continuations of crankshaft side pieces i and (9 \u2013 i ); l w( i + 1) is the distance between the counterweights on the continuations of side pieces ( i + 1) and (8 \u2013 i ); l w j is the distance between the counterweights on the continuations of crankshaft side pieces j and (9 \u2013 j ); l w( j + 1) is the distance between the counterweights on the continuations of side pieces ( j + 1) and (8 \u2013 j ). The formation and location of the torques is illus- trated in Fig. 1. Torques M 0c i and M 0c j are the geometric sums of these torques (5) On the basis of the cosine theorem, the torques are defined as (Figs. 1a and 1c) (6) (7) where \u03b3 0 i is the initial angle between the torque vectors M 0 i and M 0( i + 1) ; \u03b3 0 j is the initial angle between the torque vectors M 0 j and M 0( j + 1) . M0i mwi\u03c1ilwi\u03c9 2;= M0 i 1+( ) mw i 1+( )\u03c1 i 1+( )lw i 1+( )\u03c9 2;= M0 j mw j\u03c1 jlw j\u03c9 2;= M0 j 1+( ) mw j 1+( )\u03c1 j 1+( )lw j 1+( )\u03c9 2,= M0ci M0i M0 i 1+( ); M0c j+ M0 j M0 j 1+( ).+= = M0ci M0i 2 M0 i 1+( ) 2 2M0iM0 i 1+( ) \u03b3 0icos\u2013+( )0.5 ;= M0c j M0 j 2 M0 j 1+( ) 2 2M0 jM0 j 1+( ) \u03b3 0 jcos\u2013+( )0.5 ,= Compensating the Total Unbalanced Mass of Crankgear Parts of a V-8 Motor by Changing the Configuration of the Crankshaft Counterweights A. D. Nazarov DOI: 10.3103/S1068798X07110020 RUSSIAN ENGINEERING RESEARCH Vol. 27 No. 11 2007 COMPENSATING THE TOTAL UNBALANCED MASS 739 740 RUSSIAN ENGINEERING RESEARCH Vol. 27 No. 11 2007 NAZAROV These angles are (Fig. 1a) (8) where \u03b10i is the initial angle between the planes of the counterweights on the continuations of crankshaft side pieces i and (9 \u2013 i) and crankshafts 1 and 4; \u03b10(i + 1) is the initial angle between the planes of the counterweights on the continuation of side pieces (i + 1) and (8 \u2013 i) and crankshafts 1 and 4; \u03b20j is the initial angle between the planes of the counterweights on the continuations of side pieces j and (9 \u2013 j) and crankshafts 2 and 3; \u03b20( j + 1) is the initial angle between the planes of the counterweights on the continuation of side pieces ( j + 1) and (8 \u2013 j) and crankshafts 2 and 3", + "5;cos \u03d50itan ci \u03b10isin c i 1+( ) \u03b10 i 1+( )sin+( )= \u00d7 ci \u03b10icos c i 1+( ) \u03b10 i 1+( )cos+( ) 1\u2013 ; \u03c80 jtan c j \u03b20isin c j 1+( ) \u03b20 j 1+( )sin+( )= \u00d7 c j \u03b20 jcos c j 1+( ) \u03b20 j 1+( )cos+( ) 1\u2013 , where ci and cj are constant structural coefficients (parameters) (19) Here k(i \u2013 1) and k( j \u2013 1) are constant structural coefficients (parameters); lw1 is the distance between the counterweights on the continuations of crankshaft side pieces 1 and 8. The coefficients may be calculated from the formulas (20) where mw1 is the initial mass of the counterweight on the continuation of side piece 1; \u03c11 is the initial distance from the axis of rotation to the center of gravity of mass mw1. The resultant Mw0 of the torques of the centrifugal inertial forces due to the counterweights and its angle of action \u03b1w0 are calculated as follows (21) (22) where \u03b30 is the initial angle between the vector torques M0ci and M0cj (Fig. 1c) (23) Correspondingly, (24) Taking account of Eq. (24), it follows from Eq. (21) that (25) On the basis of Eqs. (15) and (16), we may rewrite Eq. (25) in the form (26) Here (27) For specific V-8 motors, the relevant parameters and dimensions of the crankshaft remain constant through- ci k i 1\u2013( )lwilw1 1\u2013 ; c j k j 1\u2013( )lw jlw1 1\u2013 .= = ki mw i 1+( )\u03c1 i 1+( ) mw1\u03c11( );= k j mw j 1+( )\u03c1 j 1+( ) mw1\u03c11( ),= Mw0 M0ci 2 M0c j 2 2M0ciM0c j \u03b3 0cos\u2013+( )0.5 ;= \u03b1w0tan M0ci \u03d50isin M0c j \u03c80 jcos+( )= \u00d7 M0ci \u03d50icos M0c j \u03c80 jsin+( ) 1\u2013 , \u03b3 0 90 \u03d50i \u03c80 j", + "5,= f 0 mwi\u03c1ilwi;= f 1 ci 2 c i 1+( ) 2 2cic i 1+( ) \u03b10 i 1+( ) \u03b10i\u2013( )cos\u2013+= + c ji c j 2 c j 1+( ) 2 2c jc j 1+( ) \u03b20 j 1+( ) \u03b20 j\u2013( )cos\u2013+[ ]; f 2 = 2c ji ci 2 c i 1+( ) 2 2cic i 1+( ) \u03b10 i 1+( ) \u03b10i\u2013( )cos\u2013+[ ]{ \u00d7 c j 2 c j 1+( ) 2 2c jc j 1+( ) \u03b20 j 1+( ) \u03b20 j\u2013( )cos\u2013+[ ] }0.5 . mwi mw 9 i\u2013( ); \u03c1i \u03c1 9 i\u2013( );= = mwi\u03c1i mw 9 i\u2013( )\u03c1 9 i\u2013( );= mw j mw 9 j\u2013( ); \u03c1 j \u03c1 9 j\u2013( );= = mw j\u03c1 j mw 9 j\u2013( )\u03c1 9 j\u2013( );= \u03b10i \u03b10 9 i\u2013( ); \u03b20 j \u03b20 9 j\u2013( ).= = MR0 Mw0; \u03b1R0\u2013 \u03b1w0 18\u00b026\u2032,= = = MR w\u03c92 m0c \u0394my+( ).= \u0394MR w\u03c92\u0394my.= MR Mw; \u0394MR\u2013 \u0394Mw;\u2013= = \u03b1R \u03b1R0 \u03b1w \u03b1w0,= = = For the angles on the right side in Eq. (38), we may write (39) The torques are calculated as follows (40) It is evident from Eqs. (11), (12), (15), (16), (21), (26), and (28) and from Fig. 1 that, by adjusting (increasing and decreasing) the angular position of the corresponding crankshaft counterweights, Mw0 may be increased by \u0394Mw. As a result, \u0394my is compensated, \u0394MR is eliminated, and complete equilibrium of the V-8 motor is ensured. This is confirmed by the following considerations. When using the given method for compensation of \u0394my and elimination of \u0394MR, the masses of the crankshaft counterweights and the distance from the crankshaft\u2019s center of gravity to the axis of rotation remain constant, while the positional angles change. Therefore (41) where Mi , M(i + 1), Mj , and M( j + 1) are the moments of the centrifugal inertial forces due to the counterweights on the continuation of crankshaft side pieces i, (i + 1), j, and ( j + 1) after adjustment of their positional angles (Fig. 1b). As is evident from Eqs. (11), (12), (15), and (16) and from Fig. 1b, torques M0ci and M0cj increase with increase in \u03b10i and \u03b20j by \u0394\u03b1i and \u0394\u03b2j and decrease in \u03b10( i + 1) and \u03b20( j + 1) by \u0394\u03b1(i + 1) and \u0394\u03b2( j + 1). In that case (42) (43) where Mci and Mcj are the values of M0ci and M0cj after adjusting the positional angles of the corresponding counterweights; \u03b3i is the angle between the torque vectors Mi and M(i + 1) after adjusting the corresponding counterweight positions; \u03b3j is the angle between the torque vectors Mj and M( j + 1), depending on the positional angles of the corresponding counterweights", + " 1a and 1b) (44) where \u03b1i is the angle between the planes of the counterweights on the continuations of crankshaft side pieces i and (9 \u2013 i) and crankshafts 1 and 4 after increasing \u03b10i by \u0394\u03b1i; \u03b1(i + 1) is the angle between the planes of the counterweights on the continuations of crankshaft side pieces (i + 1) and (8 \u2013 i) and crankshafts 1 and 4 after decreasing \u03b10(i + 1) by \u0394\u03b1(i + 1); \u03b2j is the angle between the planes of the counterweights on the continuations of crankshaft side pieces j and (9 \u2013 j) and crankshafts 2 and 3 after increasing \u03b20j by \u0394\u03b2j; \u03b2( j + 1) is the angle between the planes of the counterweights on the contin- \u03b1Rtan \u03b1R0tan \u03b1wtan \u03b1w0tan 1/3.= = = = \u0394MR MR MR0; \u0394Mw\u2013 Mw Mw0.\u2013= = Mi M0i; M i 1+( ) M0 i 1+( );= = M j M0 j; M j 1+( ) M0 j 1+( ),= = Mci Mi 2 M i 1+( ) 2 2MiM i 1+( ) \u03b3 icos\u2013+( )0.5 ;= Mc j M j 2 M j 1+( ) 2 2M jM j 1+( ) \u03b3 jcos\u2013+( )0.5 ,= \u03b3 i 180 \u03b1i \u03b1 i 1+( ); \u03b3 j\u2013+ 180 \u03b2 j\u2013 \u03b2 j 1+( ),+= = 742 RUSSIAN ENGINEERING RESEARCH Vol. 27 No. 11 2007 NAZAROV uations of crankshaft side pieces ( j + 1) and (8 \u2013 j) and crankshafts 2 and 3 after decreasing \u03b20( j + 1) by \u0394\u03b2( j + 1). These angles correspond to the following formulas (Fig. 1b) (45) (46) Analysis of Eqs. (42)\u2013(46) and Figs. 1a and 1b shows that increasing the given torques entails increasing the initial positional angles of the counterweights closer to the crankshaft correction planes and decreasing the corresponding angles of the counterweights farther from these planes. On the basis of the foregoing and Eqs. (9) and (10), we may write (47) (48) The angles of action \u03d5i and \u03c8j of these torques are defined as (49) (50) On the basis of Eqs. (3), (4), (19), (20), (41), (45), and (46), we may express Eqs", + " (57) and (58), the result of increasing \u03b10i by \u0394\u03b1i and \u03b20j by \u0394\u03b2j and decreasing \u03b10(i + 1) by \u0394\u03b1(i + 1) and \u03b20( j + 1) by \u0394\u03b2( j + 1) is to reduce the angles \u03b1f 0 \u2013 (\u0394\u03b1(i + 1) + \u0394\u03b1i) and \u03b2f 0 \u2013 (\u0394\u03b2( j + 1) + \u0394\u03b2j). Consequently, the cosines of these angles are reduced, and the corresponding torques increase. Taking account of Eqs. (3), (4), (19)\u2013(22), (41), (45), and (46), the resultant Mw of the torques of the centrifugal inertial forces due to the crankshaft counterweights and its angle of action \u03c9w after adjusting the positional angles are as follows (Fig. 1c) (59) (60) where \u03b3 is the angle between the torque vectors Mci and Mcj, depending on the adjustment of the positional angles of the corresponding crankshaft counterweights (Fig. 1c) (61) Taking account of Eqs. (23)\u2013(25) and (61), it follows from Eq. (59) that (62) On the basis of Eqs. (57) and (58), we may write Eqs. (60) and (62) in the form, (63) (64) Mci = f 0\u03c92 f 3 f 4 \u03b1 f 0 \u0394\u03b1 i 1+( ) \u0394\u03b1i+( )\u2013[ ]cos\u2013{ }0.5; Mc j f 0c ji\u03c9 2= \u00d7 f 5 f 6 \u03b2 f 0 \u0394\u03b2 j 1+( ) \u0394\u03b2 j+( )\u2013[ ]cos\u2013{ }0.5. Mw Mci 2 Mc j 2 2MciMc j \u03b3cos\u2013+( )0.5 ;= \u03b1wtan Mci \u03d5isin Mc j \u03c8 jcos+( )= \u00d7 Mci \u03d5icos Mc j \u03c8 jsin+( ) 1\u2013 , \u03b3 90 \u03d5i \u03c8 j.+ += Mw Mci 2 Mc j 2 2MciMc j \u03d5i \u03c8 j+( )sin+ +[ ]0.5 .= Mw f 0\u03c92 f 3 f 4 \u03b1 f 0 \u0394\u03b1 i 1+( ) \u0394\u03b1i+( )\u2013[ ]cos\u2013{ }[= + c ji 2 f 5 f 6 \u03b2 f 0 \u0394\u03b2 j 1+( ) \u0394\u03b2 j+( )\u2013[ ]cos\u2013{ } + 2c ji f 3 f 4 \u03b1 f 0 \u0394\u03b1 i 1+( ) \u0394\u03b1i+( )\u2013[ ]cos\u2013{ }[ \u00d7 f 5 f 6 \u03b2 f 0 \u0394\u03b2 j 1+( ) \u0394\u03b2 j+( )\u2013[ ]cos\u2013{ } ]0.5 \u00d7 \u03d5i \u03c8i+( ) ]0.5;sin \u03b1wtan = f 3 f 4 \u03b1 f 0 \u0394\u03b1 i 1+( ) \u0394\u03b1i+( )\u2013[ ]cos\u2013{ }0.5 \u03d5isin[ + c ji f 5 f 6 \u03b2 f 0 \u0394\u03b2 j 1+( ) \u0394\u03b2 j+( )\u2013[ ]cos\u2013{ }0.5 \u03c8 j]cos \u00d7 f 3 f 4 \u03b1 f 0 \u0394\u03b1 i 1+( ) \u0394\u03b1i+( )\u2013[ ]cos\u2013{ }0.5 \u03d5icos[ + c ji f 5 f 6 \u03b2 f 0 \u0394\u03b2 j 1+( ) \u0394\u03b2 j+( )\u2013[ ]cos\u2013{ }0.5 \u03c8 j ] 1\u2013 .sin RUSSIAN ENGINEERING RESEARCH Vol. 27 No. 11 2007 COMPENSATING THE TOTAL UNBALANCED MASS 743 In the subsequent calculations, we must take into account that (Fig. 1c) (65) where \u0394\u03d5i and \u0394\u03c8j are the increments in \u03d50i and \u03c80j, respectively, due to adjustment of the positional angles of the corresponding crankshaft counterweights. Analysis of Eqs. (17), (18), (53), and (54) shows that adjustment of the counterweight positional angles is accompanied by increase in \u03d50i , \u03c80j , \u03d5i, and \u03c8j . In that case, the sum of \u03d5i and \u03c8j increases and consequently their sine increases. As a result, as is evident from Eq. (63), the corresponding torque increases. On the basis of Eqs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000583_9780470439098.ch3-Figure3.1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000583_9780470439098.ch3-Figure3.1-1.png", + "caption": "Figure 3.1. Schematic representation of a gel in its collapsed and swollen states.", + "texts": [ + " These systems have obvious attractiveness in basic science and are potential for many applications, for example, serving as artificial muscles, sensors, microrobots, micropumps, and actuators. The most extensive investigations on SPs have been carried out on the polymer gels in which a network of long polymer molecules holds the liquid in place and so gives the gel what solidity it has. Chemically cross-linked gels can undergo a transition between a collapsed and an expanded state, that is, between a shrunk and a swollen state (Fig. 3.1). The difference between these two states could reach several orders of magnitude. This behavior has been very attractive for the applications of the gels as potential actuators, sensors, controllable membranes for separations, and modulators for drug delivery (Osada and Gong, 1998). Studies on the light-responsive gels began as early as the 1970s. Van der Veen and Prins (1971) prepared a gel system consisting of a low-molecular-weight (LMW) chrysophenine dye (1) and a water-swollen gel of poly(2-hydroxyethyl methacrylate) cross-linked with ethylene glycol dimethylacylate" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000609_978-3-540-77296-5_33-Figure10-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000609_978-3-540-77296-5_33-Figure10-1.png", + "caption": "Fig. 10. Katana is used on a variety of mobile platforms for service robotics tasks", + "texts": [], + "surrounding_texts": [ + "Precision enhancement for object placement. Many industrial applications require to pick up objects from an array of 100 or more positions and, after the object is processed by a machine, to place it in a second array of the same or different dimensions. Not every of these points is taught-in, but only the position of the edges. The intermediate points are then calculated by the kinematic model of the system. However, the precision of these intermediate points is not very high and may differ from reality up to 1 mm. Industrial applications require often a position of a tenth of a millimeter. Thus, correction algorithms are required. A measuring device mounted on the end effector allows the robot to explore the working area. At each moment, Katana stores the configuration of the joint parameters and the values measured by the instrument in the memory. These values are then processed by a learning algorithm, which will extract the mapping between a position and the error. Based on this knowledge, the system is then able to compensate the position error for any point in the space. In [4], a neural network is used to improve the position accuracy of a robot manipulator. After the position errors for all grid points of a calibration board are identified, the network is used as an interpolator to determine the errors for any location within the calibration space. The training is executed on-line using the grid points in the neighborhood of the target position as training patterns. For the calibration process of Katana, a modified version of this approach is used. The network is trained off-line over the entire set of measured points in order to save processing time during the execution. Furthermore, it allows to better control the training process and eventually to repeat it until the desired performance is achieved. The learning algorithm is composed of a neural network, which has the ability to learn and generalize from previous experiences. The first step consists of collecting a set of samples which will be used to train the network. A sample is composed by the of the end effector. Then the network is trained using Quasi-Newton [5, 10] or steepest gradient backpropagation algorithm and a validation set is used for an early stop of the training process. After the training process has reached the convergence, the network is used in feed forward mode to estimate the errors of each new target position. Then, a false target position is computed and sent to the robot, which will eventually reach the desired point. However, simply subtracting the estimated error from the target position will not lead the end effector to the desired point. This problem is illustrated in figure 3, where the dotted circle represent the false target computed by subtraction of the error. At this point, the error vector may be completely different from that estimated, specially when the robot is close to some critical configurations. As a consequence the robot will stop at a point ('X') which is far from the desired. A search algorithm based on simulated annealing [6] and genetic algorithm [7] is then used to compute the best absolute position of each joint (in encoder steps) and the corresponding position error false point in the nearing of the target position which minimize the residual error. The cost function to be minimized, used to evaluate the partial solution is defined as: C=d\u2212DK et y et . (1) Where d is the desired position of the end effector, DK et is the searched false target and y et is the output of the neural network when the joint positions et at the false target are presented as inputs. The Quasi-Newton training method is faster than a standard backpropagation method because it uses, in addition to the gradient, second order approximation of the error surface [10]. To illustrate the efficiency of this method, a simple simulation is described. A training set containing 35 patterns is used for training a network with 6 inputs, 10 units in the first hidden layer, 5 units in the second hidden layer and one output. The validation set consists of 5 patterns. The mean square error, the mean absolute error and the number of epochs required for convergence are averaged over 10 learning sessions. The results are presented in table 2. Calibration Process. An example of the method described above is the calibration process for a restricted area in space. First, a subset of points to be measured is being defined, usually placed along a grid. Eventually, specific required points can be manually added or one can let the robot choose points at random in the working space. At each point the robot measures the position error, but only if a measuring equipment is available. Otherwise the user is asked to feedback the error by placing the end effector at the real point. Once a sufficient number of samples are available, the learning process may start. The number of validation patterns plays an important role for the quality of the solution. They are used for cross-validating the performances of the network and avoid over-fitting the training samples. The application (Katana4D user software, see Figure 6) shows two graphs representing the squared error over the training (upper) and the validation set (lower curve) of the patterns in function of the learning epochs. Once the learning has terminated, the network's weight matrix corresponding to the minimum point in the validation error is saved. This is the network with the best generalization propriety. The Quasi-Newton method clearly shows a superior performance, achieving better results than the steepest descent method in about 200 times epochs less. Figure 4 and 5 show the best solution found by both algorithms. The circles represents the desired outputs, while the crosses are the responses of the neural networks for the same inputs. Self-adaptation to the real space coordinates. In most industrial applications, only a small part of the complete robot's working space is exploited. There are, in general, two or three bounded regions where the robot has to perform a task with high accuracy, the remaining portion is used by the manipulator as a transitory space. A single calibration process that is valid for the whole working space would require a large amount of time. First, hundreds of calibration points have to be collected, then a neural network of medium to large size would be necessary to generate the complex input output mapping. To overcome this problem, the application allows the creation of multiple models of local areas. In this way, a neural network may be trained only with patterns belonging to a limited region, creating a more accurate error estimator. When the robot is active the joint positions are constantly monitored and, when they fall into the input boundaries of a neural network, the position of the end effector is compensated. The environment often moves at least minimally, leading to errors in handling and thus reducing the reliability of the system. A visual system is a powerful possibility to provide autonomous feedback by continuous screening of the working space. Inverse kinematics, visual object recognition and neural error compensation as described above are combined to provide a self learning and robust pick & place system. Speech functions. Katana can listen and talk to the user. These functions make use of the improvements during the last years in both, the language understanding and the improvement of the comprehension based on training and constraints. Compared to manual control of the robot (e.g. by leading it at the forearm to a desired place, using a joystick or a keyboard), the language interface reacts slower for two reasons: 1) The word or sentence takes time to be pronounced, and the end has to be detected. 2) The danger that a wrong action is taking place because of an undesired command which the robot gets (e.g., from a dialog between humans) requires a confirmation before execution of the command. However, the language functions can be very useful to start and stop full programs or for disabled persons to interact with the robot. Katana is employed, in two major fields: a) industrial manufacturing, assembling and quality control and b) in service robotics. The applications in research and education today often target the service robotics field, although there are many interesting research topics which address fixed-place industrial applications or both. The use of the algorithms described above is different in these two major fields: where in the industrial applications the Katana arm is the main component, which also may serve as master for the control of the connected machines and automation peripherals, in service robotics the arm is often the slave which is controlled by the moving platform which may contain computer power higher or equal to that of the Katana arm. Industrial applications. The Katana robotic arm is ideally suited for pick & place, assembling and quality control tasks of light-weighted objects. Meanwhile, the system is used world-wide by small-sized, medium and large companies like BMW, Intel, Unilever, Mettler-Toledo or Maxon Motors for a variety of applications. The main advantages are the high flexibility and easiness of its use, both based on the security 2.2 Fields of Applications Service robotics. Due to the low power consumption, Katana can also be an exceptional system component in mobile robotics. The control system architecture described above makes it an interesting component for several reasons: The concept of distributed intelligence can be realized under many different aspects; the powerful and standalone real-time Linux environment opens a wide range of applications which may include movements along trained trajectories, interactive tasks with humans and other robotic systems as well as robust behavior in unknown environments based on the sensory systems available on the arm and its grippers. aspects as well as on the combination of standard and artificial intelligence powered control and interface solutions." + ] + }, + { + "image_filename": "designv11_83_0001778_08ias.2008.33-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001778_08ias.2008.33-Figure3-1.png", + "caption": "Fig. 3. Series-coupling connections.", + "texts": [ + " If the induction machine is modelled as a transformer and the equivalent circuit is represented as a transformer equivalent circuit then transformertype characterisation tests can be carried out on the machine. The test procedures include the open-circuit tests and the series-coupling tests. There are two open-circuit tests that are to be carried out on the machine: the open-circuit rotor (OR) test and the open-circuit stator (OS) test. The series-coupling tests involve electrically connecting the stator (primary) coil to the rotor (secondary) coil, as is shown in Fig. 3. This allows for two different measurement set ups to be achieved. The first is the cumulative (series-aiding) test and the second is the differential (series-opposing) test. The cumulative test occurs when the stator and rotor poles of the machine align to give an algebraic summing of the machine fluxes that corresponds to a maximum inductive point. This condition will occur naturally if a voltage is applied and the rotor is free to move. For the differential test the poles of the machine must be aligned in such a fashion that an algebraic summing of the machine fluxes that corresponds to a minimum inductive point is found" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000189_s11003-006-0038-0-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000189_s11003-006-0038-0-Figure1-1.png", + "caption": "Fig. 1. Schematic diagram of the contact problem: (1, 2) bodies.", + "texts": [ + " The distribution of stresses obtained with regard for convective heat exchange is compared with the solutions of the problems with thermal insulation of the unloaded surface and with the solutions of the corresponding isothermal problem of the theory of elasticity [16]. Bia\u0142ystok Polytechnic Institute, Poland. Translated from Fizyko-Khimichna Mekhanika Materialiv, Vol. 41, No. 6, pp. 26\u201332, November\u2013December, 2005. Original article submitted May 30, 2005. 