diff --git "a/designv11-33.json" "b/designv11-33.json" new file mode 100644--- /dev/null +++ "b/designv11-33.json" @@ -0,0 +1,10421 @@ +[ + { + "image_filename": "designv11_33_0003629_b978-0-12-816713-7.00021-0-Figure21.5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003629_b978-0-12-816713-7.00021-0-Figure21.5-1.png", + "caption": "FIGURE 21.5 Definition of joint between DIP and PIP bone. DIP, distal phalange; PIP, proximal phalange.", + "texts": [ + " First, the center of the rotation in proximal phalange bone was identified where a new local coordinate system was created. The coordinate system was then oriented in such manner that the \u201cy\u201d axis corresponds with the axis of the rotation of the joint axis. In the center of the new local coordinate system, a new reference point (RP-PIP) was created. Another reference point (RP-DIP) was also created on the surface of the distal phalange bone. Between both reference points, a rigid wire was then created, which is used to define a connector that allows rotation only in \u201cy\u201d axis at the RP-PIP point (Fig 21.5). Both reference points are connected to bones using constraints, which fix the translations and rotations of the wire relative to the bone. In this manner, a simplified, numerically stable, and biomechanically accurate joint can be defined. Other joints in the human hand are then defined analogously, considering their local features. The accuracy of joint movement and soft-tissue deformation was validated on a simplified fingertip FE model. Results showed accurate joint movement and soft-tissue deformation (Fig 21" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000777_s0025654411040017-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000777_s0025654411040017-Figure2-1.png", + "caption": "Fig. 2. Fig. 3.", + "texts": [], + "surrounding_texts": [ + "By setting \u03b4 = 0 in Eqs. (3.6) and (3.7) and by introducing new variables z1 and z2 by formulas z1 = z + a1\u03d5 and z2 = z \u2212 a2\u03d5, we obtain the equations z\u2022\u20221 + a11(c1z1 + \u03b31z \u2022 1) + a12(c2z2 + \u03b32z \u2022 2) = 0, z\u2022\u20222 + a21(c1z1 + \u03b31z \u2022 1) + a22(c2z2 + \u03b32z \u2022 2) = 0, (4.1) MECHANICS OF SOLIDS Vol. 46 No. 4 2011 a11 = 1 + a2 1 \u2212 a1f1, a12 = 1 \u2212 a1a2 \u2212 a1f2, a21 = 1 \u2212 a1a2 + a2f1, a22 = 1 + a2 2 + a2f2. (4.2) The dimensionless variables ak and fk (k = 1, 2) in formulas (4.2) are the ratios of the corresponding variables introduced in the preceding sections to the radius of inertia \u03c1; i.e., we assume that the radius of inertia \u03c1 is equal to unity. The variables z1 and z2 have the following physical meaning. These are small deviations of the points O1 and O2 from their equilibrium positions described in Sec. 3. Equations (4.1) and (4.2) and relations (1.1) with the use of the Lyapunov theorem permit solving the above stability and instability problem for any fixed v. Let us write out the characteristic polynomial of the linear system (4.1) as p4 + p3(a11\u03b31 + a22\u03b32) + p2(a11c1 + a22c2 + \u03b31\u03b32\u0394) + p(\u03b31c2 + \u03b32c1)\u0394 + c1c2\u0394 = 0, (4.3) where \u0394 = a11a22 \u2212 a12a21. Applying the Routh\u2013Hurwitz theorem to the polynomial (4.3), we arrive at the following result. Proposition 1. For the asymptotic stability of the equilibrium in the case of rectilinear and uniform motion of the vehicle, it is necessary and sufficient that the rolling friction coefficients f1 and f2 (and hence the velocity v of the vehicle translational motion determined by (1.1)) satisfy the inequalities a11\u03b31 + a22\u03b32 > 0, a11c1 + a22c2 + \u03b31\u03b32\u0394 > 0, \u0394 > 0, (4.4) (a11\u03b31 + a22\u03b32)[a11\u03b32c 2 1 + a22\u03b31c 2 2 + \u03b31\u03b32\u0394(\u03b31c2 + \u03b32c1)] \u2212 (\u03b31c2 + \u03b32c1)2\u0394 > 0, \u0394 = a11a22 \u2212 a12a21 = a(a + f2 \u2212 f1), a = a1 + a2. (4.5) The dimensionless parameters aij (i, j = 1, 2) are determined by formulas (4.2). Consider several particular cases where Proposition 1 applies. Proposition 2. If the system is completely symmetric, i.e., if \u03b3k = \u03b3 = 0, ck = c, fk = f, k = 1, 2, a1 = a2 = a0, (4.6) then inequalities (4.4) hold for each f ; i.e., the asymptotic stability of the vehicle equilibrium considered above is preserved for any wheel rolling friction coefficients and hence for any velocities of its translational motion. The proof is by a straightforward verification of inequalities (4.4) under conditions (4.6). In this case, it follows from (4.5) that \u0394 = a2 > 0 and the first two inequalities in (4.4) are satisfied, because they are reduced to the following ones: 2\u03b3(1 + a2 0) > 0, 2c(1 + a2 0) + \u03b32a2 > 0. The last inequality in (4.4) has the form c2(1 \u2212 a2 0) 2 + 4\u03b32c2a2 0(1 + a2 0) > 0. Thus, the stability conditions (4.4) are satisfied for any f \u2265 0, and the proof of Proposition 2 is complete. Proposition 3. Assume that all conditions (4.6) except for the last are satisfied; i.e., a1 =a2 (the case of incomplete symmetry; i.e., the car body center of mass is displaced). Then, for a1 < a2, the stability conditions (4.4) are satisfied for every f \u2265 0, and for a1 > a2, the stability is violated for f > f0 = 1 a1 \u2212 a2 [ 2 + a2 1 + a2 2 + \u03b5(a1 + a2)2 \u2212 \u221a \u03b52(a1 + a2)4 + 4(a1 + a2)2 ] > 0, \u03b5 = \u03b32 c . (4.7) Remark. The quantity f0 in inequality (4.7) can be arbitrarily small for some values of the system parameters. For example, as \u03b5 \u2192 0 (small damping or large rigidity of the wheel) and for a1a2 = 1 and a1 > 0, we have f0 = (a1 \u2212 1)(1 + a2 1)/(a1 + a2 1), which can be arbitrarily small as a1 \u2192 1. MECHANICS OF SOLIDS Vol. 46 No. 4 2011 The proof is by a straightforward verification. Under the conditions given in Proposition 3, the instability inequalities (4.4) become \u03b3[2 + a2 1 + a2 2 + (a2 \u2212 a1)f ] > 0, c[2 + a2 1 + a2 2 + (a2 \u2212 a1)f ] + \u03b32(a1 + a2)2 > 0, \u0394 = (a1 + a2)2 > 0, x2 + 2x\u03b5(a1 + a2)2 \u2212 4(a1 + a2)2 > 0, x = 2 + a2 1 + a2 2 + (a2 \u2212 a1)f. (4.8) For a2 > a1, these inequalities are necessarily satisfied for any f > 0. (For the first three inequalities in (4.8), this is obvious, and for the last, this follows from the inequality x > 2 + a2 1 + a2 2 > 2(a1 + a2).) But if a2 < a1, then the first inequality in (4.8) is violated for f > f1 = (2 + a2 1 + a2 2)/(a1 \u2212 a2), the second is violated for f > f2 > f1, the third is always satisfied, and the last is violated for f \u2208 [f0, f00], where f0 is given by the formula in (4.7) and f00 is given by the same formula with the plus sign at the radical. It is easily seen that 0 < f0 < f1 < f00, which implies that all inequalities in (4.8) are violated for f > f0. The proof of Proposition 3 is complete. Proposition 4. Let f2 = 0, \u03b30 = 0, and f1 = f > 0; i.e., assume that the fore (driving) wheel does not experience any rolling resistance and there is no dissipation in it. In this case, if a1a2 < 1, then the instability arises for f > a1, and if a1a2 > 1, then the instability arises for f > (a1a2 \u2212 1)/a2. Proposition 5. Let f1 = 0, \u03b31 = 0, and f2 = f > 0; i.e., assume that the rear (driven) wheel does not experience any rolling resistance and there is no dissipation in it. In this case, if a1a2 < 1, then the instability arises for f > (1 \u2212 a1a2)/a1, and if a1a2 > 1, then the system is stable for any f > 0. The proof of Propositions 4 and 5 is by a straightforward verification of inequalities (4.4) in Proposition 1 under the assumptions stated above. Let \u03b31 > 0, \u03b32 = 