734 1068\u2013820X/05/4106\u20130734 \u00a9 2005 Springer Science+Business Media, Inc. Consider the problem of contact interaction of two elastic bodies (Fig. 1). The first body is pressed to the second body with a force P and rotates about the axis of symmetry with an angular velocity \u03c9. The intensity of heat generation caused by friction in the contact region is equal to the power of the friction forces, i.e., Q ( r ) = f \u03c9 a r p ( r ), r < 1, where f is the friction coefficient, p ( r ) is contact pressure, a is the radius of the contact region, and r and z are dimensionless cylindrical coordinates related to the radius of the contact region. The thermal contact of the bodies is regarded as perfect" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002262_tia.2021.3058113-Figure15-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002262_tia.2021.3058113-Figure15-1.png", + "caption": "Fig. 15. Various views of the CRAMB prototype.", + "texts": [ + " The experimental results are obtained from a prototype CRAMB (see Fig. 2) Authorized licensed use limited to: Carleton University. Downloaded on June 04,2021 at 02:56:52 UTC from IEEE Xplore. Restrictions apply. 0093-9994 (c) 2021 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. that implements the externally biased three-pole bearing as its radial stage (shown in Fig. 15). A mechanical swivel bearing allows the shaft end opposite the magnetic bearing to spin and pivot but constrains it from displacing radially. While the prototype is designed with an optimal bias field (\u03b6 = 0.569), a compensation coil has been installed that can be re-purposed to modify the bearing\u2019s bias field when conducting static force tests. Full details of the prototype can be found in [5], [24]. Key machine parameters are provided in Table I, which includes the equivalent force model values needed to solve (4) and utilize the non-dimensional current map of Fig. 13b. Bearing suspension instabilities that can result from using either the linear control approach or a poorly biased bearing are now investigated. Two bearing models are used for the simulations: 1) the bearing prototype of Fig. 15; 2) a reduced bias version of the prototype bearing with \u03b6 = 1 4 (so that the exact force vector regulator will command discontinuous coil currents, as explained in Section III-D). These bearing models use the current maps of Fig. 13b and 13c with the machine parameter values listed in Table I. In all simulations, identical PID and PI controller gains (see Fig. 12) were used with a 700 Hz current controller bandwidth and 100V dc inverter bus. 1) Full motion control (startup and step disturbance): Fig", + "3], this phenomenon will pose stability challenges by making the system sensitive to inacuracies in the imbalance control suppression and because such strategies are typically applied above a minimum threshold shaft speed. Three sets of tests were run on the prototype to validate the exact force vector regulator and compare its performance to the standard linear force regulator. 1) Static Force Test: The case study of Section III-C1 and Fig. 5b was investigated experimentally by applying the minimum L2 coil currents of ~ic = [ic,1,\u2212 ic,1 2 ,\u2212 ic,1 2 ] and measuring forces produced. To complete this test, the CRAMB is placed in a commercial milling machine (right side of Fig. 15), with the rotor installed in the mill\u2019s spindle and the stator fixed to the mill\u2019s x-y table via a load cell. The test results are shown in Fig. 19, where it can be seen that the \u03b6 = 1 4 profile matches Fig. 5b and has the expected minimum force value and corresponding coil current anticipated by the analysis framework of Section III-B and Fig. 6. Also as expected, the optimally biased bearing does not encounter a minimum force limit on this solution within its force profile. 2) Startup and Rotational Tests: The prototype machine was operated using the proposed exact force vector regulator at speeds up to 1300 RPM" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001434_s11249-007-9276-z-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001434_s11249-007-9276-z-Figure1-1.png", + "caption": "Fig. 1 Geometry of EHL of circular contacts under pure squeeze motion", + "texts": [ + " The coupled transient modified Reynolds, elasticity deformation, and ball motion equations are solved simultaneously using the finite difference method and the Gauss-Seidel iteration method. The effects of flow rheology of the lubricant on the elastic deformation and the performance of a squeeze film are discussed under the conditions of impact and rebound. 2.1 Modified Reynolds Equation The situation of two spheres approaching one another may be expressed as the equivalent sphere approaching a plane. Consider the squeezing film mechanism as shown in Fig. 1, where an elastic sphere of radius R is approaching in an infinite plate with a velocity. The lubricant in the system is taken to be a nonNewtonian powerlaw lubricant. The equations of motion governing the axial symmetric flow of a compressible fluid under the assumptions of the lubrication theory, and neglecting the inertia terms, in one dimension, are given by dp dr \u00bc os oz \u00f01\u00de dp dz \u00bc 0 \u00f02\u00de Since the flow occurs under the pressure gradient only and the pressure is constant across the film. Mathematically, the constitutive equation of power law fluid under the present axial symmetric problem is expressed as [22] s \u00bc m ovr oz n 1 dvr dz \u00f03\u00de where s is the shear stress, qvr/qz is the shear rate, m is the consistency index, and n is the flow index" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002801_s00006-021-01119-6-Figure10-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002801_s00006-021-01119-6-Figure10-1.png", + "caption": "Figure 10. Screw term representations of a typical 3- limbed manipulator and the vectors of X l bl nl and X l 0 nl of limbs 1 and 2", + "texts": [ + " (75) are being retained/preserved and expressed as coefficients of a G6,0 multivector, we have u1 = p1, u2 = p2, u3 = p3, m4 = p4, m5 = p5 and m6 = p6. The reason for defining the line as a multivector in G6,0 is so that the maximum grade is one instead of two as in Eq. (75). Although the bivector representation for the line (as presented in Eq. (75)) is more compact and informative, the method developed in [47], which is the basis of the affectation index in this work, was defined in G6,0 and we use this signature. With reference to Fig. 10 below, given a particular end effector position (XE), Ll is thus defined as Ll = [ X l bl nl ,X l bl nl \u00d7 X l 0 nl ] (77) which is in agreement with [50], where l refers to the l-th limb of a parallel manipulator, bl refers to the base of the l-th limb and nl refers to the joint of the l-th limb that is connected to the end effector. The first tuple of the 4-tuple position element Xi j is removed to form a 3-tuple vector. In Eq. (77), the actuation line is a concatenation of two 3-tuples. The first 3-tuple (X l bl nl = (p1, p2, p3)), represents the linear components, where p1, p2 and p3 are the values of each tuple in X l bl nl " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001298_062-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001298_062-Figure2-1.png", + "caption": "Figure 2. Schematic model of thin film temperature sensor.", + "texts": [ + " For applications that require protection from the operating environment, a protective overcoat, typically aluminium oxide, is sputter deposited or evaporated onto the sensor to a thickness of approximately 2\u20133 \u00b5m. Figure 1 shows a thin film sensor developed on a glass disc and the inspection result of the sensor. The inspection was to check the shape of the sensor and to measure its dimension. The sensor measured in figure 1(b) is about 10 \u00b5m in width, 400 \u00b5m in length and 0.4 \u00b5m in thickness. Consider a parallelepiped sensor of a sandwich structure composed of three layers: substrate, sensing material and protecting layer (see figure 2). Over the protecting layer, there is the lubricant film thickness in which internal heat is generated due to friction. Part of this heat generation is absorbed by the sensor, crossing the different inner layers by conduction (the rest is both absorbed by the rotating ball and removed by the lubricant in a convective way). To obtain the temperature distribution, a transient, three-dimensional heat conduction equation must be solved in the selected computational domain with appropriate boundary and initial conditions", + " For the x direction yields\u222b t+ t t \u222b CV \u2202 \u2202y ( k \u2202T \u2202y ) dV dt = \u222b t+ t t [( kA \u2202T \u2202y ) e \u2212 ( kA \u2202T \u2202y ) w ] dt = \u222b t+ t t [ keAe TE \u2212 TP \u03b4EP \u2212 kwAw TP \u2212 TW \u03b4PW ] dt = \u222b t+ t t [aeTE + awTW \u2212 (ae + aw)TP ] dt (9) where ae = keAe \u03b4EP and aw = kwAw \u03b4PW . The following geometrical nomenclature is utilized to identify the six neighbours of each point: E, W, N, S, T and B stand for east, west, north, south, top and bottom nodes, respectively. A schematic of the geometrical nomenclature can be observed in the right-lower corner of figure 2. Distances between points P\u2013W and P\u2013E are represented by \u03b4PW and \u03b4EP, respectively. Performing the same analysis for the x and z directions, the diffusion term can be written as\u222b t+ t t \u222b CV \u2202 \u2202xi ( k \u2202T \u2202xi ) dV dt = \u222b t+ t t (\u2211 NB aNBTNB + aP TP ) dt (10) where the subscript \u2018NB\u2019 denotes neighbour value. The coefficients aNB have the form aNB = kNBANB \u03b4NB for NB = E, W, N, S, T, B. (11) The coefficient aP has the form aP = \u2212(aE + aW + aN + aS + aT + aB). (12) For the time integration the \u2018generalized trapezoidal rule\u2019 is used: \u222b t+ t t aiTi dt = [ \u03b8 ai Ti + (1 \u2212 \u03b8)a0 i T 0 i ] t (13) where the subscript i denotes any of the coefficients aNB or aP and the coefficient \u03b8 is a weighting factor varying from zero (explicit scheme) to one (implicit scheme), the Crank\u2013Nicholson scheme considers \u03b8 = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000332_50003-4-Figure3.44-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000332_50003-4-Figure3.44-1.png", + "caption": "Fig. 3.44. Development of the vehicle motion after the rear wheels get locked while the front wheels remain rolling freely.", + "texts": [ + " Consequently, the final deviation with respect to the original rectilinear path can be kept within any chosen limit and, strictly speaking, the actual system is always stable. Finally, the solutions for the non-linear system governed by the Eqs.(3.118, 3.119) and tyre characteristics according to Fig.3.41 have been established by numerical integration of the equations of motion. Now, the character of the motion may change with the level of the initial disturbance. In Figs.3.44 and THEORY OF STEADY-STATE SLIP FORCE AND MOMENT GENERATION 155 3.45, the resulting motions for two cases have been depicted. The first, without braking the front wheels (Fig.3.44) and the second with also the front wheels being braked but at a lower effort to make sure that they roll at least initially. The initial speed of travel and disturbances have been kept the same for all cases but the coefficient of friction/1 has been varied. At lower friction the time available to develop the angle of swing is larger and we see that an angle of more than 180 degrees may be reached. Then the locked wheels are moving at the front (u <0) and we have seen that that situation is stable while the motion can become oscillatory (cf. Fig.3.42). For the case of Fig.3.44 with/~ = 0.25 we indeed observe that a sign change occurs once for the yaw rate r. When the brakes are applied also at the front axle, the deceleration is larger and we have less time to come to a stop. Consequently, the final angles of Fig.3.45 are smaller than those reached in Fig.3.44. The shaded zones indicate the ranges where also the front wheels get locked. This obviously occurs when the slip angle of the initially rolling wheels 1 becomes sufficiently large and exceeds the value indicated in Fig.3.41 where the inclined straight line is reached. It has been found (not surprisingly) that the influence of raising the initial speed is qualitatively the same as the effect of reducing the coefficient of friction." + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002138_s11668-020-01090-7-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002138_s11668-020-01090-7-Figure3-1.png", + "caption": "Fig. 3 Typical dynamic model of a gear pair", + "texts": [ + " As shown in Fig. 2(a), a typical GRS consists of gear pairs, bearings, elastic shafts, drive motors and loads. Figure 2(b) shows the dynamic model of the GRS. In this dynamic model, the rotating shafts of the GRS are modeled as Timoshenko beams; the meshing gear is modeled as a pair of rigid disks connected by a spring\u2013damper with timevarying mesh stiffness. The detailed modeling process of the rotating shaft and gear pair is introduced as follows. Modeling of a Gear Pair System As illustrated in Fig. 3, a nonlinear dynamic model of a gear pair system with ten DOF is utilized in this paper. A pair of meshing gears are modeled by rigid disks representing their mass/moments of inertia. The disks are linked by line elements which represent the stiffness and the damping. The coupled dynamic equation in a matrix form can be expressed as follows Mg \u20acqg \u00fe \u00bdGg \u00fe Cg _qg \u00feKgqg \u00bc Qg \u00f0Eq 1\u00de where Mg, Cg, Gg, Kg are the inertia, damping, gyroscopic and stiffness matrix, respectively. Qg is the external force vector in the local pressure line co-ordinate system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003572_s10015-021-00694-y-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003572_s10015-021-00694-y-Figure2-1.png", + "caption": "Fig. 2 Proposed robot (CIPbot-1)", + "texts": [ + " The most challenge of development this robot is the space limitation and structure rigidity. The locomotion mechanism must be able to adjust its diameter for the target pipes and the manipulator must be powerful enough for its duties, while its workspace should cover all target area in the pipe as well. Meantime, its structure needs to be able to bear its mass and reaction force occurred, while the manipulator is operating. These mechanisms must be small enough to pass through the 150\u00a0mm pipe. The prototype and specifications of the CIPbot-1 are shown in Fig.\u00a02; Table\u00a01. This section consists of a sixfolding mechanisms and six crawlers, as shown in Fig.\u00a02. Each of them is driven by only one powerful motor. Both rotations are transferred by a gearbox at the center of the robot. The combination of wall-press type and the crawler type are used to achieve an expandability and mobility. There are 3 pairs of folding mechanisms, all 3 pairs are driven by a motor (24\u00a0V 22\u00a0W 52.1mNm 4600\u00a0rpm), the rotation is transferred through the gear train (gear ratio 1875:1) to drive each linkage. By this way, the robot can be adjusted the diameter and contact force" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002519_s42835-021-00708-6-Figure10-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002519_s42835-021-00708-6-Figure10-1.png", + "caption": "Fig. 10 Schematic of HTS field coil test setup for 35\u00a0K testing", + "texts": [ + " The measured electric field (\u00b5V/cm) as a function of DC current is shown in Fig.\u00a012. Under the 1\u00a0\u03bcV/cm Fig. 7 Magnetic field plot (3D) of HTS field coil 1 3 criterion of superconductivity, the measured critical current of HTS field coil was \u2248 63 A, which is relatively close to the theoretical estimation of 61 A. The key goal of testing the HTS field coil at 35\u00a0K using cold helium gas is to find out the critical current of the developed HTS coil at self field. The schematic diagram of the experimental setup for testing HTS field coil at 35\u00a0K is shown in Fig.\u00a010. The test facility comprises of a helium gas chamber, vacuum chamber, Stirling cryocooler (Model: SPC 4\u00a0T, Make: Stirling Cryogenics) of rated cooling capacity of 300\u00a0W at 20\u00a0K, low-voltage high-current power source capable of providing DC current up to 1000 A, dump resistor (0.1 \u2126, 500\u00a0V), cryogenic temperature monitor and a multichannel digital data acquisition system including devices such as milli voltmeter, milli Ohm meter and vacuum gauge [4]. Two calibrated platinum resistance sensors (Pt102) were mounted on the top and bottom of the HTS field coil for measuring the temperature" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002949_10775463211022489-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002949_10775463211022489-Figure1-1.png", + "caption": "Figure 1. Schematic of a flexible rotor partially filled with liquid: (a) longitudinal section view and (b) cross-sectional view.", + "texts": [ + " Thereafter, numerical results based on the analytical solutions are given for a better illustration of how the parameters influencing the critical spinning speed of a liquid-filled rotor system. Consider the steady-state vibration of a flexible rotor partially filled with liquid. It is supposed that the rotor was mounted with hinged ends. Also, the rotor rotates with constant angular velocityV around its longitudinal axis and undergoes a steady whirl with an amplitude of \u03b5(x) and an angular frequency of \u03c9, as shown in Figure 1. Comparing with the centrifugal force, the gravity force can be neglected. In the steady state, the liquid and the rotor spin as a single whole. The related parameters of the rotor and the coordinates are shown in Figure 1. In the rotating frame c-x\u2019y\u2019, the perturbation velocity and pressure of the fluid particles can be written as follows \u03c5r \u00bc ur0 \u00fe v\u03b80 (1) Ptotal \u00bc \u03c1fV 2 r2 r2i 2 \u00fe pF\u00f0r, \u03b8, t\u00de (2) where r0, \u03b80 are the unit vectors, and u and v are the radial and tangential velocity components of the fluid particles. Ptotal is the total pressure, which is composed of centrifugal force and the perturbed pressure pF. \u03c1f is the fluid density. According to the acceleration composite theorem of point in theoretical mechanics, the absolute acceleration of fluid particle can be defined as aa \u00bc ae \u00fe ar \u00fe ak (3) where ae is the convected acceleration, ar is the relative acceleration, and ak is the Coriolis acceleration" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000287_3-540-31761-9_31-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000287_3-540-31761-9_31-Figure2-1.png", + "caption": "Fig. 2. Definitions: Mechanical system with (a) and without Grosch-wheel (b).", + "texts": [], + "surrounding_texts": [ + "In literature many roughness descriptors have been proposed. Parallel to the standard height parameters (roughness average Ra, Rp, Rz, etc.) the most descriptors used are the amplitude distribution function ADF parameters (kurtosis, skewness), Schulze-surface-index and Eichhorn-parameters [7]. An investigation in [8] showed that the first group does not have a good correlation with the friction coefficient. Schulze [12] and Eichhorn [6] describe in different ways the capability of road surfaces to establish contact to rubber. In the present work ADF based parameters, the Bearing Ratio Curve descriptors, are introduced to this application field. They are known in the field of metal manufacturing and are used to evaluate the quality of surfaces. This approach (ISO 13565-1996) divides the cumulative ADF into three sections: a central one called main plateau, a section of peaks and a section of valleys. The used BRC-descriptors are main roughness Rk, peak roughness Rpk, valley roughness Rvk, fraction of peaks MR1 and the fraction of valleys MR2. The fractal descriptors build up the second group. Their is due to the ability of giving the power content of the measured roughness profile in dependency of the wave length. This is valid with the assumption of self affinity which is correct for a wide range of road surfaces [9]." + ] + }, + { + "image_filename": "designv11_83_0002993_s00170-021-07405-8-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002993_s00170-021-07405-8-Figure2-1.png", + "caption": "Fig. 2 Cold extrusion process diagrams of sun gear. a Extruding internal spline. b Extruding external gear", + "texts": [ + " Thus, the objectives of this study are (i) to propose the precision sizing process as a postoperation to enhance the forming quality of sun gears formed by cold extrusion process, and (ii) to investigate the influences of important process parameters such as the sizing amount of external gear, interference value of internal spline, die bearing length, and friction factor on dimensional accuracy of internalexternal teeth to determine the optimal parameters for commercial production of sun gears and future research. Sun gear is a kind of typical gear-like part, which is widely used in the truck transmission system. The geometry and main parameters of target sun gear in this study are shown in Fig. 1 and Table 1, respectively. It is observed that the sun gear has two kinds of tooth shapes, namely external gear and internal spline, respectively. Figure 2 shows the proposed cold extrusion process of sun gear in previous investigation. First, a ring-shaped billet 1, formed by hot-forged, was inserted into the stationary splined mandrel and die. Then, billet 1 moved downward by the punch, and the internal tooth shapes were gradually formed by the splined mandrel. When billet 1 was going to pass through the die cavity completely, billet 2 was placed on billet 1. Then, billet 1 and billet 2 were pressed by the punch, and the internal tooth shapes of billet 1 were formed completely, as displayed in Fig. 2a. Second, the forming principle of external gear was the same as that of internal spline. The billet with internal spline was put into the toothed die container and extruded by the splined punch so that the external gear was formed, as illustrated in Fig. 2b. The formed gears obtained by the experiments of cold extrusion process are shown in Fig. 3, in which the first-formed end of sun gear was defined as the small tooth width, and the big tooth width represents the last formed end. The accuracy of internal-external teeth was inspected using the gear measuring center, as denoted in Fig. 4. The total profile deviation (F\u03b1) and total helix deviation (F\u03b2) are two key indexes of external gear accuracy, and the total M value deviation (FM) represents the accuracy of internal spline" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002430_012006-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002430_012006-Figure3-1.png", + "caption": "Figure 3. Structure diagram of Pitch chain", + "texts": [], + "surrounding_texts": [ + "This article mainly focuses on the inverse dynamics problem. When solving the inverse kinematics, the speed and acceleration of the sliders and the links will also be obtained. The pitch chain is similar to the yaw chain, so the inverse kinematics takes the yaw chain as an example. In order to facilitate the analysis, the yaw chain is simplified as a planar crank slider mechanism in the plane. Its closed vector diagram is shown in Figure 4. 1) Position inverse solution By the closed vector method, we can establish the closed vector equation as: 1 1 1 1 1 1 1 3q l e n u oR R R (5) So the inverse solution of the yaw slider position is: MEMAT 2021 Journal of Physics: Conference Series 1820 (2021) 012006 IOP Publishing doi:10.1088/1742-6596/1820/1/012006 T 1 1 1 3 1 1 1 1 1 3 1 1 1( ( )) ( ( ))q l l oR R R n u oR R R n u (6) 2) Slider\u2019s speed By differentiation of equation (5) with respect to time, we can get: 3 1 1 1 1 1 1 3 \u02c6 \u02c6( ) ( )R q l v e Eu ER R (7) Multiplying the two sides of equation (7) with 1u to get: T 1 1 3 1 1T 1 1 \u02c6( ) 0 =q u ER R J u e (8) Where 3Rv is the speed of the center of mass of the rotating pair 3R ; 1 is the angular velocity of the yaw link in the base coordinate system; 0 1\u02c6 = 1 0 E ; 1J is the Jacobian matrix of the yaw chain. 3) Link\u2019s speed Multiplying the two sides of equation (7) with 1E\u0302u , and substituting equation (8) leads to: T T T 1 1 1 1 3 1 1 3 T 1 11 1 1 \u02c6 \u02c6(( ) ) ( )\u02c6 \u02c6( ) ( )1 = 0 l Eu e u ER R Eu ER R Ju e (9) According to equation (7), the speed of the center of mass of the link is: 3 1 1 1 1 1 1 3 1 1 1 \u02c6 \u02c6 \u02c6( ) 0 ( ) 2 2R v l l v v Eu ER R Eu J J (10) Where: 1vJ and 1J are the linear velocity Jacobian matrix and the rotational angular velocity Jacobian matrix of the link. 4) Slider\u2019s acceleration Differentiating both sides of equation (7) with respect to time leads to: 2 2 1 1 1 1 1 1 1 1 1 3 1 3 \u02c6 \u02c6( ) ( )q l l e Eu u ER R R R (11) Multiplying the two sides of equation (11) with 1u to get: 2 T 2 1 1 1 1 1 1 3T 1 1 1 = ( )q l J u R R u e (12) 5) Link\u2019s angular acceleration Multiplying the two sides of equation (11) with 1E\u0302u to get: T T T 1 1 1 1 3 1 1 3 T 1 1 2 1 1 21 1 1 1 1 \u02c6 \u02c6(( ) ) ( )\u02c6 \u02c6( ) ( )1 1 1 = ( ) ( ) 0 l l l Eu e u ER R Eu ER R Ju e D D D D (13) Where T 2 T 21 1 1 1 1 1 1 3T 1 1 \u02c6( ) = ( )l D Eu e u R R u e ; T 2 2 1 1 3 \u02c6( ) D Eu R R . 6) Centroid acceleration analysis of link Differentiating both sides of equation (10) with respect to time leads to: 2 21 1 1 1 1 1 2 1 3 1 \u02c6 = ( ) 2 2v l Eu v J R R uD D (14)" + ] + }, + { + "image_filename": "designv11_83_0002679_14644207211012726-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002679_14644207211012726-Figure4-1.png", + "caption": "Figure 4. Deviation from the nominal geometry: maximum material, best fit and least material diameters, cylindricity, and runout.", + "texts": [ + "33 In the present case, the samples were inspected by three-dimensional scanning with Blue LED Fringe Projection, using a Zeiss Comet 5M Scanner. Measured data were compared to the CAD files in Geomagic Control X. For each strut, the diameter of the cylinder that had the best fit for the measured points was calculated (BF). Similarly, maximum material (MM) and least material (LM) cylinders were defined as the diameters of the minimum circumscribed cylinder and the maximum inscribed cylinder, respectively, as shown in Figure 4. Geometric Dimensioning and Tolerancing (GD&T) was used to characterize the deviations due to the strut waviness and misalignment. First, cylindricity was assessed, controlling the circularity and the straightness of the strut, by evaluating the spacing of two concentric cylinders between whose the strut surface must lie, as shown in Figure 4. The deviation of the strut surface with respect to the nominal geometry was evaluated by using the Total Runout as well, i.e., by considering two concentric cylinders being coaxial with the nominal strut axis. The surface roughness was qualitatively assessed by comparing the samples using optical microscopy. In addition, some bulk samples (Figure 3) were fabricated and inspected, with a focus variation microscope Keyence VHX, obtaining the profile of the surface and calculating a typical surface roughness indicator, as follows Ra \u00bc 1 l Z l 0 z\u00f0x\u00dedx (1) Finally, quasi-static compression tests were performed on octet-truss blocks to assess the mechanical properties" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000828_3.43615-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000828_3.43615-Figure3-1.png", + "caption": "Fig. 3 Aerodynamic moments about the neutral point.", + "texts": [ + " 2, the two equations involving the two unknowns, the trim angle of attack and control surface deflection, can be written as C Lcr^tri w C A ~f~ C o (1) Equation (1) can be solved for the trim angle of attack and control surface deflection. The neutral point is defined as that point on the longitudinal axis of the body where the aerodynamic moment does not depend on angle of attack, f Transferring the equilibrium force system to the neutral point should therefore simplify the moment equation. The force system transferred to the neutral point is shown in Fig. 3. The prime on the moment indicates that the moment is that about the neutral point. The neutral point is fixed for a given configuration at the distance hnl from some reference position. The term \u00a3n is the static margin, that is, the nondimensional distance between the neutral point and the e.g. The thrust and drag forces are omitted from this figure. f The stick-fixed case only will be considered. The method can also be adapted to the stick-free case. Figure 4 is the force system of Fig. 3 nondimensionalized. The moment equation, as pointed out previously and as seen in Figs. 3 or 4, does not now depend on the angle of attack. Compared with Fig. 2, Fig. 4 more clearly shows the effect of e.g. changes or changes in the configuration for a given configuration whose aerodynamic coefficients are linear with angle of attack and control surface deflection. There is no difference in the results of analysis based on either Fig. 2 or 4. Figure 4, however, can be discussed without writing down the equations (which is a handy device for use in discussions of qualitative data)", + " As a second example, consider the effect of the control surface location (at different longitudinal positions along the body) on the trim angles of attack and control surface deflection. For this example, let the control force be considered separate from the force of the rest of the configuration. In particular, let the subscript c on a force or moment indicate the force or moment due to the control surface. The subcript \u2014 c on a force or moment will indicate the force or D ow nl oa de d by U N IV E R SI T Y O F M IC H IG A N o n Fe br ua ry 9 , 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /3 .4 36 15 (C0_;(C0_ Fig. 5 (see fig.3) Effect of e.g. location on trim angle of attack and control surface deflection. moment of the configuration less that of the control surface. The lift and moment or the configuration are then L = L_c + Lc M = M-c + I (2) where lc is the distance between the center of application of the control force and the point about which the moment is written. The force and moment system is shown in Fig. 6. The force and moment system less the control surface, transferred to the neutral point of the configuration less the control surface (NP)-" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001007_s1068798x08120113-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001007_s1068798x08120113-Figure4-1.png", + "caption": "Fig. 4. Nonlinear dynamic core of the elastic-smoothing system based on a multifunctional turning and milling center: (1) indenter; (2) surface; (3) support; (4) indenter guide pieces; R, indenter radius; m1, indenter mass; P, smoothing force; C1, rigidity coefficient of the tool\u2019s spring; B1, viscous-friction coefficient; C0, rigidity coefficient of the indenter\u2013blank contact; Ffr, frictional force in the indenter guide pieces; hr, height of the ridge of plastically deformed material at the surface.", + "texts": [ + "25Lp, we write the energy stored per oscillation cycle of the indenter as or in the first approximation Wst \u2248 and, cor- respondingly, Wst \u2248 . Hence, the main parameters influencing the energy stored per oscillation cycle are C0, k2, y, and hmax. Thus, in the dynamic machine-tool system (Fig. 1) with elastic smoothing of parts in a dynamically rigid turning and milling center, we may isolate a nonlinear dynamic core, which permits adequate identification of the self-oscillatory process. This core is shown in Fig. 4. For analysis of the dynamic characteristics of smoothing on the basis of the computer model of the nonlinear dynamic core within Vissim 5 software, it is expedient to plot a multiple-sheet phase portrait corresponding to the elastic-smoothing tool, with the following parameters: m = 0.025 kg; B = 40 N m/s; P = 120 N; Ffr = 5 N; K1 = 7 \u00d7 1012 N/m3, K2 = 0.4, R = 3 mm [4]. A multisheet phase portrait may be constructed in the coordinates Y, dy/dt by the attribution method [2]. Analytical solution of Eqs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001623_sice.2008.4654762-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001623_sice.2008.4654762-Figure3-1.png", + "caption": "Fig. 3. A simple friction model of the free rolling tire.", + "texts": [ + " Although, there\u2019s a difficulty to get experimental data that enables to guess the relationship between slip ratio and friction, the concept of relationship between slip ratio and friction is very important to estimate maximum friction coefficient while normal driving. It is because, in a normal driving condition, rolling effect is bigger than slip effect, we can not observe maximum friction coefficient until braking. Because (4) is a function of only normal force and traction force, it is relatively simple to accomplish real time estimation. Generally, the tire-road friction is formulated as an equation of slip ratio, camber angle, and normal force. Fig. 3 shows the single tire model of a vehicle. And assume that Fig. 3 is a simple model of a free rolling driven wheel. For that system, the slip ratio is defined as w xw Rw VRw \u2212 =\u03bb (5) where, R is a radius of a tire, xV is the longitudinal velocity, ww is the driven wheel angular velocity. An encoder measures the wheel angular velocity and the longitudinal velocity is assumed to be equal to the car velocity. - 784 - Driving wheel dynamics are expressed as (6). NrTrFTJ wmwxmww )(\u03bb\u00b5\u03c9 \u2212=\u2212= (6) Finally, friction estimation model for a tire can be obtained by (7). Nr JT w wwm \u03c9 \u03bb\u00b5 \u2212 =)( (7) where, mmm IKT = is a wheel driving torque that is obtained by input current to the driving motor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000583_9780470439098.ch3-Figure3.32-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000583_9780470439098.ch3-Figure3.32-1.png", + "caption": "Figure 3.32. (a) Photoinduced rolling motion of the continuous ring of the LCE film", + "texts": [ + " prepared LCE films with a densely cross-linked, twisted configuration of azobenzene moieties (Harris et al., 2005b). The films showed a large amplitude bending and coiling motion upon exposure to UV light, which results from the 901 twisted configuration of the LC alignment (Fig. 3.31). Most lately, Ikeda and coworkers prepared a continuous ring of the LCE film by connecting both ends of the film (Yamada et al., 2008). The azobenzene mesogens were aligned along the circular direction of the ring. As shown in (a) 21a + 21c and (b) 21b + 21c. Source: Mamiya et al., 2008. Fig. 3.32a, upon exposure to irradiation with UV light from the downside right and visible light from the upside right simultaneously, the ring rolled intermittently toward the actinic light source, resulting almost in a 3601 roll at room temperature. This is the first example of this kind of photoinduced motion in a single-layer film, although the rolling of the LCE ring shown here was slow and stopped when the ring was broken by irradiation. Furthermore, they prepared a plastic belt of the LCE-laminated film by connecting both ends of the film, and then placed the belt on a homemade pulley system as illustrated in Fig. 3.32b. By irradiating the belt with UV light from top right and visible light from top left simultaneously, a rotation of the belt was induced to drive the two pulleys in a counterclockwise direction at room temperature as shown in Fig. 3.32c. A plausible mechanism of the rotation is as follows: Upon exposure to UV light, a local contraction force is generated at the irradiated part of the belt near the right mation of an LCE sample upon exposure to light at l= 514 nm. (c) Mechanism of the locomotion of the dye-doped LCE sample. Source: Camacho-Lopez et al., 2004. pulley along the alignment direction of the azobenzene mesogens, which is parallel to the long axis of the belt. This contraction force acts on the right pulley, leading it to rotate in the counterclockwise direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003453_s10846-021-01454-7-Figure10-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003453_s10846-021-01454-7-Figure10-1.png", + "caption": "Fig. 10 Physical model of the proposed robot at steering", + "texts": [ + ", RF of 2:1, 3:1, 4:1, and 5:1) of left fin and (1:1) of the right one. The robot motion is recorded by a Kodak camera with high resolution specifications at 30 frame per second (fps). The recording camera, is fixed on the top-side of the swimming pool at 50 cm distance from the robot to capture the motion of the robot. A three black-labels fixed on the robot\u2019s body for motion tracking. Motion commands are sent to the controller via the HC-06 Bluetooth module, four 1.5 V AA batteries are used to supply the robot with the required energy as shown in Fig. 10. Water density is assumed to be 1000 kg/m3. For each experiment, the designed robot swam for approximately 20 s until it reaches a steady-state motion, after that, the steady-state data were recorded and extracted. Fin actuating signals are as shown in Fig. 11. The red curve represents the velocity of the right fin, and it is the same for all of the following cases. The variation of the speed for the left fin for the following four cases is denoted by green, purple, yellow, blue curves, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001429_s106193480803012x-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001429_s106193480803012x-Figure4-1.png", + "caption": "Fig. 4. Cyclic voltammogram of (1) hydroquinone and (2) pyrocatechol in a 0.1 M HCl solution (the scale was 20 mV/cm; cadd = 1 \u00d7 10\u20133 M, v = 10 mV/s).", + "texts": [ + " 63 No. 3 2008 VOLTAMMETRIC DETERMINATION OF DIHYDROXYBENZENES 263 We also tried to determine hydroquinone and pyrocatechol in their binary mixture in a 0.1 M HCl solution by the standard addition method. The CVA of hydroquinone was recorded, while the region of pyrocatechol oxidation potentials was cut off. Each component was determined from a separate aliquot of the test solution in the supporting electrolyte. To improve the parameters of hydroquinone peaks, its CVA was swept from +0.2 to +0.43 V (Fig. 4). This somewhat reduced the effect of the pyrocatechol oxidation\u2013reduction process and improved the cathodic peak of hydroquinone: its height slightly increased. The anodic peak of hydroquinone remained unchanged. As it has already been mentioned, the concentration of pyrocatechol was determined from its cathodic peak, which was recorded from +0.7 V by sweeping potential to the beginning of the reduction of p-benzoquinone. The data obtained using this procedure showed that the results of determining hydroquinone from the anodic peak were unstable, and the tendency to overestimating them by 5\u201320 rel % was observed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001084_detc2007-34070-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001084_detc2007-34070-Figure3-1.png", + "caption": "Figure 3. Definition of the local reference mark", + "texts": [ + "org/about-asme/terms-of-use For this, expressions of normal force F and moment M are linearized relative to generalized displacements dn-s and \u03bbs: = \u2212 s snd kk kk M F \u03bb2221 1211 (6) from the numerical evaluations k11, k12, k21 and k22 defined by: ssnssn M k d M k F k d F k \u03bb\u03bb \u2202 \u2202= \u2202 \u2202= \u2202 \u2202= \u2202 \u2202= \u2212\u2212 22211211 ,,, (7) Meshing stiffness terms k11, k12, k21 and k22 depend on the angular position of the driving wheel \u03981, Under stationary operating conditions, they are periodic functions kij(t) which frequency corresponds to meshing frequency. Stiffness matrix that couples the 6 degrees of freedom of the pinion with the 6 degrees of freedom of the driven wheel is deduced from the calculation of elastic potential energy: >=< \u2212 \u03bb \u03bb n ssnp d kk kk dU 2221 1211, (8) Introducing equations that couple the 12 degrees of freedom (vector q) with generalized displacements dn and \u03bb in the local reference mark (X, Y, Z) defined on figure 3: q T T qT d t t tn .. 2 1 == \u03bb (9) with tq = < u1, v1, w1, \u03c61, \u03c81, \u03b81, u2, v2, w2, \u03c62, \u03c82, \u03b82 > (10) tT1 = [0, cos\u03b2, sin\u03b2, Rb1 tan\u03b1 sin\u03b2, Rb1 sin\u03b2, Rb1 cos\u03b2, 0,-cos\u03b2,-sin\u03b2, Rb2 tan\u03b1 sin\u03b2,-Rb2 sin\u03b2, Rb2 cos\u03b2] (Z parallel to wheel axes, Y along the apparent line of action, X from pinion to the driven wheel) Gear stiffness matrix is written in the following form: T kk kk TK t = 2221 1211 (12) Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/03/2016 T At last, the mean values of meshing stiffness terms kij-m are introduced in equation (9), instead of kij(t), in order to obtain the modal characteristics of the gearbox" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002772_978-3-030-67750-3_14-Figure30.9-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002772_978-3-030-67750-3_14-Figure30.9-1.png", + "caption": "Fig. 30.9 Top view of loader bucket", + "texts": [], + "surrounding_texts": [ + "bucket, which is 100 mm linear actuator and 150 mm linear actuator depending on the length of the arm needed to move. All these actuators are controlled by a motor located at the wheelbarrow body tray and near to the handle where the user can easily manage to operate. The connection harness attached under the wheelbarrow bucket is for safety reason. All the items can be referred in Figs. 30.8 and 30.9 that shows the detail view of the loader bucket." + ] + }, + { + "image_filename": "designv11_83_0003132_978-981-16-0119-4_20-Figure58.6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003132_978-981-16-0119-4_20-Figure58.6-1.png", + "caption": "Fig. 58.6. 3D model made in \u2018Fusion 360\u2019 software ( Source Author)", + "texts": [ + " The prototype was designed and after a practical research it was finalized that the bamboo will rotate as per the design and the laser could move in single plan, i.e. parallel to the ground. The conceptualization started through sketches, but that was not enough as imagining the product and its working was not that easy for the producers (See Fig. 58.5). Prototyping Mentors at Fab-Lab, Nagpur guided students to use \u201cFusion 360\u201d, an integrated software built to bridge designer and producer. The project was first created virtually in the software clearing the details in the design of outer body and the machine (See Fig. 58.6). Frommajor working like access toworking spacemachine, bamboo fixing, etc. To smaller details like display and interaction panels, the 3Dmodel became the reference for prototyping. With use of power tools, the body of machine were made with ply and acrylic. Outer body and compartment for machine was made with plywood, whereas doors to access the working area was designed with plywood and acrylic for easy governess of laser. Power tools were used to cut the plywood and laser to cut acrylic. In the machine, bamboo is fixed and rotated as per design" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001888_2008-01-2630-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001888_2008-01-2630-Figure1-1.png", + "caption": "Fig. 1 Overview of Fiat-131 vehicle gearbox in the first gear-shift", + "texts": [ + " Mathematically, N i i NxxKKurtosis 1 44 /1)( (2) where 4 is the variance square, N is the number of samples, x' is the mean value of samples and xi is an individual sample. A normal distribution has a kurtosis value of 3 and it shows the good condition. Helical gears are almost always used in automotive transmissions. The meshing stiffness of a helical tooth pair is time-varying [10], and was modeled as a series of suggested spur gears so that the simulation techniques for spur gears can be applied. Overview of Fiat-131 vehicle gearbox in the first gear-shift is shown in Fig. 1,where M is Module (mm), b is Face width (mm), is pressure angle (deg), is helix angle (deg) and D1 is pitch diameter (mm). Fig. 2 shows the equivalent gear system in the first gear-shift, where the main parameters for the gear system of Fiat-131 gearbox and the equivalent gear system in the first gear-shift are also shown in the figures. The process of identification the tooth pitting or wear of gear tooth requires building a mathematical model. In most of the gear pair systems, the coupling between the vibration modes is controlled by mesh stiffness therefore a 2-degree-of-freedom semi-definite mathematical model representing only the torsional vibration may yield a quite accurate result in most practical cases" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000999_bf02875875-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000999_bf02875875-Figure1-1.png", + "caption": "Figure 1. Padder roller system action schematic diagram.", + "texts": [ + "20 Keywords: Padder roller system, Mathematical modeling and control, Hamilton principle, Transcental transfer function Introduction In dyeing industry, the process of applying the padder roller system to dye and process the surface of the produced cloth to make the cloth surface luster and uniform is referred to as the cloth surface padding and dyeing. In this process, the fabrics will be immersed into the dye solution for a short period of time. Then the roller of a padder will be employed to squeeze the solution into the lacuna of the tissue of both the fabric and the yarn. Meanwhile, the excess of the solution will be eliminated so as to ensure that the dye solution is evenly distributed over the fabric. In this way, the cloth\u2019s added value can be further increased, as indicated in Figure 1. To facilitate the massive and continuous padding and dyeing treatment to the cloth surface, the padder is utilized and plays a very important role in automatic and continuous production procedure of the padding and dyeing project. However, since the padder has to endure great pressure while in operation, the roller may be deformed, as illustrated in Figure 2. When the roller is deformed under great pressure, and has to continue operation, unbalance vibration will take place and it will lead to the uniformity in the padding and dyeing quality as well as the lower rate of fine products" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003572_s10015-021-00694-y-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003572_s10015-021-00694-y-Figure1-1.png", + "caption": "Fig. 1 Types of in-pipe robot", + "texts": [ + " This research mainly focuses on development the compact in-pipe robot to service the small sewers between 150 and 300\u00a0mm in diameter. These pipes are hard for worker to access. The robot is expected to use to reduce a chance of sewerage system failure and work instead of workers to help them work easier, reduces cost, time, and the risk from flooding causes. There are many researchers have studied about the in-pipe robot in recent decades, each robot was proposed in different mechanism and different purposes. The in-pipe robot can be classified to 7 types, as shown in Fig.\u00a01. Each mechanism has its advantage and disadvantage, depends on purpose and environment. The PIG type is the only one type which is driven by fluid pressure. Mostly it is applied in an industrial to service the long pipeline, such as gas, oil, or water pipeline. The wheel type is usually used for inspection and cleaning in smooth surface, such as gas or air pipe [1\u20134]. The crawler type is used to achieve the rough surface, Nagaya K at el had developed the inspection robot to climb on a vertical steel pipe using the advantage This work was presented in part at the 26th International Symposium on Artificial Life and Robotics (Online, January 21- 23, 2021)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002191_j.matpr.2020.12.112-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002191_j.matpr.2020.12.112-Figure2-1.png", + "caption": "Fig. 2. 3D modelling.", + "texts": [ + " This point bargains about the compelled part appearing of the posts and imitating them by contributing specific physical representations and limit situations to reenact the fascinating situation plans ended likely. The segment stayed demonstrated utilizing ANSYS 14.0 software Plan Modeler programming as a strong perfect and ANSYS 14.0 worktable was utilized aimed at the assess- ment in the static examination. The key work method of a Limited Component Examination system is tended to like in Fig. 1(A) Fig. 2 Fig. 3 Fig. 4 Fig. 5 Fig. 6 Fig. 7 Fig. 8 Fig. 9 Fig. 10. The stainless harden fabric property Bulk (q): 7950 kg/m3 Young\u2019s modulus (E):206000Mpa Poisson\u2019s ratio: 0.3 The CAD typical of the bar is fit hooked on a limited quantity of components utilizing ANSYS 14.0 software inherent lattice calculation. The pillar contact locale is fit and interlinked to empower estimation of the power collaboration between them limited component investigation or FEA representing to a genuine task as a \u2018\u2018work\u201d a progression of little, consistently formed tetrahedron associated components, as appeared in the above fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003080_s40430-021-02961-2-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003080_s40430-021-02961-2-Figure1-1.png", + "caption": "Fig. 1 Quarter car model for braking and suspension dynamics", + "texts": [ + " The measured vertical axle acceleration and braking acceleration are used to estimate dynamic wheel load with artificial neural network (ANN). For this aim, ANN was designed by using multilayer perceptron (MLP) networks. Also, the measured axle height was employed to determine the direction of damper motions. 2 The modeling of\u00a0the\u00a0integration between\u00a0suspension and\u00a0braking dynamics The quarter car model is used to investigate the effects of wheel load oscillations on brake pressure changes of ABS as shown in Fig.\u00a01. This model composes braking and suspension systems. In the model, the unsprung and sprung masses are reduced into two masses. Both of these masses are connected to each other by means of the shock absorber and spring as shown in Fig.\u00a01. Therefore, the effects of suspension dynamics are reflected on braking dynamics through the spring and damper. In this study, the variations in damper level are only considered to reflect effect of vertical dynamics of suspension into braking dynamics. In this study, suspension dynamics are integrated to braking dynamics through vertical wheel load, since the vertical load acting on wheel is the common parameter for both suspension and brake dynamics. Thus, the changes in wheel loads connect suspension dynamics to braking dynamics", + "\u00a0(2a) as follows: The wheel load changes are calculated from Eq.\u00a0(3) as follows: where \u0394Pb(t) is the brake pressure change rate, \u0394?\u0308?(t) is the change in wheel acceleration, and \u0394Fz(t) is the change in wheel load. (1)MV\u0307x = \u2212\ud835\udf07(\ud835\udf06)Fz(t) (2)Iw?\u0308?(t) = \ud835\udf07(\ud835\udf06)ReFz(t) \u2212 \ud835\udf07bAbrbPb(t) (2a)Iw?\u0308?(t) = \ud835\udf07(\ud835\udf06)ReFz(t) \u2212 KbPb(t) (3)Fz(t) = Iw?