0, and \u03b32 > 0. Then the following assertion holds. Proposition 6. Under the above conditions, the asymptotic stability domain D1 in the plane of the parameters (f1, f2) is given by the following inequalities. If 0 < a1a2 < 1, then 0 < f2 < 1 \u2212 a1a2 a1 , 0 < f1, c1a1f1 \u2212 c2a2f1 < c0, f1 \u2212 f2 < a1 + a2. (4.9) If 1 < a1a2 < \u221e, then 0 < f1 < a1a2 \u2212 1 a2 , 0 < f2, c1a1f1 \u2212 c2a2f2 < c0, (4.10) c0 = c1(1 + a2 1) + c2(1 + a2 2). (4.11) The proof of Proposition 6 is by a straightforward verification of inequalities (4.4) under the restrictions given in the statement. Note that if one of the inequalities in conditions (4.10), (4.11) is not satisfied, then the vibrations under study are unstable. Now let \u03b31 = \u03b32 = 0; i.e., assume that there is no dissipation in the wheels. In this case, Eq. (4.3) is biquadratic, the stability of system (4.1) can be only nonasymptotic, and the conditions for such stability are equivalent to the conditions that this biquadratic equation has only pure imaginary roots. These considerations lead to the following result. Proposition 7. For \u03b31 = \u03b32 = 0, the stability domain D0 in the plane of parameters (f1, f2) is given by the following inequalities: f1 > 0, f2 > 0, c0 > c1a1f1 \u2212 c2a2f2, a1 + a2 > f1 \u2212 f2, [c0 \u2212 (c1a1f1 \u2212 c2a2f2)]2 > 4c1c2(a1 + a2)(a1 + a2 \u2212 f1 + f2). (4.12) The proof of Proposition 7 follows from well-known necessary and sufficient conditions for the existence of pure imaginary roots of a biquadratic equation. Proposition 8. The inclusion D1 \u2282 D0 holds. MECHANICS OF SOLIDS Vol. 46 No. 4 2011 The proof of Proposition 8 is by a straightforward substitution of the extreme values in the linear inequalities (4.9) or (4.10) into inequalities (4.12) and verification that the latter are satisfied. Remark. Proposition 8 means that this nonsymmetric addition of dissipation narrows down the stability domain in the plane of the parameters (f1, f2). Graphically (schematically), the domains D1 and D0 in various possible cases are shown in Figs. 2\u201315, where digit 1 corresponds to the line f1 \u2212 f2 = a1 + a2, digit 2 corresponds to the line c1a1f1 \u2212 c2a2f2 = c0 = c1(1 + a2 1) + c2(1 + a2 2), digit 3 corresponds to the line f2 = |1 \u2212 a1a2|/a1, digit 4 corresponds to the line f1 = |1 \u2212 a1a2|/a2, digit 5 corresponds to the line f1 \u2212 f2 = \u2212|1 \u2212 a1a2|(a1 + a2)/(a1a2), and digit 6 corresponds to the parabola (c0 \u2212 c1a1f1 + c2a2f2)2 \u2212 4c1c2(a1 + a2)(a1 + a2 \u2212 f1 + f2) = 0. Figures 2, 4, 6, 8, 10, 12, and 14 correspond to the case of \u03b31 = \u03b32 = 0, while Figs. 3, 5, 7, 9, 11, 13, and 15 correspond to the case of \u03b31 = 0, \u03b32 = 0 or \u03b31 = 0, \u03b32 = 0 for all possible distinct values of the parameters a1, a2, c1, and c2. Now consider the case of weighted wheels, m1 = 0 and m2 = 0. Since we assume that the car body mass is m = 1, the quantities m1 and m2 are the respective dimensionless ratios of the wheel masses to the car body mass. Simple calculations show that the form of the vibration equations (4.1) remains the same and the coefficients aij (i, j = 1, 2) are determined by the functions a11 = \u03c12 \u2212 (a1 + am2)(\u2212a1 + f1) \u03c12 0 , a12 = \u03c12 \u2212 (a1 + am2)(a2 + f2) \u03c12 0 , a21 = \u03c12 + (a2 + am1)(\u2212a1 + f1) \u03c12 0 , a22 = \u03c12 + (a2 + am1)(a2 + f2) \u03c12 0 , where \u03c12 0 = m0\u03c1 2 and m0 = 1 + m1 + m2. Then one can readily compute that \u0394 = a11a22 \u2212 a12a21 = a(a + f2 \u2212 f1)/\u03c12 0. Proposition 1 remains valid under the above variations in aij , i, j = 1, 2, and \u0394. In a similar way, one can restate Propositions 2\u20138. Equations (4.1) and (4.2) describing the vehicle vibrations under the action of the rolling