\u0308?(t) + KbPb(t) \ud835\udf07(\ud835\udf06)Re (4)\u0394Fz(t) = Iw(\u0394?\u0308?(t)) + Kb(\u0394Pb(t)) \ud835\udf07(\ud835\udf06)Re The suspension system between sprung mass and unsprung mass is modeled with two degree of freedoms as shown in Fig.\u00a01. Therefore, the motion equations of suspension model can be written as follows: where k2 and k1 are the spring coefficients of the suspension and tire, c2 and c1 are the damping coefficients, and g is the acceleration of gravity. In this study, the damping coefficient of the tire c1 is neglected, because the damping capacity of the rubber is very low for small size tires used in passenger car. The right two terms of Eq.\u00a0(6) describe the dynamic and static wheel loads, respectively. The dynamic and static wheel loads are described as follows: When Eq", + "\u00a0(6) becomes as follows: where FkB is the body spring force, FcB(vd) is the damper force produced by active damper, and it is function of the piston velocity of damper vd. Thereby, the damping force is described as follows: Hence, the wheel load change is specified as follows: where Fz is the actual wheel load, Fz,st is the static wheel load, and Fz,dyn is dynamic wheel load. In here, static wheel load has constant value and the actual wheel load takes only positive values due to downward vertical load acting on front axle, while the dynamic wheel load can take positive and negative values due to the oscillations between tire and rim as shown in Fig.\u00a01. So that, the dynamic wheel load determines the direction of wheel load changes as the following: Therefore, the wheel load changes are expressed by substituting Eq.\u00a0(11) into Eq.\u00a0(10) as follows: (5)m2z\u03082 + c2(z\u03072 \u2212 z\u03071) + k2(z2 \u2212 z1) = m2g (6)m1z\u03081 \u2212 c2(z\u03072 \u2212 z\u03071) \u2212 k2(z2 \u2212 z1) = \u2212k1(z1 \u2212 z0) + m1g (7)Fz,st(t) = m1g (8)Fz,dyn(t) = \u2212k1(z1(t) \u2212 z0(t)) (9)m1z\u03081 = (FcB(vd) + FkB) + (Fz,dyn + Fz,st) (10)FcB(vd) + FkB = m1z\u03081 \u2212 ( Fz,st + Fz,dyn ) (11)\u0394Fz(t) = m1z\u03081 \u2212 ( Fz,st + Fz,dyn(t) ) (12) If Fz,dyn > 0, \u0394Fz < 0 If Fz,dyn < 0, \u0394Fz > 0 Journal of the Brazilian Society of Mechanical Sciences and Engineering (2021) 43:367 Page 5 of 21 367 In Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002927_j.ijleo.2021.167285-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002927_j.ijleo.2021.167285-Figure3-1.png", + "caption": "Fig. 3. Temperature compensated tip-force sensing epidural needle.", + "texts": [ + " Two Fabry-Perot interference (FPI) sensors are employed for force sensing function of the needle. To produce a qualified FPI sensor, the procedure is separated into six steps as shown in Fig. 2. A cavity (named FP cavity) is formed by two ends of a cleaved fiber, which is cut by a fiber cleaving instrument and fixed inside a glass capillary. One fiber with a connector was cut and cleaved into two parts and fixed with epoxy into a glass capillary of 250 \u00b5m in outer diameter to form a FP cavity with a cavity length of around 20 \u00b5m. Fig. 3 shows the sensing needle structural design, where two FPI sensors are placed at tip of the needle. As the FPI sensor can be severely affected by temperature variation, temperature compensation is required in vivo applications. The one adhering to the needle Z. Mo et al. Optik 242 (2021) 167285 wall serves as force sensor, which could sense the hybrid signal change of force and temperature. Another one, served as reference FPI sensor, was placed close to needle tip, hung as a cantilever inside of the needle and mounted on a metal support beam, providing temperature change information to the force sensor while avoid force influence applied on the needle tip" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001586_iembs.2008.4649475-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001586_iembs.2008.4649475-Figure1-1.png", + "caption": "Fig. 1. Pro/ENGINEER Wildfire depiction of the prototype mouse shell.", + "texts": [ + " The inclined surface was intended to slightly supinate the hand, yielding a more stable orientation of the user\u2019s forearm bones. The design offered two primary challenges: the creation of a working tracking system and new mouse button functionality. After several failed attempts, the team designed the hardware as a shell around a commercial mouse, thereby eliminating the need to redesign the optical tracking and push-button elements. Pro/ENGINEER Wildfire 2.0 was used to create computer models of the mouse shell (see Fig 1). Given these three-dimensional computer-aided-design models, a Stratasys Dimension BST (Breakaway Support Technology) rapid prototyping machine was used to create proof-of-concept hardware (see Fig. 2). Note that software filters to remove shaking from the mouse coordinates sent to the operating system were also considered. These may be addressed in future efforts. After meeting with the arthritis support group, it was clear that battery removal, whether from a TV remote control or a CD player, was difficult for these individuals" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003135_012006-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003135_012006-Figure1-1.png", + "caption": "Figure 1. Structure of human palm joints Figure 2. Bionic hand", + "texts": [ + " In this paper, the artificial fish swarm algorithm (AFSA) [8] is used to solve the inverse kinematics problem of the bionic hand with the potential of approaching the global minimum. Also, the optimal real solution of the angle is obtained by simulating the behavior of the fish swarm and searching the optimal angle based on the expression of the given position, and how the parameters of the AFSA affect its solution accuracy is analyzed. The bones of the human hand consist of carpal bones, metacarpal bones, and phalanges, as shown in Figure 1. Five metacarpal bones form the palm, and each finger consists of three phalanges, that is, metacarpophalangeal joint, proximal interphalangeal joint, and distal interphalangeal joint. In addition, the thumb consists of a metacarpal joint and two phalangeal joints. Based on the human hand structure, this paper proposes a bionic robotic arm consisting of rigid links and rotational joints, as shown in Figure 2. From Figure 2, it can be seen that the robotic hand consists of a finger prosthesis, a palm, fishing lines, spring pieces, and a motor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002295_j.robot.2020.103715-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002295_j.robot.2020.103715-Figure3-1.png", + "caption": "Fig. 3. (a) shows an image of a single segment. It consists of a main body that contains the servo drive and the sensor board. The lateral extensions increase the stability of the robot and span the distance between the main bodies of consecutive segments within the robot chain. (b) shows the geometric representation of a segment within the simulation. Corresponding to the sensorization of the real segment, those areas used for tactile feedback are marked in yellow. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)", + "texts": [ + " This design approach has already been realized by Chiu et al. [10]. In the following, first the mechanical setup of the robot, its sensorization, and the representation within a physics simulation will be described in Section 2. Section 3 covers the control concept that allows the robot to adapt to its environment. Finally, in Section 4, simulation data is compared to experiments using the physical robot. In its current state, the robot consists of 12 segments, which are connected as closed kinematic chain (see Fig. 1). In Fig. 3 (a), a single segment of this chain is depicted. Each segment is equipped with a servo motor which is used to adjust the angle to the next segment. The outer surfaces of the segments are covered with contact sensors based on the piezoresistive principle [11, see Fig. 4]. In addition to the contact sensors, four segments are equipped with acceleration sensors to determine the orientation of the robot.2 This data is used to maintain stability during locomotion. In parallel to the development of the hardware, a simulation of the robot was created based on ROS (Robot Operating System; ros", + "org). This simulation was used to test and evaluate the controller before its application to the real robot. As mentioned, the robot consists of 12 nearly identical segments (four of the segments are equipped with acceleration sensors in addition to the contact sensors) that are linked as closed kinematic chain. Each of these segments contains a servo drive (Dynamixel AX-12A; ROBOTIS Co. Ltd., South Korea) whose mechanical output is connected to the next segment of the chain. As can be seen in Fig. 3 (a), the servo is located in the central compartment of the segment. The two lateral extensions, which are shifted relative to the central part, have been added to expand the sensorized area. Also, they increase the stability of the system against sideward tilting. The central compartment as well as the lateral extensions were printed in ABS using a 3D Printer (Replicator 2; MakerBot Industries LLC, USA). The piezoresistive contact sensors [11, see Fig. 4 (a, b)] consist of comblike electrodes covered by conductive foam (4450", + " Residues of the resist were removed using acetone-soaked tissues. The sensors were cut out of the laminate sheet and to prevent oxidation of the copper, they were chemically tinned. Both, sensor flex-PCB and ESD-foam were attached to the segments by screws. Thus, the foam is pressed against the flex-PCB at the location of the screws. Since this contact would be detected as collision, these areas were isolated using adhesive Kapton R\u20dd tape. As mentioned before, the segment hull consists of the central ompartment and two lateral extensions (see Fig. 3). For each of hese three sections individual sensor flex-PCBs were manufacured. The resistance of the electrode-foam contacts is evaluated ocally by an A/D-converter (LTC2309; Linear Technology Corpoation, USA) circuit located at the inner side of each segment. As epicted in Fig. 6, the A/D-converters and acceleration sensors ADXL312; Analog Devices Inc., USA) are connected via I2C to a icrocontroller board (Teensy 3.2; PJRC.COM LLC, USA) in a ring opology. The servos are controlled via a TTL-serial interface that as added to the microcontroller board" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003065_s12206-021-0607-z-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003065_s12206-021-0607-z-Figure4-1.png", + "caption": "Fig. 4. Assembly model of the pinion and the gear.", + "texts": [ + " Substituting Eq. (10) into Eq. (11), the meshing equation is represented by: ( , ) ( , ) ( , ) ( , ) ( , ) ( , ) ( , ) ( , ) ( , ) ( , ) ( , ) ( , ) ( , ) ( , ) ( , , ) ( ) ( ) ( ) 0 c c c c c c c c c c c c i j i j i j i j i j i j i j c j k k k k k k lk i j i j i j i j i j i j i j k k k j k k k k j f k z x x z x y z z y y r z x x z \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03d5 \u03d5 = \u2212 \u2212 \u2212 + \u2212 = . (12) From Eqs. (10) and (12) and Table 1, the assembly model of the pinion and the gear with S-shape surfaces is shown in Fig. 4. This section evaluated the kinematic errors based on the assembly errors of the S-shaped surface gear pairs with asymmetric involute teeth. As shown in Fig. 5, ( , , )f f f fS x y z was a fixed coordinate system. The coordinate systems ( , , )h h h hS x y z , ( , , )a a a aS x y z and ( , , )b b b bS x y z were auxiliary coordinate systems used for describing the assembly errors. The symbols rx\u0394 and ry\u0394 were used to set the horizontal and vertical axes of the two gears. The error of the center distance between points gO and fO is the position vector c\u0394 , and c\u0394 = x y zd d d+ +i j k " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002993_s00170-021-07405-8-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002993_s00170-021-07405-8-Figure1-1.png", + "caption": "Fig. 1 Geometry of sun gear", + "texts": [ + " Thus, the objectives of this study are (i) to propose the precision sizing process as a postoperation to enhance the forming quality of sun gears formed by cold extrusion process, and (ii) to investigate the influences of important process parameters such as the sizing amount of external gear, interference value of internal spline, die bearing length, and friction factor on dimensional accuracy of internalexternal teeth to determine the optimal parameters for commercial production of sun gears and future research. Sun gear is a kind of typical gear-like part, which is widely used in the truck transmission system. The geometry and main parameters of target sun gear in this study are shown in Fig. 1 and Table 1, respectively. It is observed that the sun gear has two kinds of tooth shapes, namely external gear and internal spline, respectively. Figure 2 shows the proposed cold extrusion process of sun gear in previous investigation. First, a ring-shaped billet 1, formed by hot-forged, was inserted into the stationary splined mandrel and die. Then, billet 1 moved downward by the punch, and the internal tooth shapes were gradually formed by the splined mandrel. When billet 1 was going to pass through the die cavity completely, billet 2 was placed on billet 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002128_s13369-020-05100-6-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002128_s13369-020-05100-6-Figure6-1.png", + "caption": "Fig. 6 Block diagram of the SMD system, PC interface program, control units and the CADWD machine", + "texts": [ + " Afterward, a translational transform between points is created (X\u2013Y path), the layers are built by increase Z axes by layer height for the X\u2013Y path. The solid model is expressed as points (x, y, z). The machine is represented by 3 axes Cartesian robot for simulation and visualization, from solid model points, the Cartesian trajectory is generated (x, y, z) and (x\u0307, y\u0307, z\u0307). That means the points of motion and velocities are specified, then the machining process is animated in a graphical procedure. The pattern points can be saved and can be loaded on the machine user interface program for experiment implementation. Figure 6 represents the computer-aided design (CAD) program. The CADWDM control system contains an innovative program created through the MATLAB program. MATLAB is a high-level technical computing language, and it is an interactive environment for algorithm development, data visualization. The MATLAB is used because of the system requirement to work at the online process that is able by MATLAB real-time signal processing. The program allows us to draw a CADmodel for the part inMATLAB rather than utilizing commercial CAD packages", + " Simultaneously, the current program divides the part that was drawn by the CAD program directly from the base layer toward the top based on the thickness of the painted layer and the number of layers. The currentCADWDMsystemused the online programing method via three-dimensional CAD data for components to create and simulate CADWDM software. Online programing has been used because it is available at a low cost compared to off-line programing(OLP)software packages, especially for small size components. The system constructed by the PC interface program, control unit, and CADWD machine. The block diagram for the system is shown in Fig. 6. The PC interface program is a user interface panel, it is designedwith the graphical user interface development environment (GUIDE) of MATLAB\u00ae. The PC interface program is a computer-aidedmanufacturing (CAM) program as shown in Fig. 7, it\u2019s used for controllingmachine torchmovements relative tomachineworkspace, the program has received the CAD output file and sent it to low-level machine program (the brain of machine C++ code), so CAM program reads and follows CAD program and change it to commands that will send to the microcomputer (Arduino) inside the machine (which has low-level machine code) and is exchanged the step of the process with the control unit to satisfy the required transportation for machine torch" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002772_978-3-030-67750-3_14-Figure28.8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002772_978-3-030-67750-3_14-Figure28.8-1.png", + "caption": "Fig. 28.8 Outlet 2\u2014velocity reading of Mode 2 nozzle", + "texts": [ + "9 presents the pressure curve that is affected by the length of the nozzle. It clearly can be seen that the pressure at outlet 2 is inversely proportional to the nozzle length. As the nozzle length increases, the pressure will be decreased. The graph pattern is in a gradually decreasing manner. This is because, as the powder goes through along the nozzle, the pressure will continue to decrease due to the loss of potential energy. Figure 28.7 shows the pressure reading during Mode 2 nozzle was set. While, the velocity reading of Mode 2 nozzle set as illustrated in Fig. 28.8. Presented in Fig. 28.10 is the velocity of the outlet 2 against the nozzle length. It shows that as the length of the nozzle increases the velocity happening on outlet two also increases. This is because the velocity is directly proportional to the length of the nozzle. Based on this graph, it is concluded that the higher the length and the diameter of the outlet, the faster the velocity. The Taguchi method was conducted using the Minitab software (Khavekar et al. 2017; Arnold 2006; Zaman et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002772_978-3-030-67750-3_14-Figure17.3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002772_978-3-030-67750-3_14-Figure17.3-1.png", + "caption": "Fig. 17.3 Rolling resistance", + "texts": [ + " The tractive force must fulfill the requirements of vehicle dynamics to overcome forces such as the rolling resistance, gravitational, and aerodynamically which are summed together as the road load force FRL as shown in Fig. 17.2 (Arof et al. 2020a, b, c, d). Thegravitational force Fg dependson the slopeof the roadway, as shown inEq. (17.1). Fg = mg sin \u03b1 (17.1) \u03b1 is the grade angle, m is the total mass of the vehicle, g is the gravity constant. The hysteresis of the tire material causes it at the contact surfaces with the roadway. The centroid of the vertical forces on the wheel moves forward when the tire rolls. Therefore, from beneath the axle toward the direction of motion by the vehicle, as shown in Fig. 17.3. Tractive Force The tractive forcewas used to overcome the Froll force alongwith the gravity force and the aerodynamic drag force. The rolling resistance has been minimized by keeping the tires as inflated as possible by reducing the hysteresis. The ratio of retarding forces due to rolling resistance and the vertical load on the wheel known as the coefficient of rolling resistance C0. The rolling resistance force is given by Eroll = \u23a7 \u23a8 \u23a9 sgn[vxt ]mg(C0 + C1v2xT ) if vxt = 0 FTR \u2212 FgxT if Vxt = 0 and \u2223 \u2223FTR \u2212 FgxT \u2223 \u2223 \u2264 C0mg sgn(FTR \u2212 FgxT )(C0mg) if Vxt = 0 and \u2223 \u2223FTR \u2212 FgxT \u2223 \u2223 > C0mg \u23ab \u23ac \u23ad (17" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002714_tmag.2021.3076084-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002714_tmag.2021.3076084-Figure5-1.png", + "caption": "Fig. 5. Measuring system.", + "texts": [ + " _ 6 6 _ _ 6 6 _ _ 6 6 _ _ 6 6 _ _ 6 6 _ _ 6 6 _ _ 6 6 _ _ 6 6 _ 1 1 6 6 5 5 6 6 5 5 6 6 1 1 6 6 0 cos sin cos sin 1 03 cos sin cos sin A d p p r q p p r B d p p r q p p r C D E d p p r q p p r F d p p r q p p r I i i I i iI I I i i I i i _12 6 _ _12 6 _ _12 6 _ _12 6 _ _12 6 _ _12 6 _ _12 6 _ _12 6 _ _12 6 _ _12 6 _ 1 1 6 6 5 5 6 6 5 5 6 6 1 1 6 6 0 cos sin 3 cos 3 sin 1 2 cos 2 sin3 3 cos 3 sin cos sin d p p r q p p r d p p r q p p r d p p r q p p r d p p r q p p r d p p r q p p r i i i i i i i i i i (14) The amplitude and phase difference of (14) are the same as (12). IV. VERIFICATION OF CHARACTERISTICS In order to verify the effectiveness of the proposed faulttolerant control method for a two-controllable-rotor motor, the torque ripples and load characteristics are computed and measured using FEA and a prototype, respectively, where the measuring system is shown in Fig. 5. A. Back-EMF The computed and measured back-EMFs are compared, where one rotor is rotated at 60 rpm and the other rotor is fixed. Fig. 6 shows the measured back-EMF waveform, and Fig. 7 shows the components of the computed and measured backEMFs. From Fig. 7, the measured fundamental component is smaller than the computed fundamental component, and the measured harmonic components are larger than the computed harmonic components. These are due to the assembly and dimension errors of the prototype" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000474_bfb0031450-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000474_bfb0031450-Figure3-1.png", + "caption": "Fig. 3 : Slider-crank mechanism", + "texts": [ + " To determine the rotational component of the phalanx speed vector, an additional condition has to be formulated. One possible way is stating that the rotational speed of the phalanx equals the rotational speed of the object. This corresponds to the previously explained theory, where specific object movements are generated by identical movements of the contact plane (3 D) or contact line (2 D). The proposed method states that the contact line movement is a pure translation. The required phalanx rotation in frame Cld can be obtained when considering the system as a slider-crank mechanism (figure 3). The slider is moving on the contact line and the crank is formed by the line connecting the contact point 2,i with the centre of curvature p corresponding to the contact 1 j. The rod connecting P to the contact point 1,j is (instantaneously) fixed to the index. Figure 4 shows the mathematical equivalent, in this case the object has a zero curvature at point 2,i. When this is not the case, an additional (contact line) velocity component has to be taken into account. The length of the driving rod is n" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000024_2005-01-3371-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000024_2005-01-3371-Figure3-1.png", + "caption": "Fig. 3 Normal radial stress distribution over the gear ; (a) radial stress (b) tangential stress", + "texts": [ + " This computed solution, referred to as the true solution, is obtained for radial, tangential and shear stresses all over the gear. With this computed solution, the locations of the radial and tangential stress extrema are known, as shown in Fig. 2(a-b). We thus train our machine learning optimization algorithm on the stress and position data around these locations as well as define an optimal sensor layout over the gear for damage detection. 2 U.S. patent application filed Using this strategy, we train the algorithm on about 10% of these data. A prediction of normal radial stress all over the gear is shown in Fig. 3.(a). Comparing Figs. 2(a) and 3(a), it is shown that the prediction follows the true solution closely. The rms error over the entire grid is about 5% . Fig. 3(b) shows the predicted normal tangential stress distribution. Again, comparison of Fig. 3(b) with Fig. 2(b) shows that the predicted and true tangential stresses are in close agreement. Extending the learning models thus constructed to test damage data, with one gear tooth made less rigid than the rest of the teeth, predictions of radial stress are made by the thus learned model. These predictions are shown in Fig. 4(a) as the difference between the true normal radial stress and the just predicted damage radial stress distributions. Inspection of Fig. 4(a) clearly shows a discrepancy in the radial stress distribution in a given region, and this local region turns out to be correspondent to be the damaged tooth" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001616_cit.2008.4594755-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001616_cit.2008.4594755-Figure1-1.png", + "caption": "Figure 1. The devices of swarm-bots required for implementing exploring robots", + "texts": [ + " Table 1 shows additional packages that are required for using Pyro. JPEG Runtime in Table 1 is used for capturing the screenshot of the simulator or transmitting an image to the camera on a robot system. OpenCV is adopted to process the image from the camera on a robot. TCL/TK imaging and Python imaging packages is used for image processing of the user interface module. In our experiment, our test Pyro robot module was assumed to be equipped with two sonar sensors. Also, we added two virtual GPS module and the virtual motor wheel system in the simulation. Fig. 1 shows the devices required for implementing exploring swarm-bots. One camera sensor is used for identifying other robots and targets. And a number of sonar sensors are used to detect the surrounding barriers. The distance information to the barrier is used in the map construction algorithm embedded in the robot. The detailed algorithm will be explained in Section 5. A GPS device is used for searching a destination and a target location using grid fitting. Finally, motor device is responsible for moving a robot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003139_aced50605.2021.9462301-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003139_aced50605.2021.9462301-Figure4-1.png", + "caption": "Fig. 4. Quarter Finite Element (FE) model of the WRSESM.", + "texts": [ + " In order to investigate the performance of the DFC, a Wound Round Separated Excited Synchronous Machine (WRSESM) was modelled in ANSYS MAXWELL as a Finite Element design. The WRSESM taken as basis for the modelling is manufactured by WUEKRO GmbH. The mentioned company supplied the real WRSESM alongside with detailed information of the machine e.g. machine geometry drawings, magnetic materials and windings scheme among many other data. The general characteristics of the WRSEM are stated in Table 1. The WRSESM has a four pole symmetric structure and therefore we have created in ANSYS Maxwell just one quarter of the machine (see Fig. 4). By modelling just one period of the periodic structure, the size of the FE design is reduced, thus the calculation complexity is diminished and the FE solution time is shorted without losing the high fidelity brought by the FEM simulation [12]- [14]. Next, the validation process for the FE model was carried out. Three different tests were performed on the software model and on the real machine and a relative error was calculated for each measurement. The first test was the open circuit test that shows a maximum error of 7% between the FE model and the real machine for field winding currents larger than 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001804_robot.2007.364178-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001804_robot.2007.364178-Figure1-1.png", + "caption": "Fig. 1. 2-1 Euler Angle Rotation", + "texts": [ + " Using the approach of Robinett et al [12], Starr et al [3] chose objective function matrices to enforce a minimum energy condition for the optimization of movements by a gantry crane system with one degree of freedom. We have applied this same methodology to a two degree of freedom extension of the system for investigating curvilinear motions. The system used for our investigation is a simple pendulum modeled as a two degree of freedom system with a 2-1 Euler angle rotation sequence, depicted in Fig. 1. While this rotation sequence does have singularities, they occur when \u03c6 = \u00b1\u03c0 radians, far beyond the range of motion seen in our simulations or experiments. Equations of motion for the 1-4244-0602-1/07/$20.00 \u00a92007 IEEE. 4537 system are as follows: \u03b8\u0308 = 2\u03b8\u0307\u03c6\u0307 tan\u03c6\u2212 g l sin \u03b8 cos\u03c6 + 1 l x\u0308 cos \u03b8 cos\u03c6 (1) \u03c6\u0308 = \u2212\u03b8\u03072 cos\u03c6 sin\u03c6\u2212 g l cos \u03b8 sin\u03c6 (2) \u2212 1 l (x\u0308 sin \u03b8 sin\u03c6+ y\u0308 cos\u03c6). Initially, we considered elliptical spatial trajectories for this work. However, an angle of inclination and exactly four points must be used to define elliptical segments, and only a small fraction of possible point sets can be used to define an ellipse" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001057_20080706-5-kr-1001.01557-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001057_20080706-5-kr-1001.01557-Figure2-1.png", + "caption": "Fig. 2. A regular triangular array of optical mice with N=3.", + "texts": [ + " 1c), three velocity vectors become different in direction but the same in magnitude. Theses observations tells that a different traveling pattern of a mobile robot results in a set of different velocity readings of an array of optical mice. Reversely, it is possible to estimate the linear and angular velocities of a traveling mobile robot from the velocity readings of optical mice. Assume that N optical mice are installed at the vertices, P i, i= 1,\u2026,N , of a regular polygon that is centered at the center, O b , of a mobile robot traveling on the xy plane. Fig. 2 shows an example of a regular triangular array of optical mice with N= 3 . Let u x= [ 1 0 ] t and u y= [ 0 1 ] t be the unit vectors along the x axis and the y axis, respectively. The position vector, p i = [ p ix p iy ] t , i= 1,\u2026,N , from O b to P i, can be expressed as p i = [ ]p ixp iy = \ua00e \ua01a \ufe33 \ufe33 \ufe33\ufe33 \ua00f \ua01b \ufe33 \ufe33 \ufe33\ufe33 r cos {\u03b8+ ( i-1)\u00d7 2\u03c0 N } r sin {\u03b8+ ( i-1)\u00d7 2\u03c0 N } (1) where \u03b8 represents the heading angle of a mobile robot with the forwarding direction aligned with p 1 , and r represents the distal distance of each optical mouse", + " Let v b= [ v bx v by ] t and w b be the linear velocity and the angular velocity at the center O b of a mobile robot, respectively. And, let v i= [ v ix v iy ] t , i= 1,\u2026,N , be the linear velocity at the vertex P i, which corresponds to the velocity readings of the i th optical mouse. Then, there holds the following velocity relationship: v b + \u03c9 b q i = v i (3) Premultiplied by u x t and u y t, (3) gives u x t v b + \u03c9 b u x t q i = u x t v i (4) u y t v b + \u03c9 b u y t q i = u y t v i (5) respectively. Referring to Fig. 2, (4) and (5) can be rewritten as v bx - wb\u00d7p iy = v ix (6) v by + wb\u00d7p ix = v iy (7) From (6) and (7), the velocity mapping from a mobile robot to an array of optical mice can be represented as A x \u0307 = \u0398 \u0307 (8) where x\u0307 = \ua00e \ua01a \ufe33 \ufe33 \ua00f \ua01b \ufe33 \ufe33 vbx vby \u03c9 b (9) \u0398 \u0307 = \ua00e \ua01a \ufe33 \ufe33 \ufe33\ufe33 \ufe33 \ufe33 \ua00f \ua01b \ufe33 \ufe33 \ufe33\ufe33 \ufe33 \ufe33 \u0398 \u0307 1 \u0398 \u0307 2 \u22ef \u0398 \u0307N \u2208 R 2N\u00d71 with \u0398 \u0307 i = [ ]v ix v iy (10) A = \ua00e \ua01a \ufe33 \ufe33 \ufe33\ufe33 \ufe33 \ua00f \ua01b \ufe33 \ufe33 \ufe33\ufe33 \ufe33 A1 A2 \u22ef AN \u2208 R 2N\u00d73 (11) with A i = [ ] 1 0 -p iy 0 1 p ix (12) Note that the expression of A is quite simple as a function of the position vectors, p i=[ p ix p iy ] t, i=1,\u2026,N" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001429_s106193480803012x-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001429_s106193480803012x-Figure3-1.png", + "caption": "Fig. 3. Cyclic voltammogram of (1) hydroquinone and (2) pyrocatechol in the boric acid supporting electrolyte with pH 8.3 (cadd = 1 \u00d7 10\u20133 M, v = 10 mV/s).", + "texts": [ + " The studies of the effect of the pH of boric acid solution in the range from 4.9 to 8.9 (the electric conductivity of the solution was increased by the addition of a 0.1 M Na2SO4 solution) on the shapes of hydroquinone and pyrocatechol peaks revealed the following. At pH 4.9, anodic peaks of both components retained their sharp shapes, whereas cathodic peaks had wide plateaus instead of peaks. At pH 7.2, both anodic peaks decreased in heights and widened in bases, while cathodic peaks became sharp. In the pH range 7.7\u20138.9, CVAs of both components had the same shape (Fig. 3). Both peaks of 262 JOURNAL OF ANALYTICAL CHEMISTRY Vol. 63 No. 3 2008 SKVORTSOVA et al. hydroquinone were sharp, the heights of pyrocatechol anodic and cathodic peaks almost halved as compared to their heights in a 0.1 M HCl solution, and only the cathodic peak retained its sharp form. Thus, in the borate buffer solution with pH > 8, we managed to significantly decrease the currents of pyrocatechol anodic and cathodic peaks and to retain the parameters of hydroquinone peaks. The currents of both hydroquinone peaks at different stationary concentrations of pyrocatechol in the borate buffer solution with pH 8", + " The concentration of hydroquinone was found by measuring the currents of anodic and cathodic peaks in the boric acid supporting electrolyte with pH 8.3. The concentration of pyrocatechol was found from its reduction peak in a 0.1 M HCl solution. We managed to reduce the effect of pyrocatechol on the CVA of hydroquinone to some extent by the cessation of sweeping potential to the anodic region before the beginning of pyrocatechol oxidation and by changing the direction of the sweep, as is shown in Fig. 3, CVA 1'. The results of determining the components were verified in model solutions by the standard addition method and from the calibration graphs. The results of hydroquinone determination by the standard addition method from the measurements of both peaks in the solution that contained the equal amount of pyrocatechol (0.400 g/L) were overestimated by 10\u201350 rel %. Such a systematic error was evidently due to the unproportional growth of oxidation and reduction peaks of hydroquinone when its standard addition was added to the solution" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001866_j.tcs.2008.04.001-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001866_j.tcs.2008.04.001-Figure1-1.png", + "caption": "Fig. 1. A membrane structure.", + "texts": [ + " There were investigatedmany classes of P systems and most of them are computationally complete; when an exponential workspace can be created in a polynomial time, e.g., by membrane division, then polynomial (often, linear) solutions to computationally hard problems (typically, NP-complete problems) can be devised.\" This paragraph can be found almost in this form in the introduction of many papers in the membrane computing area; very frequent was and still is the illustration of the notion of amembrane structure (and of the associated terminology) from Fig. 1, a sort of logo of the domain. To a great extent, these phrases and this figure capture the essence of membrane computing. It should be added another slogan of the field, stating that \u201cthe rules are used in the non-deterministic maximally parallel way\", and that recently membrane computing proved to be a very promising framework for devising models for biology (and other areas, such as economics and linguistics) \u2013 with surprising applications in unexpected areas, such as approximate optimization, in the sense of evolutionary computing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001108_j.ijsolstr.2008.03.010-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001108_j.ijsolstr.2008.03.010-Figure1-1.png", + "caption": "Fig. 1. (a) Honeycomb constitutive law, (b) Al foam constitutive law (c) discrete-chain, (d) contact surface (footprint) (e) 2D strip with corner impedance and mobility.", + "texts": [ + " Finally, a 2D strip model is developed in Section 4 that includes intermediate flexible supports to simulate thin electronic boards carrying lumped masses and held by a stiff frame mounted on the crushable material. The strip responds to decelerations from the collapsing crushable material, transmitted to the strip boundaries and intermediate supports by the holding frame. The goal is to evaluate how transmitted deceleration changes from strip inertia, boundary conditions and intermediate supports discussed in Section 5.4. The non-linear constitutive law of the material is made of three stages (see Fig. 1(a)\u2013(b)). The first stage is linear elastic till critical or collapse stress rcr is reached. The second stage lies over a relatively wide slant or flat plateau during collapse of the crushable material. The plateau average line happens over a constant stress for honeycomb (Fig. 1(a)), or follows a strainhardening curve for metal foams (Fig. 1(b)). Once most cells have collapsed, a stage of densification starts and the material reverts to its bulk properties. Let rf(ef) be the constitutive law of the crushable material where ef is engineering strain. Consider a 1D layer of crushable material with thickness hf and relative density qf < qs where qs is density of the parent material. Divide hf into nf cells where each cell is hc = hf/nf long with mass mi = mc = hcqf. The impact of a rigid mass M attached to the chain of masses mi and springs ki (see Fig. 1(c)), striking a rigid boundary with an initial velocity Vo, starts the process of collapse. The first mass in the chain is M. The coordinates zoi and zi measure initial and instantaneous distance between mi and impact surface. Let ui be the displacement of mi and ei the strain in spring ki connecting mi to mi+1 D~u1 \u00bc \u00f0u1 u2\u00de=\u00f0hc=2\u00de; D~unf\u00fe1 \u00bc \u00f0unf\u00fe1 ub\u00de=\u00f0hc=2\u00de D~ui \u00bc \u00f0ui ui\u00fe1\u00de=hc; i \u00bc 2; ::;nf z1 \u00bc zo1 u1; zi \u00bc zoi ui; i \u00bc 2;3; ::;nf \u00fe 1 zo1 \u00bc hf ; zoi \u00bc hf \u00f0\u00f02i 3\u00de=2\u00dehc \u00f01\u00de At the rigid boundary, the displacement ub vanishes. In (1), note that the undeformed length of the first and last cells in the chain is hc/2 (see Fig. 1(c)). The dynamic equations are Mottu1 \u00fe rf \u00f0D~u1\u00de \u00fe 1tot\u00f0u1 u2\u00de \u00bc 0 miottui \u00fe rf \u00f0D~ui\u00de rf \u00f0D~ui 1\u00de \u00fe 1tot\u00f02ui ui\u00fe1 ui 1\u00de \u00bc 0; i \u00bc 2;3; ::; nf \u00fe 1 \u00f02\u00de with initial conditions otui\u00f00\u00de \u00bc Vo; i \u00bc 1;2; . . . ;nf \u00fe 1 \u00f03\u00de Bt is a constant viscous damping impedance. Reverse motion is not allowed meaning that for the ith cell ui\u00f0t\u00de \u00bc ui\u00f0t Dt\u00de if otui < 0 \u00f04\u00de where Dt is time interval in the numerical integration. Mindlin plate equations (Mindlin, 1951) may be written in vector form as D\u00bd\u00f01 m\u00de$2W\u00fe \u00f01\u00fe m\u00de$U =2 jGh\u00f0W\u00fe $w\u00de \u00bc \u00f0qh3 =12\u00deottW \u00f05\u00de jGh\u00f0r2w\u00fe U\u00de \u00fe p \u00bc qhottw \u00f06\u00de where U = $ W, D = Eh3/12 (1 m2), W = {wr,wh}T is the vector of rotations, w is transverse displacement, (q,m) are density and Poisson ratio, (E,G) are Young and shear moduli, j is shear constant, h is thickness, t is time, p is applied pressure, $2 is the Laplacian and $ is the gradient operator", + " Expand {wr,wh,w}T in its eigenset wr\u00f0r; h; t\u00de \u00bc XN n\u00bc0 XM j\u00bc1 anj\u00f0t\u00degrnj\u00f0r\u00de cos nh wh\u00f0r; h; t\u00de \u00bc XN n\u00bc0 XM j\u00bc1 anj\u00f0t\u00deghnj\u00f0r\u00de sin nh w\u00f0r; h; t\u00de \u00bc XN n\u00bc0 XM j\u00bc1 anj\u00f0t\u00deunj\u00f0r\u00de cos nh \u00f028\u00de where \u00f0M;N\u00de is the number of radial and circumferential modes in the expansion. Substituting (28) in (5) and (6) and enforcing orthogonality of the eigenfunctions yields uncoupled differential equations in the generalized coordinates anj(t) \u20acanj\u00f0t\u00de \u00fe x2 njanj\u00f0t\u00de \u00bc pnj\u00f0t\u00de=Nnj Nnj \u00bc \u00f01\u00fe dn0\u00depqh Z rd 0 \u00f0u2 nj \u00fe h2\u00f0g2 rnj \u00fe g2 hnj\u00de=12\u00derdr \u00f029\u00de where \u00f0 \u00de is derivative w.r.t. t and dn0 is the Kronecker delta. Let y be the instantaneous circular segment defining the footprint (0 6 y 6 2rd) (see Fig. 1(d)). For y 6 rd, the generalized force pnj(t) in (29) takes the form pnj\u00f0t; y\u00de \u00bc p\u00f0t\u00de pnj\u00f0y\u00de \u00bc 2p\u00f0t\u00de Z hy 0 Z rd r1\u00f0h\u00de unj\u00f0r\u00derdr cos\u00f0nh\u00dedh hy \u00bc cos 1\u00f0\u00f0rd y\u00de=rd\u00de; r1\u00f0h\u00de \u00bc \u00f0rd y\u00de= cos h \u00f030a\u00de p(t) is the instantaneous uniform applied pressure over the footprint sector. Also, for y > rd pnj\u00f0y\u00de \u00bc \u00f0 1\u00den\u00fe1 pnj\u00f02rd y\u00de \u00f030b\u00de _y is related to obliquity angle a and V(t) by _y \u00bc V\u00f0t\u00de= tan a \u00f031\u00de V(t) is the instantaneous rigid body translation velocity of the disk during collapse of the crushable material. For a free disk, (5) and (6) admit a rigid body translation p00\u00f0t; y\u00de p\u00f0t\u00der2 d cos 1 1 f\u00f0 \u00de \u00f01 f\u00de 1 \u00f01 f\u00de2 1=2 \u00bc N00rd tan\u00f0a\u00de\u20acf N00 \u00bc pqhr2 d; f \u00bc y=rd \u00f032\u00de and a rigid body rotation # about a diameter p10\u00f0t; y\u00de 2 3 p\u00f0t\u00der2 df\u00f02 f\u00de 1 \u00f01 f\u00de2 1=2 \u00bc N10rd \u20ac# N10 \u00bc pqhr2 d\u00f01\u00fe \u00f0h=rd\u00de2=6\u00de=4 \u00f033\u00de In a 2D plate or \u2018\u2018strip\u201d model, let (X,Y) be a global coordinate system in the plane of the plate, where X is finite along the strip and Y is infinite (see Fig. 1(e)), and x is a local running coordinate with origin at each corner. Although the strip itself is continuous, the analysis segments it into branches joined at corners. The kth corner joins branches k to k 1 by a (2 2) mobility matrix Yck, and connects to a rigid base by a (2 2) impedance matrix Zck. For example, in the frequency domain where time dependence is sinusoidal with radian frequency x, the Z\u00f01;1\u00deck component of Zck has the form Z\u00f01;1\u00deck \u00bc \u00f0 mcx 2 \u00fe kc \u00fe i\u0302x1c\u00dek; i\u0302 \u00bc ffiffiffiffiffiffiffi 1 p \u00f034\u00de In (34), mc, kc and Bc are lumped mass, spring stiffness and viscous damping of a 1-degree of freedom element at the kth corner", + " Results of the discrete-chain are presented first where the plate is assumed rigid. Results on response including plate elasto-dynamics and comparison with measured data follow. The case of \u2018\u2018fixed-crushable\u201d is discussed and compared to the \u2018\u2018moving-crushable\u201d case. Finally, response of a strip simulating an electronic board is presented. In all discussions to follow, disk displacement w is the same as u in the discrete-chain. Before discussing the inelastic response of the crushable material by the discrete-chain in Fig. 1(c), a comparison between results from the linear discrete-chain model and those from an analytical model El-Raheb (1993) is presented to evaluate convergence of the discrete-chain model. Consider the stack of two layers: 3.8 cm steel on 15.2 cm of linear elastic foam moving at Vo = 40 m/s and striking a rigid boundary. Properties of steel and linearized foam are listed below Steel : Es \u00bc 207 GPa; qs \u00bc 8 g=cm3; Zs \u00bc ffiffiffiffiffiffiffiffiffi Esqs p \u00bc 40:6 MPa=\u00f0m=s\u00de Foam : Ef \u00bc 6:9 MPa; qf \u00bc 1:3 g=cm3; Zf \u00bc ffiffiffiffiffiffiffiffiffi Ef qf q \u00bc 0:095 MPa=\u00f0m=s\u00de \u00f048\u00de Zs, Zf are acoustic impedances", + " Both approximations have no physical basis and are intended to allow for some energy dissipation. The comparison above suggests that 200 cells suffice for convergence of the discrete-chain model. Before discussing results of the discrete chain model with non-linear springs, a description of the constitutive models of honeycomb and metal foam is necessary. For clarity in what follows, \u2018\u2018material\u201d will refer to the parent material, while \u2018\u2018structure\u201d will refer to the collapsible structure made of that parent material. For honeycomb, a typical constitutive law is shown in Fig. 1(a). For small strain e and up to critical stress rcr, the construction is linear elastic r = Eee with modulus Ee that is somewhat lower than that of the parent material E because it is a thin structure rather than a bulk material (Gong et al. (2005)). For e > ecr, the structure collapses plastically and r follows an average constant value rcr until e = ed the densification strain, then changes to r Ede where Ed is densification modulus that is larger than Ee but smaller than the parent material bulk modulus Eb = E/3(1 2m). One peculiarity of this constitutive law is that e may vary in the range ecr 6 e 6 ed without changing r, and therefore if it is assumed that after collapse the structure does not resist tension, then e may rise then fall without affecting r. Clearly the mechanism of how the structure unloads for e > ecr dictates motion without affecting r. For metal foam, a typical constitutive law is shown in Fig. 1(b). Still, for small e and up to rcr, the structure is linear elastic r = Eee. But for e > ecr, modulus drops to almost zero at ecr yet r rises gradually and smoothly with e as with strain-hardening to join the state of densification when local modulus reaches Ed. In this way and unlike honeycomb, r and e are always uniquely related. The discrete-chain model with non-linear springs is now applied to a rigid mass on crushable material striking a rigid boundary. The first example assumes a 300 kg/m2 mass on honeycomb with a typical constitutive law shown in Fig. 1(a), and with properties listed as \u2018\u2018Example 1\u201d in Table 1. Fig. 4 plots _u and \u20acu histories at three stations along the layer. Fig. 4(a1) and (b1) shows _u and \u00fc at the mass. \u00fc is almost constant till arrest at tstop = 8.6 ms. A minor dip in \u00fc occurs near tdip = 5.5 ms (Fig. 4(b1)), caused by collapse waves from the rigid boundary traveling with speed Vc and arriving at the mass after tc hf/Vc = 6.5ms. Equating tdip with tc gives Vc/Vo = 1.18 > 1. An explanation of the nature of the \u2018\u2018collapse wave\u201d now follows", + " The time tend it takes the footprint to wet the whole plate depends on a and Vo as expressed by the approximate relation tend Ltana/Vo where L is side length of the square plate. The larger Vo/tana is, the shorter tend becomes. The problem with this expression is that Vo drops since collapse pressure decelerates the plate, and pressure eccentricity produces a moment that reduces a. An alternative to the square geometry yielding to analysis is the disk. To conserve mass, disk radius is taken as rd \u00bc L= ffiffiffi p p . The expanding footprint shown as a circular segment in Fig. 1(c) produces a time dependent asymmetric forcing function with generalized force given by (30a,b). For the free disk, Fig. 7 plots pnm\u00f0y\u00de=r2 d versus y/2rd, the normalized chord length (Fig. 1(d)), for the rigid body and a few elastic modes. Fig. 7(a) plots pn;0\u00f0y\u00de=r2 d for the translational n = 0 and the rotational n = 1 rigid body modes. Note that for the translational mode, p0;0 is highest followed by the rotational mode p1;0 that is symmetric about y = rd. As n increases, pn;m diminishes with n and m meaning that lower wave-number modes have a stronger effect on response than higher wave-number modes. For odd n, pn;m is symmetric about y = rd while for even n it is anti-symmetric", + " The sharp \u2018\u2018drop\u201d in prescribed deceleration raises magnitude of post-pulse response but does not affect response during the pulse. The author thanks Dr James Zwissler of the Jet Propulsion Laboratory for his support during this study and for providing the test data. For sinusoidal time dependence with radian frequency x, the linear Mindlin elasto-dynamic equations of a strip are (Mindlin (1951)) Dsoxxw jGshs w\u00fe oxw\u00f0 \u00de \u00bc qs\u00f0h 3 s =12\u00dex2w jGshs\u00f0oxxw\u00fe oxw\u00de \u00bc qshsx 2w; Ds \u00bc Esh 3 s \u00f01 m2 s \u00de=12 \u00f0A1\u00de w, w are displacement and rotation (see Fig. 1e), Es, Gs, ms, qs are Young and shear moduli, Poisson ratio, and mass density, hs is thickness and j is shear constant. The constitutive relations are Q \u00bc jGshs\u00f0oxw\u00fe w\u00de; M \u00bc Dsoxw \u00f0A2\u00de Q, M are shear force and moment. Eq. (A1) admits the solution w\u00f0x\u00de w\u00f0x\u00de \u00bc ek1x ek2x d1ek1x d2ek2x C1 C2 ; dk \u00bc \u00f0k2 k \u00fe k2 s \u00de=kk \u00f0A3\u00de k1,2 satisfy the dispersion relation \u00f0 k2 j \u00fe k2 e \u00de\u00f0 k2 j \u00fe k2 s \u00de 12k2 e =h2 s \u00bc 0; ks \u00bc x=cs; cs \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffi jGs=qs p ; ke \u00bc x=ce; ce \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi Es=qs\u00f01 m2 s \u00de q \u00f0A4\u00de Substituting (A3) in (A2), in terms of the state vector S = {f,g}T S\u00f0x\u00de \u00bc B\u00f0x\u00deC \u00f0A5\u00de where f = {Q,M}T and g = {w,w}T, B(x) is a 2 2 matrix of exponentials andC = {C1,C2}T" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002226_012033-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002226_012033-Figure1-1.png", + "caption": "Figure 1. The tail shaft of the helicopter Mi-26:", + "texts": [ + " The creation of a physical and mathematical model of a full-scale system in 2, 3 and n-different scales lets us limit the results of volume of full-scale studies of the NTS to fixation of one or two leading parameters. A new method has been developed for assessing the elastic-dissipative characteristics of the frictional interaction of the spline couplings of the tail transmission of the Mi-26 helicopter on the example of heavily loaded friction pairs of helicopters, in order to increase their reliability and durability when operating at low temperatures (Figure 1). I - splined coupling with flanges in the bearing; 1, 3 - hollow shafts; 2 - bearings of intermediate hearings; 4 - temperature sensors The operational reliability and efficiency of heavily loaded helicopter couplers is determined not only by speed and load parameters, but also by physico-mechanical, physico-chemical, and tribological parameters of frictional processes in frictional subsystems. The basic purpose of the tail shaft is to transfer the rotating moment from the main gearbox to the tail rotor by means of series-connected elastic elements having certain masses and moments of inertia", + " The analysis of the research has shown that the main faults of the spline couplings of the helicopter transmission are the following: the formation of cracks and delamination of the rubber bearing cage; leakage of lubricant, which stimulates overheating of the coupling and its bearing; deformation and wear products generation of the coupling components; the formation of a sideways clearance in the coupling joints; misalignment of tail shaft bearings; increased outrun of the shaft tube, as well as axle fracture or shaft twisting. During the launch of the helicopter transmission under extremely low temperatures conditions, the stationary temperature monitoring sensors 4 installed in the couplings (Fig. 1) are unable to inform the pilots in a timely manner about any emerging problems. A more advanced splined joint diagnostics technology is required in order to allow real-time identification of any emergency