friction forces admit the following analogies (interpretations) to well-known systems in mechanics. Each of the coefficient matrices multiplying zk and z\u2022k (k = 1, 2) in (4.1) with (4.2) taken into account consists of MECHANICS OF SOLIDS Vol. 46 No. 4 2011 two parts. The first part (it is said to be unperturbed) does not contain the friction coefficients f1 and f2. The second (perturbed) part depends on these coefficients linearly. As follows from the above derivation of the equations of motion, the origin of the perturbed part significantly depends on the following two external causes. One of them is the interaction between the deformable wheels and the rough road, which causes the rolling resistance torques Mk1 and Mk2; the second cause is the torque M1 from the engine (external energy source) to the driving wheels ensuring their (and the driven wheels) permanent rolling if there are resistance torques mentioned above. If these causes are absent, then the system motion occurs for f1 = f2 = 0, and Eqs. (4.1) contain only the first parts of the above-mentioned matrices for which the solutions have the property of (asymptotic) stability (a conservative system MECHANICS OF SOLIDS Vol. 46 No. 4 2011 under the action of some forces with complete or incomplete dissipation). But if f2 1 + f2 2 = 0, then these causes result in the appearance of additional terms in Eqs. (4.1), which, according to [6], can be interpreted as positional (circular) forces (for the corresponding terms with z1 and z2); gyroscopic and nonconservative (depending on the velocities) forces (for the corresponding terms with z\u20221 and z\u20222); conservative positional forces (for the corresponding terms with z1 and z2); and dissipative (Rayleigh) MECHANICS OF SOLIDS Vol. 46 No. 4 2011 MECHANICS OF SOLIDS Vol. 46 No. 4 2011 forces (for the corresponding terms with z\u20221 and z\u20222). In the general case, such a decomposition of forces was performed in [7]. Thus, the existence of the above-mentioned external causes leads to the appearance of additional forces of the described structure, whose destructive character is well known in mechanics of gyroscopic systems, structural mechanics (in particular, in mechanics of bridges), and other adjacent fields; it is described in the literature [8\u201310]. Even if there forces do not actually lead to the vehicle instability, they still significantly change the frequencies of its vibrations in the vertical plane, and this fact cannot be ignored when solving the corresponding problems. These facts were already considered earlier by the authors in the paper [11]." + ] + }, + { + "image_filename": "designv11_33_0002063_s00170-017-0656-8-Figure13-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002063_s00170-017-0656-8-Figure13-1.png", + "caption": "Fig. 13 The multitooth dynamic cutting model", + "texts": [ + " Suppose the resultant force F3 from the shearing plane on the chip in coordinate system S3 is [\u2212F3x 0 F3z] T, the resultant force F3 in coordinate system S2 can be written as F2 \u00bc M 23 \u03b7S\u00f0 \u00deF3 \u00f031\u00de Suppose the resultant forceF5 from the rake face of the tool on the chip in coordinate systemS5 is F5x 0 \u2212F5z\u00bd T , the resultant force F5 in coordinate system S2 can be written as F* 2 \u00bc M 21 \u03c6r\u00f0 \u00deM14 \u03b3n\u00f0 \u00deM 45 \u03b7c\u00f0 \u00deF5 \u00f032\u00de Merchant\u2019s formula [19] is universal; the frictional angle in oblique cutting can be calculated as follows: \u03b2 \u00bc \u03c0=2\u00fe \u03b3n\u22122\u03d5r \u00f033\u00de Based on the force balance principle, the following relationship can be obtained according to formulas (28), (31), (32), and (33) F3x \u00bc Fs F3z \u00bc f \u03d5r; \u03b3n; \u03b7s; \u03b7c;\u03b2\u00f0 \u00deF3x \u00f034\u00de The resultant force from the shearing plane on the chip in coordinate system S0(X0, Y0, Z0) can be written as F0 \u00bc M 01 i\u00f0 \u00deM 12 \u03c6r\u00f0 \u00deM 23 \u03b7S\u00f0 \u00deF3 \u00f035\u00de The components of F0 in X0 , Y0 , Z0 directions are the tangential force, radial force, and axial force. Deferent from the single tooth cutting theory established earlier, the research of the cutting force in this paper is carried out based on the actual situation. Therefore, at any time within the scope of the cutting tool effective angle, there are one orFig. 