situations, which would improve the safety of long-distance piloting. Dynamics of Technical Systems (DTS 2020) IOP Conf. Series: Materials Science and Engineering 1029 (2021) 012033 IOP Publishing doi:10.1088/1757-899X/1029/1/012033 The tribological system of the spline joint of the transmission of the Mi-26 helicopter (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002969_j.matpr.2021.05.300-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002969_j.matpr.2021.05.300-Figure1-1.png", + "caption": "Fig. 1. Al alloy clutch plate heat flux results.", + "texts": [ + " Calculation of stresses, and greatest weights utilizing EulerLagrange conditions. Developing 3D model of MPCs utilizing Solid edge programming. Exporting the model of MPCs from strong edge to ANSYS 19.2 work bench for FEA. Static and dynamic examination of MPCs utilizing ANSYS work seat at various working condition. Determining the wear, most extreme disfigurement and equal worries for every material utilizing FEA Showing the connection of disfigurement and obstruction property of every material utilizing figures and tables (Fig. 1. Fig. 2. Fig. 3. Fig. 4. Fig. 5. Fig. 6. Fig. 7. Fig. 8. Fig. 9. Fig. 10. Fig. 11. Fig. 12. Table 1). Comparing and conversation of results for choice of better covering materials. Fig. 5. Cast iron clutch plate heat flux results. Fig. 6. Cast iron clutch plate heat flux results. Grip disappointment and harm because of extreme frictional warmth and warmth changes to the grasp counter mate circle frequently happens to a car grips. This circumstance adds to warm weakness to the segment which causes the grasp counter mate plate to break and twist" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003296_j.mechmachtheory.2021.104440-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003296_j.mechmachtheory.2021.104440-Figure4-1.png", + "caption": "Fig. 4. Stiffness and damping arrangement of the bearings, as viewed along the negative z-axis", + "texts": [ + " They have \u201czero\u201d thickness, therefore the total length of a shaft is taken up entirely by the flexible shaft elements. The mass moment of inertia about the transverse (bending) axes (x and y) are taken as half of the polar mass moment of inertia (about the axis of twist, z). The mass and gyroscopic matrices for a rigid disk are as described by Rao et al. [28]. As each pinion and wheel is located at a node, these matrices are of order 5. For the purposes of the eigenvalue problem, the rolling element bearings are simplified to massless elements with constant lateral stiffness and viscous damping (Fig. 4). In reality, the total stiffness of a rolling element bearing changes periodically as a function of the frequency of rotation of the cage, where the rollers are retained. The constant stiffness values used in the eigenproblem are mean values of such a periodic function. In the present investigation, the bearing bending stiffness is neglected. The bearings contribute only to the stiffness and damping of the system. Similar to the gear disks, their individual matrices are of order 5. All matrices are presented in Appendix 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002473_tmag.2021.3068131-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002473_tmag.2021.3068131-Figure3-1.png", + "caption": "Fig. 3. Concept of the proposed method. (a) Motor cross section. (b) Flux density waveforms.", + "texts": [ + " This procedure is iteratively carried out until Hreac is converged. If the above-mentioned method is simply applied to induction motors, vast computational time will be required because the time period of the electromagnetic field in the rotor is often dozens to hundreds times longer than that of the power supply due to the rotor slip. Consequently, the considerable number of minor loops should be considered to solve the rotor field. To overcome this problem, we proposed a fast calculation method based on the symmetric rotor shape of induction motors. Fig. 3 shows the concept of the proposed method. This method is valid when the slip s satisfies the following expression: s = n/(mNbar) (3) where Nbar is the number of rotor bars per pole pair, n and m are arbitrary integers. Fig. 3 indicates the case for n = m = 1. In this case, after one time period of power supply T , the point P1 shown in Fig. 3(a) reaches the location, where point P2 exists before T due to s. Consequently, the variation in the flux density at P1 during the next time period is identical to that of P2 during the previous time period. Therefore, the flux density waveform of P1 for one time period of the rotor can be obtained by connecting the waveforms of P1 to PNbar . In addition, the waveforms of P2 to PNbar can be obtained only by shifting the phase angle, as shown in Fig. 3(b). In this case, the hysteresis loops generated at P1 to PNbar are identical to each other. These relationships can be generalized for arbitrary m and n under the condition of (3), as follows: B(\u03b8, t) = B ( \u03b8 \u2212 n 2\u03c0 pNbar , t \u2212 mT ) (4) H reac(\u03b8, t) = H reac ( \u03b8 \u2212 n 2\u03c0 pNbar , t \u2212 mT ) (5) where p is the number of pole pairs. Therefore, in the proposed method, 2-D FEA is required only for m times period of power supply, whereas 1-D FEA is required only for 1 bar pitch region of rotor core. This method can be applied to most of induction motors, in which the same shaped bars are arranged in equal spaces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000629_9780470612286.ch1-Figure1.4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000629_9780470612286.ch1-Figure1.4-1.png", + "caption": "Figure 1.4. Transformations between the end-effector frame and the world reference frame", + "texts": [ + " Whenever calculating an explicit form of the inverse geometric model is not possible, we can calculate a particular solution through numeric procedures [PIE 68], [WHI 69], [FOU 80], [FEA 83], [WOL 84], [GOL 85] [SCI 86]. In this chapter, we present Paul\u2019s method; Pieper\u2019s method, and Raghavan and Roth\u2019s method are detailed in [KHA 02]. 1.2.3.1. Stating the problem Let fTEd be the homogenous transformation matrix representing the desired location of the end-effector frame RE with respect to the world frame Rf. In general cases, fTEd can be expressed in the following form: fTEd = Z 0Tn(q) E [1.10] where (see Figure 1.4): \u2013 Z is the transformation matrix defining the location of the robot frame R0 in the world reference frame Rf; \u2013 0Tn is the transformation matrix of the terminal link frame Rn with respect to frame R0 in terms of the joint coordinates q; \u2013 E is the transformation matrix defining the end-effector frame RE in the terminal frame Rn. When n \u2265 6, we can write the following relation by grouping on the right hand side all known terms: 0Tn(q) = Z-1 fTEd E-1 [1.11] When n < 6, the robot\u2019s operational space is less than six" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001072_s0129183107011364-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001072_s0129183107011364-Figure1-1.png", + "caption": "Fig. 1. The Gershgorin Circle of L in S plane.", + "texts": [ + " , n . (4) Suppose that the network is connected in the sense that there are no isolate clusters, then L(L = \u2212A) is a symmetric and irreducible real number matrix. It can be shown that zero is an eigenvalue of L with multiplicity 1 and all the other eigenvalues of L are strictly positive. In this case, according to Gershgorin theorem,18 all eigenvalues of L in the complex plane are located in a closed disk centered at D + 0j with a radius of D = maxi ki, i.e., the maximum degree of a graph (see Fig. 1). Therefore, the set of eigenvalues of L can be ordered sequentially in an ascending order as 0 = \u03bb1 < \u03bb2 \u2264 \u03bb3 \u2264 \u00b7 \u00b7 \u00b7 \u2264 \u03bbn \u2264 2D . (5) For a network with time-delay, the linear consensus algorithm takes the following form:14 x\u0307i(t) = \u2211 j\u2208Ni (xj(t \u2212 \u03c4) \u2212 xi(t \u2212 \u03c4)) , i = 1, 2, . . . , n (6) with a collective dynamics x\u0307(t) = \u2212Lx(t \u2212 \u03c4)) = Ax(t \u2212 \u03c4) . (7) Here, we assume that the time-delay in all links is equal to \u03c4 . Before further studying, we present a result and a definition, which are essential throughout the paper" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000775_detc2007-35917-Figure8-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000775_detc2007-35917-Figure8-1.png", + "caption": "Figure 8. Two flexible beams applied with heat flux", + "texts": [ + " It can be also shown that the result obtained by linear model fails to explain for the initial stiffening and the subsequent thermal softening effects. Figure 7 shows the axial stress of the midpoint on the neutral axis. It is clearly seen that without considering the nonlinear strain, the stress result obtained by linear model neglects the positive stretch strain due to the initial stiffening phenomenon. 6 Copyright \u00a9 2007 by ASME erms of Use: http://www.asme.org/about-asme/terms-of-use Downl 6.2 Thermal bending for flexible beam system: Two flexible rectangular beams applied with heat flux are shown in figure 8. B2 is connected to B1, and B1 is connected to the ground by rotational joints. The beams have cantilevered-free boundary condition. The material data of each beam are the same as those given in table 1. The geometric property is: length l = 4m, height h = 0.02m, thickness b = 0.02m. For both two beams, heat flux at upper and lower surfaces are given by q1 = 1\u00d7106 w/m2, q2 = 0. In this simulation, the heat generation and the body force are not taken into account. Initially, the system is in static state without deformation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000250_ichr.2004.1442689-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000250_ichr.2004.1442689-Figure1-1.png", + "caption": "Figure 1 : Reference frames of the biped robot using DH convention", + "texts": [ + " Aiming to combine our intuitive knowledge of walking, with the concept of having separate modules that deals independently with stability and motion, the present paper proposes a control scheme based on fuzzy logic 11 3 1. The dual-support phase of a biped robot was selected in order to investigate the potential o f fuzzy logic in solving the over actuation problem without having to optimize some energy cost function in the process. The objectives of the research presented in this paper are to establish and test a control scheme that is not based on a trajectory following approach, while avoiding at the same time the integration of complex models into the controlIer. Figure 1 shows the 6 DOF biped robot used in the analysis. The orientation and position of the frames at each joint are determined based on the Denavitt Hartenburg convention commonly used in the analysis of serial kinematic chains. The reference frame for the overall system is taken as &gYopZo, located at the foot of the support leg. The non-support leg will be referred as the free leg. The stability of the biped will be analyzed when the free leg is at various positions with respect to the support leg" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000250_ichr.2004.1442689-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000250_ichr.2004.1442689-Figure2-1.png", + "caption": "Figure 2: Biped robot in the dual support phase", + "texts": [ + " During the dual-support phase, where both feet are on the ground, the dynamics of the biped is subject to holonomic constraints if we assume that the contact points between the ground and feet are fmed within the reference frame. These constraints can be described in the dynamic model by using the Lagrange multiplier A j as folIows: 843 where L is the Lagrangian of the system L=K-U, K is the kinetic energy of the system and U is the potential energy, qi is the generalized coordinate of joint i, is the input torque at joint i , and c(q] is the mathematical relation that defines the geometrical constraints. Figure 2 shows the relations between the joint angles when the biped is in the dual support mode. Assuming that a1 = a4 and a2 = a3 , the x and y components of the vector linking the first frame (i.e. left foot) to the last frame (i.e. right foot) must satisfy the following conditions 844 and a1 (sW71) - sin(% + 92 + q 3 + q4 + 95 I) + a2 (sin(q1 + 42 1 - W q 1 + 42 + q3 + 9 4 ) ) d (3) re-arranging the above two equations and denoting by ci = coS(qi) , si = sin(qi), cy = cos(q, + q j ) , and sij = s h ( g i +qi), the geometric constrains can be written in matrix form as Denoting by C the 2x6 matrix defined by C = - - a], and after a 4 n evaluating the Lagrangian of the biped robot, the joint accelerations can be obtained from Equation 1 and written as q = M-1 (r - CTR - N(g , 4) - P(q)) ( 5 ) The CTA terms correspond to the joint torques caused by the reaction forces between the feet and the ground" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002208_012043-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002208_012043-Figure6-1.png", + "caption": "Figure 6. Scheme for calculating the action of forces in the contact zone of the lubricating rods", + "texts": [], + "surrounding_texts": [ + "The considered technology, technological equipment and consumables (optimized aluminum alloy) can significantly improve the technical, economic and ecological characteristics of the friction subsystem \u201clocomotive wheel \u2013 rail\u201d, as well as the entire system \u201ctrack \u2013 rolling stock\u201d. Thermal metal cladding of the working surfaces of the wheel tread of the locomotive and use of the anisotropy of the Al-Fe friction bond allows the following: 5\u20136-fold reduction of the level of traction energy losses associated with the lateral creep forces when the locomotive moves in a pushing mode, when overcoming the wind load, and when moving in curved track sections; a possibility of turning locomotive wheel set treads due to such a defect as rolling along the rolling circle of a wheel set of a tread. Compared to turning due to a defect in the limiting value for the thickness of the flange, it ensures 3\u20134-fold reduction of the cost for the measures for turning locomotive wheel treads; thermal metal cladding of the rolling circle surface of the wheel tread of a locomotive wheel helps to increase the trailer load by 20\u201325% by increasing the adhesion coefficient from 0.27 in case of the frictional coupling of the wheel to the rail in the traditional form Fe-Fe to 0.45 \u2013 0.5 when the Fe-Al friction bond is applied; thermal metal cladding with an optimized aluminum alloy of the working (frictional) surface of the wheel flange allows increasing the draw pull by 15\u201320% in the presence of unbalanced acceleration. It compensates for the resistance forces during the movement of the locomotive (for example, the forces of resistance to the movement of the rolling stock as a \"carriage\" in curved track sections); thermal metal cladding of the working surfaces of the wheel tread of the locomotive allows excluding (reduction) of using sand as an activator of the adhesion coefficient. It excludes the phenomenon of oversanding the ballast prism of the railway track and, as a result, 1.5\u20131.8-fold increase in the interrepair time; it also reduces the expenses of traction energy by 2\u20133%; Dynamics of Technical Systems (DTS 2020) IOP Conf. Series: Materials Science and Engineering 1029 (2021) 012043 IOP Publishing doi:10.1088/1757-899X/1029/1/012043 the application of the technology of thermal metal cladding helps to increase the mileage of a locomotive from 1.5 thousand km with its one-time servicing to 8\u201310 thousand km when filling it with briquettes of the TMC modifier." + ] + }, + { + "image_filename": "designv11_83_0002796_j.matpr.2021.04.285-Figure4-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002796_j.matpr.2021.04.285-Figure4-1.png", + "caption": "Fig. 4. Modified tooth Spur Gear Profile.", + "texts": [ + " One DOF in the axial direction of the gear was arrested, slippage in the gear was reduced, and misalignment due to gear shaft inclination was reduced with this type of arrangement. The groove cut into the tooth must be higher than the gears\u2019 contact line. The extrusion must be positioned under the contact line. On each tooth of the gear, an extrusion and groove were formed. To reduce noise and increase smoothness, the extrusion and cut can use either concave or convex profiles. The force acting on the extrusion was counteracted by the same force acting on the gears in the opposite direction. The Fig. 4 shows the spur gear with modified tooth profile. Extrusion radius from the centre of the space breadth and on the pitch circle, Er = 0.75 m Groove radius from the Centre of the circular thickness on the pitch circle, Gr = 0.75 m Length of the groove from the addendum circle, Gl = 0.25 m Extrusion length from the clearance circle, El = 0.25 m Contact length, Cl = 0.25 m Extrusion length from the dedendum circle Edl = 0.5 m Clearance during meshing Cm = 0.25 m Where, m = module of gear Width of the groove and extrusion, Gw = Ew = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003453_s10846-021-01454-7-Figure13-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003453_s10846-021-01454-7-Figure13-1.png", + "caption": "Fig. 13 Turning period at case 1", + "texts": [], + "surrounding_texts": [ + "The proposed design of the swimming robot is simulated by SolidWorks\u00ae software and validated experimentally. The starting angle of rotation is set to 50o. The robot will oscillate in a 100o amplitude (i.e. from 50o to -50o) as in ref. [31]. A computational domain of exact dimensions of the swimming pool (i.e. 1 m \u00d7 0.65 m \u00d7 0.65 m) has been used throughout this experiment as shown in Fig. 9. The flow type is set to have (laminar and turbulent) option. The static pressure of 101,325 Pa at 293.2 K is used, and a local mesh of 6 levels of refinement cells are used throughout this simulation." + ] + }, + { + "image_filename": "designv11_83_0002581_icees51510.2021.9383730-Figure12-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002581_icees51510.2021.9383730-Figure12-1.png", + "caption": "Fig 12. Contours o f Temperature with 10mm Fins", + "texts": [ + " By introducing rectangular fins, the upmost temperature on the terminal windings will be reducing. The temperatures of the windings are reduced by a further increase of the fin\u2019s depth. By using the rectangular fins, the highest temperature of the motor\u2019s critical components will be reduced. Rectangular fins of 2 mm thick and 5 mm deep are introduced. This reduces the upmost temperature on the terminal windings by about 11%, as in fig. 11. The depth of the fins if further increased to decreases the windings temperature up to 17%. This is evident from fig. 12. However, in designing the fins, utmost care should be taken, as it should not comprise the rigid structure of the housing. 0,025 0,075 In the end windings there is no significant effect on the orientation of the fins on the temperature as seen in fig. 13. In this case, fins of 10 mm depth, with axial orientation were employed as contrary to the radial orientation considered previously. S.No T em perature and H eat transfer rate Fins (mm) Temperature(\u00b0C) Rate o f keat transfer(W) 1 5 mm (or) 0.005 m 59" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000368_detc2004-57064-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000368_detc2004-57064-Figure7-1.png", + "caption": "Figure 7. A Parallel Manipulator Without Relative Motion Between the Platforms and With Self-Motion.", + "texts": [ + "asmedigitalcollection.asme.org/pdfaccess.ashx?u namely, they are the subalgebras associated with planar subgroups. Moreover, the set e\u03021 e\u03022 e\u03023 is linearly independent. Furthermore, Am f a 3 j 1 Am f j 0 Therefore, there is no relative movement between the fixed platform and the moving platform. Furthermore, since 3 \u2211 i 1 f ji 3 dim Vm f j dim Am f j j 1 2 3 It follows that Fp \u2211 f ji 3 \u2211 j 1 Vm f j 0 Hence, there are no passive degrees of freedom; i.e. there is no self-motion. 5.2 Fixed Platform With Self-Motion. Figure 7 shows a parallel platform, in which there is no relative movement between the moving and fixed plataforms, this class of platforms was analyzed in Section 2.1. Two of the serial connecting legs are formed by three revolute pairs whose axes are parallel. The remaining serial connecting chain is formed by Copyright \u00a9 2004 by ASME rl=/data/conferences/idetc/cie2004/71535/ on 05/01/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Do four revolute pairs whose axes are parallel. Therefore, the closure algebra associated with each leg are given by Am k j ple\u0302 j ; namely, they are the subalgebra associated with the planar subgroup" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002969_j.matpr.2021.05.300-Figure9-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002969_j.matpr.2021.05.300-Figure9-1.png", + "caption": "Fig. 9. Kevlar material clutch plate temperature results.", + "texts": [ + " Developing 3D model of MPCs utilizing Solid edge programming. Exporting the model of MPCs from strong edge to ANSYS 19.2 work bench for FEA. Static and dynamic examination of MPCs utilizing ANSYS work seat at various working condition. Determining the wear, most extreme disfigurement and equal worries for every material utilizing FEA Showing the connection of disfigurement and obstruction property of every material utilizing figures and tables (Fig. 1. Fig. 2. Fig. 3. Fig. 4. Fig. 5. Fig. 6. Fig. 7. Fig. 8. Fig. 9. Fig. 10. Fig. 11. Fig. 12. Table 1). Comparing and conversation of results for choice of better covering materials. Fig. 5. Cast iron clutch plate heat flux results. Fig. 6. Cast iron clutch plate heat flux results. Grip disappointment and harm because of extreme frictional warmth and warmth changes to the grasp counter mate circle frequently happens to a car grips. This circumstance adds to warm weakness to the segment which causes the grasp counter mate plate to break and twist. This later will make issues, for example, grip slip, grasp drag or disappointment of hold to separate appropriately and grip shaking just as shortening the lifecycle of the part" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001844_cp:20080483-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001844_cp:20080483-Figure3-1.png", + "caption": "Figure 3. Generation of the reference vectors in a MC using the SVM based indirect modulation model:", + "texts": [ + " Further advantages related to the improved efficiency and installed power per kW are investigated to justify the merits of this proposal. amplitude and angle. The proportion between the duty-cycles of the two adjacent vectors gives the direction and the duty cycle of the zero-vector determines the magnitude of the reference vector., A widely used method to implement modulation on MCs is the indirect approach [2], which considers the MC as having two virtual stages. The input current vector lin (actually only its angle) is the reference of the rectification stage (Fig.3a) and the output voltage vector VOul is the reference of the inversion stage (Fig.3b), but will also apply for the auxiliary VSI. The duty-cycles of the active switching vectors for the rectification stage, I r' 18 are given by (1) and the duty-cycles of active switching vectors for the inversion stage, Va' Vp are given by (2) as proposed in [4]. results that the average voltage over a switching period delivered by the virtual rectifier stage (3) reaches a minimum of 0.866 the peak line-to-line voltage, justifying therefore the theoretical MC voltage transfer ratio limit. The inverter stage may use a double-sided asymmetric PWM switching sequence (01 - a1 - /31+2 - a2 - O2), but with unequal sides (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002736_978-3-030-70493-3_5-Figure6.9-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002736_978-3-030-70493-3_5-Figure6.9-1.png", + "caption": "Fig. 6.9 Gating system concepts", + "texts": [ + "3 Production flow processes for the bunk bed . . . . . . . . . . . . . . . . . . . 116 Fig. 6.4 Schematic of the reorganised furniture manufacturing plant . . . . . . . 118 Fig. 6.5 Five-stage process flows for industrial pallets . . . . . . . . . . . . . . . . . . 120 Fig. 6.6 Ten-stage process flows for domestic baby tenders . . . . . . . . . . . . . . 120 Fig. 6.7 Mathematical and simulation model for an s-stage process flows . . . 121 Fig. 6.8 Process and waiting times before and after optimisation . . . . . . . . . . 130 Fig. 6.9 Gating system concepts . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 132 Fig. 6.10 The platinum company\u2019s comminution and flotation circuits . . . . . . 137 Fig. 6.11 (a) Comminution resource utilisation, (b) flotation resource utilisation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 139 List of Figures xxii Fig. 7.1 Relationships of industry with higher education institutions . . . . . . . 151 Fig. 7.2 Joint projects and sharing of resources between industry and HEIs ", + " The challenges observed included difficulties in directing molten metal from the ladle into the mould resulting in splashing and loss of molten metal. In addition, a considerable amount of waste also solidified in the runners and gates, apart from the evident turbulence and vortex formation in the flow of molten metal due to the uniform diameter of the cylindrical down-gate, often resulting in distortions and defects in the grinding balls. As part of the way to resolve these challenges in order to optimise the casting technology, three possible concepts of the gating system were considered, as shown in Fig.\u00a06.9, modelled using Solid Works software. Molten metal can be easily directed and poured using the conical funnel in Concept (a) in order to reduce splashes. The tapered down-gate also reduced the chances for vortex formation and turbulence owing to the constricted opening. According to Bernoulli\u2019s principle of fluid dynamics, this would also increase the 6.5 Case Studies 132 speed of flow. The overflow bores on the side of the chamber would be useful to indicate that the mould cavities were full. Although the second concept (b) was almost similar to the first, it had a wide offset pouring basin which directed molten metal into the sprue to avoid the ladle coming close or in contact with the sprue" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001750_icelmach.2008.4799891-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001750_icelmach.2008.4799891-Figure5-1.png", + "caption": "Fig. 5. Contours of flux density.", + "texts": [ + " 4 shows the principle of operation of the spherical resonant actuator. The magnetic flux caused by the permanent magnet is also shown in Fig. 4. The armature is stable in the center because of balanced magnetic flux in the gap between the armature and the stator. The torque is generated on the armature when current is excited to the coil as shown in Fig. 4, and then the armature rotates in y-axis. Table I shows the analysis conditions. In this calculation, the rectangular voltage is applied to the coil. \u03b8rotation Fig. 5 shows the contours of flux density in the spherical resonant actuator. As shown in Fig. 5(a), the flux density in the stator core is small with the rotation angle of 0\u00ba, and the flux density in the stator core becomes large at the rotation angle of 3\u00ba. As shown in Fig. 5(b), the flux density in the stator core becomes large as the rotation angle of the armature increases. Fig. 6 shows the calculated and measured static torque characteristics with the excitation of 0AT and 100AT. The actuator is stable at the rotation angle of zero because of the balanced magnetic flux mentioned above. The torque remains nearly constant towards rotation angle. The average torque constant is 0.3mN\u00b7m/AT. The calculated results well agree with the measured ones. Fig. 7 shows the waveforms of the voltage and the current" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002654_s42417-021-00299-6-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002654_s42417-021-00299-6-Figure1-1.png", + "caption": "Fig. 1 Physical model of the DI-SO system", + "texts": [ + "ol.:(0123456789) Keywords DI-SO helical gears\u00a0\u00b7 Unbalanced load inputs\u00a0\u00b7 Time-varying parameters\u00a0\u00b7 Nonlinear behaviors\u00a0\u00b7 Evolution laws DI-SO helical gear system with unsymmetrical input loads has been widely used in the propulsion devices of large-scale warships, submarines, and aircraft with the advantages of high-power density and compact structure. The 3D physical model of the DI-SO helical gear system supported by rigid mounts is shown in Fig.\u00a01. However, the nonlinear instability phenomena of meshing separation and impact have become the key problems restricting the safety and stability of such gear system, which is confirmed by some ship tests. On the one hand, due to the coupling of various nonlinear factors such as gear backlash, multi-state meshing of drive-side tooth meshing, tooth separation and back-side tooth meshing exist. On the other hand, for the unsymmetrical inputs, with the increase of the load ratio between the two inputs, the driving gear of the low-load input changes from energyconsuming to do-work state, which means this gear pair changes from driven to driving meshing state and dynamic instability happens in the process, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003270_s41403-021-00254-7-Figure10-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003270_s41403-021-00254-7-Figure10-1.png", + "caption": "Fig. 10 LPBF route 3D printed Ti6Al4V-ELI\u00a0 end connectors after machining of interfaces", + "texts": [ + " The anisotropy strongly depends on the microstructure evolution during the LPBF process. Pores or micro-cracks are the types of defects present in 3D printed parts which can result in unacceptable mechanical performance particularly in terms of fatigue resistance. The martensitic \u03b1\u2019 phase present in Ti6Al4V components due to a high cooling rate results in a higher strength compared to that in conventionally processed wrought alloy. It is often accompanied by a decrease in ductility. Dimensional Inspection End connectors and brackets were machined at the interfaces as shown in Fig.\u00a010 for facilitating its assembly for the intended application. The dimensional inspection was carried out in a SURFCOM-make 130A-model coordinate measuring machine (CMM). The surface roughness (Ra) values also measured are found to be 8\u201312\u00a0\u00b5m in the as 3D printed and shot-blasted condition and Ra of 0.4\u00a0\u00b5m was achieved at the machined interfaces. Parts were found to be dimensionally acceptable for their application in launch vehicle inter-stage. 1. 3D printing of end connectors and brackets in Ti6Al4VELI material was successfully carried out through LPBF AM process" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002838_012021-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002838_012021-Figure1-1.png", + "caption": "Figure 1. Wide drop center of the hump wheel [13]", + "texts": [ + "1088/1757-899X/1116/1/012021 tyres and many times can cause structural failure [10]. So, focusing on these issues we have some more promising and long-lasting solution without compromising to safety [11]. This analysis will help to predict the best material for the rim out of the listed materials without compromising the performance and safety. The Automotive wheel is one of the indispensable parts of the vehicle frame at a different offset of wheel [12]. The offset of the wheel rim is the distance between the hub surface to the center of the wheel, as illustrated in figure 1. 1.1. Materials Selected Different types of materials have been selected and used for a 3D model of the rim. List of materials used for the analysis are: 1. Al 6061 T6 2. Structural Steel 3. Kevlar 29 4. S-Glass 5. E- Glass 6. Basalt Fiber 7. Carbon Fiber 8. HSLA Steel Properties of selected materials for the analysis of rim as shown in Table 1. 1.2. Modeling of Rim FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 Rim was designed by using 3D modeling software solid works, 2019 version [14]- [15]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002822_s12034-021-02466-7-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002822_s12034-021-02466-7-Figure2-1.png", + "caption": "Figure 2. Stress nephogram.", + "texts": [ + " The liquid lens can produce a large tuning range in the condition of large deformation. There are also problems, for example, the volume of encapsulated liquid is usually less than 5 ml, which results in a small tuning range; two different lens curvatures would lead to unclear imaging and unmeasurable focal length. These problems could be effectively solved by adjusting refractive index and choosing suitable lens size. The software of ABQUS is utilized to simulate the lens structure and the deformation. As in figure 2, the lens becomes an approximate spherical shape after actuating. Meanwhile, the deformation is manifested by increase in lens curvature. The stress nephogram reflects the stress of each part under voltages in the figure. The red part illustrates the larger stress at the junction between the DEA and the frame. Figure 3 demonstrates the relationship between the lens projection area on the DE central region and the voltage on the DEA. The lens diameter 2A is 2, 2.4 cm or 3 cm before actuating, which equals the inner diameter of the DEA" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000930_20080706-5-kr-1001.00143-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000930_20080706-5-kr-1001.00143-Figure1-1.png", + "caption": "Fig. 1. Schematic of the dual-stage hard drive", + "texts": [ + " Although this paper does not consider multiobjective control design, all of the techniques presented here could be trivially extended to allow for control design with multiple guaranteed `2 semi-norm performance objectives. In particular, we would apply Theorem 5 to each objective, i.e. each choice of (L,R). In this case, since our controller reconstruction does not depend on the values on Wk, FAB k , FCD k , P 1 k , P 2 k , and P 3 k , these variables can be allowed to be constraint-dependent, i.e. each application of Theorem 5 with (11) could have different values for these variables. Figure 1 shows the structure of a hard disk drive with dual-stage actuation. The model of our system in discrete time with uncertain parameters is given by[ p(z) y(z) ] = Hgen(z) [ w(z) u(z) ] (14) where p := yh \u2212 r yh \u2212 ys uv um w := wr wa nPES nRPES y := [ yh \u2212 r + nPES yh \u2212 ys + nRPES ] u := [ uv um ] [ yh(z) ys(z) ] := 4\u2211 i=1 [ 1 0 c1,i c2,i ] ( zI \u2212 [ a1,i 1 a2,i 0 ])\u22121 Bi [ wa(z) uv(z) um(z) ] aj,i := aj,i + mj,i\u03b4j , \u03b4j \u2208 [\u22121, 1] r(z) := ( 2.355 z \u2212 0.9627 + 0.557z + 0.5541 z2 \u2212 1.984z + 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003062_14644207211026696-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003062_14644207211026696-Figure6-1.png", + "caption": "Figure 6. Determination of critical root section in asymmetric gear.20 Source: reproduced with permission from Prabhu and Muthuveerappan, 2015.20", + "texts": [ + " Symmetric gear Asymmetric gear Number of teeth 24 Module (mm) 3 Number of manufactured teeth 8 Tooth width (mm) 6 Pressure angle ( ) 20/20 20/22\u201320/25 Addendum circle diameter (mm) 78 Dedendum circle diameter (mm) 64.5 Material AISI 4140 Hardness (HRC) 38 1.5 Yield stress (MPa) 1150 YS \u00bc \u00f01;2\u00fe 0;13L\u00deqas (3) L \u00bc sF hf (4) qs \u00bc sF 2qf (5) a \u00bc \u00bd1;21\u00fe 2;3=L 1 (6) fFc \u00bc 30\u00fe ad ac (7) In the asymmetric gear, tooth root stress is calculated considering the new tooth form factor that changes with the geometric form (Figure 6). The new tooth form factor is as follows YF0 \u00bc 6 hf mn cosaL sf2cosan mn sinaL sf cosan (8) Tooth root stress rF0 is included in the calculation, taking into account the new YF value rF0 \u00bc Ft YF0 Ys b mn (9) Single tooth bending fatigue failure tests were carried out to evaluate the fatigue strength of symmetrical and asymmetrical gears. Test results are presented in graphs for the same torque values where the experiments were carried out. In the tests, the load reduction was reduced by 10%, similar to the staircase method, and, according to the test results, intermediate value loadings and retests were performed to verify the results of the tests and to see the intermediate value results" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001105_0022-4898(65)90129-1-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001105_0022-4898(65)90129-1-Figure1-1.png", + "caption": "FIG. 1. Mathematical model of the army M37 3/4-ton truck.", + "texts": [ + " The engineering programmer can reduce the work required to develop a model for various vehicles by starting with a generalized mathematical model. This model consists of a series of mathematical expressions which describe the total motion of the vehicle assembly while traversing a cross-country terrain and can be written in general enough terms so the solution can be applied to any tracked or wheeled vehicle. In the case of a wheeled vehicle with sprung differential assemblies, additional bounce and roll equations are required for each differential assembly. A typical mathematical model of a two axle wheeled truck can be seen in Fig. 1. The stiffness of the vehicle body and differential assemblies are very high compared to the leaf springs and tyres. The body and axle springiness has been omitted from the mathematical model to simplify its development with little or no sacrifice in solution accuracy. The vertical bounce of the body of the vehicle with respect to its centre of gravity can be evaluated by the following differential equation of motion. 2n 2n n 2i a j ~ l j = l i = 1 1=2 i - -1 t = 1 2n #=1 2n n 2i n kjlj - \"~ ct - , k j - t = 1 i = l j = 2 i - - 1 i = 1 n 2i n 2i g i = l #=2i - -1 i=1 j = 2 / - - 1 2i i = 2 i - 1 2i x,c, j = 2 i - 1 The positive forces are directed toward the ground, and the equation includes all of the leaf spring, shock absorber and accelerating forces which are acting on the vehicle frame" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002919_012093-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002919_012093-Figure3-1.png", + "caption": "Figure 3 Monitoring points", + "texts": [ + "3 Fluid-solid-thermal interaction equation In the fluid-solid-thermal bidirectional coupling calculation, the data transferred in the fluid-solid interface need to satisfy the pressure, displacement, heat fluxes and temperature conservation: f f s s f s f s f s n n u u q q T T ( 7 ) where f represents the fluid, s represents the solid. Tilting pad thrust bearing used in this paper is consisted of 10 fan-shaped pads, the model is shown in Figure 1, the structure parameters is shown in Table 1, which is as same as reference [10], and the numerical model is shown in Figure 2. Bearing lubrication performance is monitored by monitoring points, including 4 monitoring points at pad surface (p1 ~ p4), 4 monitoring points at collar surface (c1 ~ c4), and 4 temperature monitoring points (t1 ~ t4), as shown in Figure 3. 9. 30th IAHR Symposium on Hydraulic Machinery and Systems IOP Conf. Series: Earth and Environmental Science 774 (2021) 012093 IOP Publishing doi:10.1088/1755-1315/774/1/012093 Pad angle \u03b8/deg 31 Pad width B/mm 560 Pad thickness H1/mm 203 Babbitt layer thickness, H2/mm 3 Number of pads n 10 Radial pivot position Or/mm 1065 Circumferential pivot eccentricity Oc/% 50 Collar outer radius R3/mm 1335 Collar inner radius R4/mm 775 Collar thickness H2/mm 660 Solid226 Structural-Thermal element is used in solid domain, each node contains 4 degrees of freedom (Ux, Uy, Uz and Temperature)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003132_978-981-16-0119-4_20-Figure19.10-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003132_978-981-16-0119-4_20-Figure19.10-1.png", + "caption": "Fig. 19.10 Stresses due to loading", + "texts": [], + "surrounding_texts": [ + "Amanufacturing team needs a drawing and specifications communicating the design (to the extent it impacts\u2014functional performance) (see Figs. 19.11 and 19.12).\nThe framework described above leverages technology to drive efficient design.", + "238 S. K. Mukherjee", + "It provides an example of proposing concepts that are a suitable option from strength and stiffness standpoint, while meeting form and functional requirements.\nStudents love to design products which consider human factors and are pleasant to use. The framework embodied in this example could be applied to other products and/or aggregates as an integral attempt to propose a design that is \u2018desired by customer\u2019 and which is appreciated by engineers and manufacturers. For those interested in designing bicycles for design competitions, this case study could be adapted to comply with the \u2018approval protocol\u2019 of UCI.8\nOptimization as a process has been used in industry for over a decade [14]. Virtual testing has been used in automotive industry to estimate \u2018fatigue life\u2019 [15]. The hypothesis that \u2018drawings are important in mechanical design process\u2019 has been evaluated, and support for the hypothesis has been reported [16].\nThere may be other case studies that already provide frameworks which leverage technology to incubate aspects of the \u2018applied sciences\u2019 in design of products. Considering the existing focus on inter-disciplinary studies in product design among researchers, this case study is an addition to these efforts.\n8UCI is an abbreviation of \u2018Union Cycliste Internationale\u2019\u2014Document: \u2018Approval Protocol for Frames and Forks\u2019. See: https://www.uci.org." + ] + }, + { + "image_filename": "designv11_83_0000146_bf02329037-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000146_bf02329037-Figure2-1.png", + "caption": "Fig. 2--Restoring-force characteristic of a bilinear spring", + "texts": [ + " However, the dynamic analysis of nonl inear continuous systems has received less at tent ion than nonl inear s ing le -degree -o f - f reedom systems. This paper deals wi th the t ransverse vibrat ions of a beam having one end clamped and the other supported on a nonl inear spring. The spring-supported end also carr ies a concentra ted mass. Figure 1 shows the schematic d iagram of the system. The beam is being exci ted t ransverse ly by a sinusoidal force F sin ~t. The spring has a bi l inear characterist ic as shown in Fig. 2. The solution of the system has been obtained by extending and modifying the methods avai l - able for nonl inear s ing le -degree-of - f reedom systems. \"~ The system considered in the present paper may be a simple representat ion of s tructures where one end of the beam has sufficiently rigid support but the other end has compara t ive ly less stiff support. The range of deflection m a y be such that the motion of the beam can be described by a l inear par t ia l-different ial equat ion but the assumption of a l inear support may not be reasonable" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002796_j.matpr.2021.04.285-Figure5-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002796_j.matpr.2021.04.285-Figure5-1.png", + "caption": "Fig. 5. Meshed view of Spur Gear.", + "texts": [ + " It is, in reality, a computational approach for solving structural problems. However, by increasing the number of equations before the desired precision is achieved, errors in the estimated solution can be reduced. This is an alternative to empirical approaches for obtaining precise answers to analysis problems. In FEM, complex structures are broken down into a finite number of tiny regions called elements. The field quantity is expressed as unknown values at the nodes, which are the corners of these elements shows in Fig. 5, and the property of the material is shown in Tables 2 & 3 and prevailing equations are considered for these elements. A torque of 21.517 Nm was applied to the two faces of the driving gear. For the pair of spur gears, the gear ratio is 2. From the Name Equivalent Elastic Strain Equivalent Stress Total Deformation State Solved Results Minimum 3.0131e-10 m/m 36.811 Pa 1.3647e-6 m Maximum 0.00049254 m/m 9.3355e7 Pa 7.9789e-6 m obtained results, the torque applied over the driven gear is converted into two times the torque acting on the driving gear is shown in Table 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001848_robot.2007.363783-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001848_robot.2007.363783-Figure2-1.png", + "caption": "Fig. 2: Intermediate coordinates description for the left", + "texts": [ + " The moving platform coordinate and the base coordinate are schematically shown in Fig. 1. The transformation matrix of the moving platform D 1-4244-0602-1/07/$20.00 \u00a92007 IEEE. 175 coordinate w.r.t the base frame coordinate using X-Y-Z Eulerian angles of Rotation is =TB P \u2212 \u2212+ +\u2212 1000 zCCSCS ySCCSSCCSSSCS xSSCSCCSSSCCC \u03b3\u03b2\u03b3\u03b2\u03b2 \u03b3\u03b1\u03b3\u03b2\u03b1\u03b3\u03b1\u03b3\u03b2\u03b1\u03b2\u03b1 \u03b3\u03b1\u03b3\u03b2\u03b1\u03b3\u03b1\u03b3\u03b2\u03b1\u03b2\u03b1 For the inverse and forward pose kinematic analysis, the standard Denavit-Hartenberg parameters are required. The selected intermediate coordinates based on the DenavitHartenberg notation are shown in Figure 2. The spherical, prismatic and universal joints need three, one and two coordinate frames, respectively. The constant platform parameters as well as the Denavit-Hartenberg parameters for the left and right legs of the manipulator are: B: Distance between two spherical joints. P : Distance between two universal joints. Rf ll , : Length of the left and right legs, respectively. ),,(),,,( 321321 RRRfff \u03b8\u03b8\u03b8\u03b8\u03b8\u03b8 : Spherical joints variables for the left and right legs, respectively. leg The X and Z axes for each coordinate frame are presented in Fig. 2 and the Y axis can be found via the right-hand rule. The Denavit-Hartenberg Parameters for the selected intermediate coordinates are given in table1. IV. INVERSE POSE KINEMATIC In the inverse pose kinematic the desired actuator variables which are the spherical joint variables, should be calculated having TB P . First the basic idea which leads to the calculation of the active joint variables with no need for the evaluation of the passive joint ones will be discussed and later its equivalent mathematical repression will be introduced", + " Since the direction of leg and 6z is known, the direction of 5z will be available by cross producting 6z and the leg\u2019s direction. It is to be noted that, the direction of the leg is available having 21 ,\u03b8\u03b8 . Also, the 6z direction is the same as the direction of the normal vector of the manipulator which is known. In the following the mathematical equivalent of the above geometrical interpretation is provided. The index f in the following formulation indicates that the related matrix evaluated for the left leg. The transformation matrix between coordinates (0) and (6) of the left leg, Fig. 2, using the Denavit-Hartenberg parameters given in table 1, is: ff TTTTTTT )()( 5 6 4 5 3 4 2 3 1 2 0 1 0 6 = (1) This transformation is also: fP B Pf B f TTTT )()()( 161 0 0 6 \u2212\u2212= (2) In which TT P B 6 0 , are: = 1000 0100 0010 5.001 )( 0 B fBT \u2212 \u2212 = 1000 0100 0010 5.001 )( 6 P fPT Equaling Eqs. (1) and (2), by the fact that TB P is completely known , we have: )2.(, 161 0 3 2 1 1 0 0 0 )()( 1 Eqofcolumnforthknowncompletely fP B Pf B TTT k k k = \u2212\u2212 = )1", + " This second order polynomial is: \u2212= \u2212= \u2212+= \u2212\u2212\u2212= \u2212+= =++ EHI EGF DBFFc CBIFAFIb IAIa cba ff / / 22 21 0)( 2 2 2 (11) Solving Eq. (11) for f , R will be found from Eq. (9). To completely address the direct kinematics problem, TB P has to be calculated which would be done in the remainder of this section. left B X )( 6 is equal to: left B X )( 6 = P AB AB AB \u2032\u2032 = \u2032\u2032 \u2032\u2032 (12) In which AB \u2032\u2032 is: \u2212 + +\u2212 = \u2212 \u2212 \u2212 =\u2032\u2032 \u2032\u2032 \u2032\u2032 \u2032\u2032 RRff RRRfff RRRBfff BA BA BA SS CSCS CCCC zz yy xx AB 22 2121 2121 \u03b8\u03b8 \u03b8\u03b8\u03b8\u03b8 \u03b8\u03b8\u03b8\u03b8 (13) expressing left B X )( 6 in the coordinate (3) (See Fig. 2): =leftX )( 6 3 RB 3 left B X )( 6 (14) Also RRRR 5 6 4 5 3 4 3 6 = is: \u2212 \u2212 == fffff ff fffff leftleft SSCCC CS CSSCS RRRR 12121 22 12121 5 6 4 5 3 4 3 6 0)( \u03c6\u03c6\u03c6\u03c6\u03c6 \u03c6\u03c6 \u03c6\u03c6\u03c6\u03c6\u03c6 (15) Equaling the first column of Eq. (15) to the Eq. (14), (since both represent leftX )( 6 3 ) the desired ff 21 ,\u03c6\u03c6 could be determined as below: = = 3 2 1 n n n n knownCompletely Bleft B left RXX 3 66 3 )()( = = \u2212 ff f ff CC S CS 21 2 21 \u03c6\u03c6 \u03c6 \u03c6\u03c6 (16) So: =f1\u03c6 atan2 ( ), 31 nn \u2212 (17) and f2\u03c6 is: =f2\u03c6 atan2 ( )/, 132 fcnn \u03c6\u2212 if \u03c0\u03c6 orf 01 = (18) =f2\u03c6 atan2 ( )/, 112 fSnn \u03c6 others (19) Now TB P can be computed as below: TB P = TTTTTT P B 65 6 4 5 3 4 0 30 (20) )(: ),(:, )(: )(: )(: 6 21 5 6 4 5 3 4 0 3 0 knownisionconfiguratthebecauseknownT dcalcualetearebacuaseknownTT calcualtedisbecauseknownT knownareinputsbecauseknownT knownisionconfiguratthebecauseKnownT P ff f B \u03c6\u03c6 Finally, since Eq", + " (17) gives two possible answers for f1\u03c6 , for each pair of ( f , R ) there are two possible answers for the forward pose solution. Moreover, Eq. (11) can lead to two answers for f thus two pairs of ( f , R ) can be acceptable and as a result there exit four possible solutions for the forward pose problem. VI. INVERSE RATE KINEMATICS In inverse rate Kinematics, having the linear velocity of point G shown in Fig. 4 as well as the angular velocity of the moving platform, all \u03b8 have to be evaluated. Consider the following definitions: GARqzlSzz /3 ,\u02c6,\u02c6\u02c6 \u2032=== where, 3z\u0302 is shown in Fig. 2, l is the length of the link and GAR /\u2032 is the position vector of point A\u2032w.r.t point G . The velocity of point A\u2032 in Fig. 4 is: qtS P \u00d7+= \u03c9 (21) where t and P\u03c9 are the linear velocity of point G and angular velocity of the moving platform, respectively and are known. Also S from the leg view is: SslS L Z \u00d7++= \u03c9 \u03c9\u03c9 )(\u02c6 (22a) where L\u03c9 is the angular velocity of the leg, Z\u03c9 is the component of L\u03c9 along 3z and\u03c9 is composed of the components of L\u03c9 which are in the plane that has 3z as the normal vector" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000999_bf02875875-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000999_bf02875875-Figure6-1.png", + "caption": "Figure 6. A differential element of the padder roller system.", + "texts": [ + " Since the external force borne by the roller system is on a symmetrical plane, the force on the upper roller can be turned into an evenly distributed force per unit length, Q, on the lower roller. 2. The bearings supporting the roller are placed at the end points of cylinder pipe and the pressing pressure can be considered as the simply supported ends. 3. Because the length is much larger than the cross-section area, the Euler-Bernoulli beam modal is applied [1]. Figure 5 represents the arbitrary configuration of a roller operational system, where XYZ is the inertia reference frame. Figure 6 is a differential element of the roller corresponding to moving coordinates x, y, z. The x, y, z, are parallel to the inertial reference frame, and G is the center of the mass. When roller is rotating by a angular velocity \u2126 in z direction, the displacement in the Y direction is (1) From Euler\u2019s beam theory, the flexural angle can be expressed as (2) and in the Y direction, there is a angular velocity. The resultant angular velocity of the element can be expressed as (3) where i, j, k, are the unit vectors in the x, y, z axis respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002581_icees51510.2021.9383730-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002581_icees51510.2021.9383730-Figure3-1.png", + "caption": "Fig 3. Simplified Motor Geometry", + "texts": [ + " of stator slots * width of tooth * net iron length/P, No. of stator slots = 36; width of tooth = 0.0078 meters, Net iron length = 0.9*L. D. Geometry Generated The motor geometry was created using ANSYS FLUENT\u2019s and also solid works. The detail of the model includes; the geometry of the front cover, external housing, geometry of stator core that includes detail of slot, windings of stator, geometry of rotor that includes the fan, the shaft, fan cover and permanent magnets. For the air inlet, the inlet slots are included in the fan cover. The fig 3 shows the simplified motor geometry. E. Stator Geometry Stator is made up of dynamo grade laminations which are of thickness 0.35 mm or 0.5 mm. For motors of larger size cores are made of segmented laminations. To give most economical balance between the costs of dies, the peripheral length usually between 0.3 m to 0.6 m is chosen for one segment, the left over amount of scrap from the lamination cuttings from steel strips and the assembly cost. In the flux paths of alternating poles, to provide an equal number of turns, the total number of segments is chosen" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001722_ifost.2008.4602951-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001722_ifost.2008.4602951-Figure2-1.png", + "caption": "Fig. 2. Hinging rod model compressed by axial force.", + "texts": [ + " The concept of a method is the selection of such values of unknown initial condition on the left end in Cauchy problem that with the certain accuracy to satisfy to known boundary conditions on the right end of a rod. Cauchy problem is solved with RungeKutta method of the fourth order. III. NUMERICAL RESULTS As an example the model of hinging rod compressed by longitudinal force is considered. It is known, that the least critical load 2 2CL EJP \u03c0= , and dependence of a square of first own frequency on compressing force P is linear [1] (fig.2). Using the developed technique, dependences of own frequencies on value of compressing force have been calculated at large displacements the loaded end face (fig.3.). On a site of change of force P from zero up to CLP a square of first own frequency behaves linearly and crosses an abscissa in a point CLP P= . Further different situations are possible. If the rod remains rectilinear and is exposed to the further compression, a square of frequency becomes negative that signals that such configuration is unstable" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002277_j.apm.2021.01.045-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002277_j.apm.2021.01.045-Figure1-1.png", + "caption": "Fig. 1. Initial structure.", + "texts": [ + " (19) , the left side of the equation is the inverse of A , and the right side is the inverse of a and the decomposition of a obtained from the initial analysis. This avoids re-inverting A. Because the finite element method is a discretized continuum for analysis, when the structure of the continuum is changed, the value of the stiffness matrix obtained by the discretization changes continuously, and the change of the adjacent element always gradually becomes smaller and does not appear. Stepped fault change, that is, there will be no change in i , and it will become 0 when i + 1 . Fig. 1 shows the initial truss structure, and Fig. 2 shows a schematic diagram of the structure after modification, in which blue represents the modified unit. Figs. 3 and 4 represent the stiffness or mass matrix, respectively. The size of the black dot represents the size of the matrix after the structure is modified, and change value is ij = | a ij \u2212 A ij |. is continuous and always increases or decreases gradually. The black dot in Fig. 4 always changes slowly from large to small. In order to determine the size of the affected area block in the modified matrix, the concept of change threshold \u03bb is introduced" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002838_012021-Figure7-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002838_012021-Figure7-1.png", + "caption": "Figure 7. Stress Distribution in Al 6061 T6", + "texts": [ + " Analysing Testing Result of Al 6061 T6 3.1.1. Total Deformation The Max. And Min. Total Deformation in Al 6061 T6 is 0.2684 mm and 0 mm respectively shown in Figure 6.The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.1.2. Stress Distribution The Max. and Min. Stress Distribution in Al 6061 T6 is 63.7 MPa and 0.3898 MPa respectively shown in Figure 7. 3.1.3. Strain Distribution The Max. and Min. Strain Distribution in Al 6061 T6 is 0.00096883 and 0.0000076477 respectively shown in Figure 8. 3.2. Analysing Testing Result of Structural Steel 3.2.1. Total Deformation The Max. and Min. Total Deformation in Structural Steel is 0.18644 mm and 0 mm respectively shown in Figure 9. The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002016_mmvip.2007.4430716-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002016_mmvip.2007.4430716-Figure3-1.png", + "caption": "Fig. 3. Ultra-precision states driven by linear motor (SPIDER).", + "texts": [ + " Step 3: By converting the discrete-time model to the continuous time model, the pole and gain of the plant model are derived. Step 4: The DIMC controller is recursively realized using identified pole and gain. In step 4, the DIMC controller is calculated as a function of the pole and gain beforehand. After step 4, it returns to step 1, and the procedure is repeated. IV. INVESTIGATED LINEAR ACTUATOR-DRIVEN PRECISION STAGE The controlled object is a 100 mm-stroke precision stage driven by the linear actuator-type synchronous piezoelectric devise driver (SPIDER) [8]. Fig. 3 shows SPIDER, the actuator properties, the drive sequence and the SPIDERdriven stage. SPIDER has eight legs. Each leg is structured by eight-stacked piezoelectric plates, the upper four plates are for the expand motion and the bottom four plates are for the shear motion. The friction tip of the SPIDER contacts the guide plate installed at the side of the stage. By applying the sinusoidal input voltage to the both motions\u2019 piezoelectric plates, it is possible to driven the elliptical orbit in the friction tip" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001825_iros.2008.4651119-Figure2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001825_iros.2008.4651119-Figure2-1.png", + "caption": "Fig. 2. Target System (N = 2)", + "texts": [ + " [10] discussed a dynamic manipulability [1]. However, these analyses are not in velocity domain but in acceleration domain. Since power is not included/considered, the obtained generable accelerations are not associated with the generable velocities (obtained in this paper). They discussed only fingertip grasp. For the complexity, the computation load is very large. The other information such as inertia tensor is needed. In addition, in [10], internal force is fixed and then applicable class is limited. The target system is shown in Fig.2. In this paper, we consider a general grasping system where an arbitrary shaped rigid object is grasped by N fingers of a robotic hand. The nomenclatures are listed at appendix. We define that the contact state is any of the following four states: 1) F-point : the contact point with static friction, 2) N-point : the contact point without friction, 3) S-point : the contact point with kinetic friction, 4) D-point : the point about to detach. We note that the contact force is zero at D-points. Also, we assume that every contact position, every contact state, every frictional coefficient are all given" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002993_s00170-021-07405-8-Figure6-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002993_s00170-021-07405-8-Figure6-1.png", + "caption": "Fig. 6 Schematic diagram of precision sizing method", + "texts": [ + " For the accuracy of external gear, with the rise of tooth width, the tooth thickness of external gear decreased firstly, and then grew gradually. Besides, the profile accuracy of external gear was ninth class and tooth lead accuracy was tenth class. The measurement results indicate that the tooth accuracy of extruded sun gear is poor and fails tomeet the final product requirements. Therefore, it is necessary to add the precision sizing process to improve the teeth accuracy of extruded sun gear. As illustrated in Fig. 6, the precision sizing method was designed according to the teeth accuracy requirements of target sun gear, where \u25b3L1 represents the sizing amount of external gear, and \u25b3L2 is the interference value of internal spline. The internal-external teeth will be reshaped by the whole tooth sizing method. This means that the tooth top, tooth flank, and tooth root will be all finished. The principle of precision sizing method is that the material radial flow is promoted by the finishing of external gear, resulting in the internal tooth fitting the splined mandrel more closely" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002776_03093247211018820-Figure3-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002776_03093247211018820-Figure3-1.png", + "caption": "Figure 3. Mesh model and constraint conditions.", + "texts": [ + " Meshing equation was established and the solving results are shown in Figure 2 based on the parameters in Tables 1 and 2. The contact ellipse major axis is considered when the gears are under a small load, and the elastic deformation is 0.00635mm. The contact path is calculated and the relative transmission error curve is calculated without elastic deformation. The transmission error unit is arc-second and unit of rotational angle is degree. According to TCA results and modeling method, the mesh model of spiral bevel gear pair could be obtained meeting assembly conditions, as shown in Figure 3. After installation, the gear and pinion both take the right meshing position. Nodes inside body have been imposed restrictions (as highlighted in Figure 5): both pinion and gear release the rotational DOF of rotational axis while rotational speed and load torque are applied on pinion and gear respectively. These models are calculated with the following material properties: elastic modulus E=2:063105 MPa, Poisson\u2019s ratio m=0:29, and applied torques on gear are 1500Nm. Since these analyses are carried out under static condition, the friction between teeth surfaces had a less effect and was ignored. Verification of the proposed method with low-speed condition After installation, nodes inside body have been imposed restrictions (as highlighted in Figure 3): both pinion and gear release the rotational DOF of rotational axis while rotational speed and load torque are applied on pinion and gear respectively. Pinion rotational speed is 500 r/min for low-speed dynamic analysis. SBG shows different dynamic characteristic under different working conditions. Since it is difficult to conduct high-speed simulation for FEM, this paper presented two different working conditions to verify the accuracy and efficiency of proposed method: low-speed and high-speed conditions with the same load torque" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002838_012021-Figure27-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002838_012021-Figure27-1.png", + "caption": "Figure 27. Total Deformation in HSLA Steel", + "texts": [ + " Stress Distribution in Carbon Fiber is 97.378 MPa and 0.28227 MPa respectively shown in Figure 25. 3.7.3. Strain Distribution The Max. And Min. Strain Distribution in Carbon Fiber is 0.0024296 and 0.000021423 respectively shown in Figure 26. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 3.8. Analysing Testing Result of HSLA Steel 3.8.1. Total Deformation The Max. And Min. Total Deformation in HSLA Steel is 0.19371 mm and 0 mm respectively shown in Figure 27. The Minimum deformation is 0 mm because there is no deformation occur at the hub part of the rim. 3.8.2. Stress Distribution The Max. And Min. Stress Distribution in HSLA Steel is 181.96 MPa and 1.2159 MPa respectively shown in Figure 28. 3.8.3. Strain Distribution The Max. And Min. Strain Distribution in HSLA Steel is 0.0011085 and 0.00000815 respectively shown in Figure 29. FSAET 2020 IOP Conf. Series: Materials Science and Engineering 1116 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1116/1/012021 The Analysis report of maximum total deformation, maximum equivalent stress and maximum equivalent strain is presented below in bar graphs: 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0003436_j.apacoust.2021.108345-Figure9-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0003436_j.apacoust.2021.108345-Figure9-1.png", + "caption": "Fig. 9. Sketch of the possible", + "texts": [ + "1 dB(A) for the simulated signal. The prediction model allows studying the sound signal generated by pairs of spur gears. In this section, the prediction model is used to study the signal produced by a specific example of a proposed MAVAS. This proposal is designed to have a mechanical assembly as simple as possible. Three pairs of spur gears without lubrication are used to produce a more complex sound, all of them with the same centre separation to facilitate the arrangement of the system. As can be seen in Fig. 9 the gear driving wheels are mounted on the same driving shaft, while the gear driven wheels are mounted on bearings over a fixed axis. The gears are made of steel and have a thickness of 16 mm. The module and number of teeth of each pair can be consulted in Table 2. Bearings allow each driven wheel to rotate at its speed without load. The proposed system can be installed in a box, which has the purpose of increasing the directivity of the sound emission in addition to allowing the positioning of the axes as shown in Fig. 9. However, the acoustic behaviour of the box has not been studied yet. According to regulation [11], the speed range for the AVAS operation is the range of greater than 0 up to and inclusive to 20 km/h. For this reason, it is considered that the MAVAS drive system provides a rotation speed to the driving shaft from 0 to 120 rpm proportionally to the vehicle speed when it circulates from 0 to 20 km/ h, as it could be assumed the same angular speed of a wheel shaft of a commercial vehicle. In addition, the regulation includes third-octave band requirements for test speeds of 10 and 20 km/h, therefore the sound produced by the three gear pairs at 60 and 120 rpm has been simulated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002772_978-3-030-67750-3_14-Figure28.2-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002772_978-3-030-67750-3_14-Figure28.2-1.png", + "caption": "Fig. 28.2 (a) and (b) Mode 1 Fig. 28.2(a) and Mode 2 Fig. 28.2(b) working principles", + "texts": [ + " The fire extinguisher nozzle experiment was conducted in a closed huge environment to ensure the data obtained covered the whole space and avoid wind, temperature and pressure disturbance. The large space allows the researchers to closely monitor the dust particles coverage. It also allows the researchers to measure the distance should be because the particles can be seen and covered the large area. The safety and health issues is also important because the fine dry powder particles can be hazardous to human respiratory system. These factors can give a significant impact on the experiment result. The bi-nozzle system operates in dual mode. Mode one, as shown in Fig. 28.2(a), works with only one outlet, which is the center outlet with 12 mm of diameter. The working principle behind this mode is when the nozzle is fully closed, the pathway for the dry powder to travel through is only limited to one. The aim is to create a high pressure and velocity stream of the fine particles flow out through the outlet, thus creating a narrow spray distance. As for mode two, as presented in Fig. 28.2. Figure 28.2(b), the dry powder particles for this mode flow two direction through outlet 1 and outlet 2. For the current nozzle, outlet 2 has a bigger diameter with 52 mm. The working principle of these outlets requires the nozzle to turn counter clockwise to allow the particles travel through two pathways. This is to allow the fluid to pass through a more significant diameter outlet (outlet 2) that create a wider spray angle with a slow velocity. Figure 28.3 illustrate the evolution of the bi-nozzle design to the existing one" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0002262_tia.2021.3058113-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0002262_tia.2021.3058113-Figure1-1.png", + "caption": "Fig. 1. Three-pole AMB and Wye coil connection used in this paper.", + "texts": [ + " The wide variation of three-pole design topologies was reviewed in [5] which concludes that designs can be categorized by whether a bias field is used and whether zero-sequence currents are used in the control coils. Options for the bias field include heteropolar airgap fields, i.e. [6]\u2013[10] and homopolar airgap fields, i.e. [17]\u2013[20] (several of which are also able to create axial forces) as well as no bias field, i.e. [14]\u2013[16]. The operating principles of all three-pole bearing topologies can be derived from Fig. 1, which has three electromagnetic actuators separated spatially by 120 degrees that can each impart a force on the rotor. Three-pole bearings that use an external bias field suffer from large force vector error (often referred to as \u201cinterference forces\u201d) [9], [10], [21], [22] which reduces the bearing\u2019s stability [23]. In previous work [5], the authors have shown that 1) the force density of a three-pole bearing can be increased by 15% by using an optimal external bias field; 2) it is possible to eliminate the force vector error from externallybiased bearings by determining the necessary coil currents from a more detailed force-current model", + " Finally, the proposed regulator is investigated and validated through a simulation study and an experimental prototype of a combined radial axial magnetic bearing (CRAMB). Authorized licensed use limited to: Carleton University. Downloaded on June 04,2021 at 02:56:52 UTC from IEEE Xplore. Restrictions apply. 0093-9994 (c) 2021 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. This paper assumes the three pole bearing structure depicted in Fig. 1 is operated by the conventional three phase inverter typical of electric motor drives. The coils are connected in a Wye configuration, requiring that their currents sum to zero. In addition to airgap \u201ccontrol\u201d fields created by the coil currents, this paper assumes that an external flux source provides a bias field to the airgap. An example configuration is shown in Fig. 2, which depicts a combined radial-axial magnetic bearing (CRAMB) utilizing axially-magnetized permanent magnets to generate the bias flux. The \u201cradial stage\u201d is implemented as the three-pole bearing of Fig. 1. This particular topology is discussed at length in [24]. Other common geometry configurations to realize this are reviewed in a previous work by the authors [5], which also developed a generalized force model for the three-pole bearing and proposed an \u201cexact solution\u201d to calculate the currents necessary to create a desired force vector. These results form the basis of the present paper and are now summarized in this section. Radial forces produced on the bearing\u2019s shaft can be modeled by (1). Here, F\u2032 is a complex force vector normalized by the maximum force Fmax = k1B 2 max that each bearing pole can produce, Bc,i is the control component of the airgap field in front of each pole (proportional to the pole coil\u2019s control current ic,i), Bmax is the maximum allowable airgap field, and SV(~v) is a space-vector operator that operates on a cartesian vector ~v as SV(~v) = [1, a, a2]~v, where a = e j2\u03c0 3 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001087_ijtc2007-44351-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001087_ijtc2007-44351-Figure1-1.png", + "caption": "Fig. 1 Brush seal geometry.", + "texts": [ + " This work presents a summary of a 3-D computational brush seal tip force and wear analysis. The analysis models a representative brush segment with bristles formed by 3-D beam elements. Bristle interlocking and frictional interactions (interbristle, bristle-backing plate and bristle-rotor) are included to better calculate resulting seal stiffness and tip forces. Results are compared to stiffness measurements and full scale seal wear tests. The brush seal consists of a set of fine diameter fibers densely packed between retaining and backing plates. As illustrated in Figure 1, the backing plate is positioned downstream of the bristles to provide mechanical support under differential pressure loads. The bristles touch the rotor with a lay angle in the direction of the rotor rotation allowing them to bend rather than buckle during rotor excursions. Last few decades, brush seals have been extensively used in secondary flow sealing in turbo-machinery applications. They have demonstrated excellent leakage characteristics. Differential pressure across a seal pushes the bristles against the backing plate" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0001585_icelmach.2008.4799976-Figure1-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0001585_icelmach.2008.4799976-Figure1-1.png", + "caption": "Fig. 1. Finite element mesh of the induction motor.", + "texts": [ + " External circuit equations are derived that allow having voltage fed electromechanical devices, and these extra equations are solved simultaneously with the field equations. The resulting set of equations is: \u23a5 \u23a6 \u23a4 \u23a2 \u23a3 \u23a1 =\u23a5 \u23a6 \u23a4 \u23a2 \u23a3 \u23a1 \u23a5 \u23a6 \u23a4 \u23a2 \u23a3 \u23a1 + \u2212+ ew VI A LjRQj PNjS 0 \u03c9\u03c9 \u03c9 (2) where S is the FE stiffness matrix, P and Q are weighting winding matrices, R is the external resistance, L external inductance, Iw a vector of winding currents, Ve denotes the external potential sources. The FE model mesh, solution and post-processing was carried out using a commercial software [7]. Fig. 1 illustrates the resulting FE mesh where it can clearly be seen that periodic boundary conditions were employed to reduce the number of unknowns and consequently the computational burden. The skew effect of the motor is not considered, but this can be easily taken into account using a multisliced FE model [7]. Fig. 2 shows the electrical circuit connected to an external voltage. The circuit can be straightforward built using a friendly FE software [7]. In Fig. 2, V1, V2 and V3 are the voltage sources, R1-3 denote the end-winding stator resistances, L1-3 are the end-winding inductances for each phase" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_83_0000171_detc2004-57029-Figure9-1.png", + "original_path": "designv11-83/openalex_figure/designv11_83_0000171_detc2004-57029-Figure9-1.png", + "caption": "Fig. 9 Workspace in three dimensions( ) o25=\u03c9", + "texts": [ + " So is the rcond2. When both rcond1 and rcond2 are 1, the dexterity of the PMT is best. 3. Examples The main structural parameters of the prototype of PMT developed in our laboratory are as follows rA=720mm \uff0c W \uff1d 780 mm \uff0c 2/\u03c0\u03b8 = \uff0c mm\uff0c200=Br 5/2\u03c0\u03c6 = \uff0ca4=180mm 3.1 The workspace of the prototype Based on the given structural parameters, through numerical calculation, some graphical representations of the workspace are obtained in three dimensions. The workspace graphs in line with and are shown in Fig.8 and Fig.9. It is obvious that the bigger the agile cutter sloping angle is the smaller the workspace. The workspace corresponding to the bigger cutter sloping angle is contained by the workspace corresponding to the smaller one. o0=\u03c9 o25=\u03c9 3.2 Dexterity analysis of the prototype The reciprocals of condition numbers on the section and pose 1000=x 0== \u03b2\u03b1 are shown in Fig.10. The reciprocals of condition numbers on nloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/09/2016 Te the section 1000=x and pose are shown in Fig" + ], + "surrounding_texts": [] + } +] \ No newline at end of file