12 The illustration of the shearing force and the cutting force several cutting blades at work at the same time. The multitooth dynamic cutting model is established in Fig. 13. Subscripts of i ,m , o represent the parameters related to the inner, outer, and middle cutting blades. Axial forces of FZi , FZm , FZo are perpendicular to the paper plane; the radial force of inner cutting blades is opposite to that of outer cutting blades. According to the established model in Sect. 3.1, the values of \u03b8i , \u03b8m , \u03b8o could be calculated, which are within the scope of the cutting tool effective angle at any time. The three direction forces in the coordinate system established in Fig", + " Figures 17 and 21 demonstrate that the change trend of the calculation results is the same as the simulation results, and the magnitude of forces is basically in line; the accuracy of the multitooth dynamic cutting force model is verified. In order to further verify the cutting force calculation theory and the simulated model established, the corresponding machining experiment is carried out. The experiment platform is set up as shown in Fig. 22. During processing, the experimental parameters (cutting tool rotating speed, feeding speed, etc.) are the same with the simulation parameters. The coordinate system azimuth agrees with that in Fig. 13. The cutting forces collected in three directions are shown in Fig. 23. As shown in Fig. 23, the three direction cutting forces all present an increasing trend. The Y direction cutting force changes most and the X direction cutting force changes least; the biggest cutting force is in the Y direction. Besides, the cutting force is discontinuous in each direction, because the number of the cutting blade in the tooth slot changes alternately. If it changes suddenly, revulsion of the cutting force will occur" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000704_cp.2012.0407-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000704_cp.2012.0407-Figure1-1.png", + "caption": "Figure 1: Arc segments with C2 contact intermediate vertex", + "texts": [ + " (4) To generate the road using on-board visual sensor measurements or approximate the real road from a given road-map, this study uses constant curvature segments. Assuming that some of vertices on the road can be obtained, those vertices are connected by line segments of a constant curvature with C2 contact at the vertices. In order to ensure connection between line segments, both position and tangent end point constraints should be met. By introducing intermediate vertex vi, two arcs of different curvature can connect two configurations as shown in Fig. 1. The entire road-map can then be modelled by a set of road segments ri, i = 1, . . . , n, and for each road segment, the centre position of the road curve and its curvature are given by the approximation algorithm. The mathematical details of the construction of the line curvature between vertices are described in [13]. Figure 2(a) shows a sample road-network of Devizes, Wiltshire, United Kingdom with GIS satellite data [3]. Given roads of interest as blue line in the figure, Fig. 2(b) shows the approximated road" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002260_j.ifacol.2017.08.764-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002260_j.ifacol.2017.08.764-Figure2-1.png", + "caption": "Fig. 2. Quadrotor-manipulator system and communication topology used in simulation", + "texts": [ + " Consider the system (30) with (62) in a ringstructured communication topology, \u03b1i \u2192 \u03b1d,i, \u03b1\u0307i \u2192 0, \u2200i \u2208 N as t \u2192 \u221e, if K1 \u03b1,i > 0, K2 \u03b1,i > 0 and K1 \u03b1,i \u2212 2K2 \u03b1,i > 0. Proof. The closed-loop system after applying the manipulator controller (62) to (30) can be written as follows: M33,is\u0307\u03b1,i+C33,is\u03b1,i +K1 \u03b1,is\u03b1,i \u2212K2 \u03b1,is\u03b1,i\u22121 \u2212K2 \u03b1,is\u03b1,i+1= 0. (63) The following proof is the same as in Qi et al. (2016). To illustrate the performance of the control scheme, let us consider four quadrotors equipped with a 2-DOF manipulator as shown in Fig. 2(a) hover at [\u221215, 0, 5] m, [15, 0.5, 5] m, [0.5,\u221215, 5] m and [0, 15, 5] m. The desired formation is a rhombus and the formation offsets are designed as \u03b41 = [10, 0, 0] m, \u03b42 = [\u221210, 0, 0] m, \u03b43 = [0, 10, 0] m and \u03b44 = [0,\u221210, 0] m. In this simulation, we choose R = 8 m, r = 2 m and desired angles of manipulator are \u03b1d,1 = [\u03c0, \u03c0/2] T rad, \u03b1d,2 = [\u03c0, \u03c0/2] T rad, \u03b1d,3 = [\u03c0, \u03c0/2] T rad and \u03b1d,4 = [\u03c0, \u03c0/2]T rad. The two-way ring network structure for the four quadrotor-manipulator systems is shown in Fig. 2(b). The physical parameters of the integrated system are given in Table 1. The simulation results are presented from Fig. 3 to Fig. 5. For the control gains in the equation (36), (55) and (62), we choose Kp,i = diag [5, 5, 5], K\u03c6,i = diag [10, 10, 10], \u039b\u03c6 = diag [20, 20, 20], K1 \u03b1,i = diag [5, 5], K2 \u03b1,i = diag [2, 2] Proceedings of the 20th IFAC World Congress Toulouse, France, July 9-14, 2017 7928 Yuhua Qi et al. / IFAC PapersOnLine 50-1 (2017) 7923\u20137928 and \u039b\u03b1 = diag [10, 10, 10] for each quadrotor-manipulator system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003903_s11370-019-00306-6-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003903_s11370-019-00306-6-Figure5-1.png", + "caption": "Fig. 5 Coupling schematic of the pitch and yaw joints of the microinstrument", + "texts": [ + " 3 Dynamic analysis diagram of the cable-driven joint unit reset motor traction motor z0 x0 y0 \u03b8 1 P1 displacement -P1 tension sensors 1T2 1T1 set of pilot wheels 1Ff 2 1Ff 1 1Ft2 1Ft1 \u03c4 f 1\u03c4 1 r0 initial position F1z F1x F1y g 1 3 According to the coordinate system, the transformation matrix between the base coordinate system and the tool coordinate system is obtained, and the forward kinematics are obtained by where si \u225c sin i, ci \u225c cos i , m = a, b. Let the target positions of portions a and b of the forceps be m = [ pmX pmY pmZ ]T , then the inverse kinematics are obtained by Since the pitch and the yaw axis of the microinstrument have an offset in space, there is a kinematic coupling, so decoupling control is needed. The pitch and the yaw joint axes are placed on a plane, and the coupling schematic is shown in Fig.\u00a05, where a1 is the offset distance between the yaw joint and the pitch joint, A, B and B\u02b9 are the tangent points of the cable and pitch drive wheel, C, C\u02b9, D and D\u02b9 are the tangent points of the cable and yaw drive wheel, and the effective radii of the pitch and yaw drive wheels are r0 and r1, respectively. In the initial position of the pitch joint o0o1, where \u03b81 = 0, the length of the cable is l = | | | | \u2322 AB | | | | + |BC| + | | | | \u2322 CD | | | | ; after the 0A2m = 0A1 1A2m = \u23a1 \u23a2 \u23a2 \u23a2 \u23a3 c1c2m \u2212c1s2m s1 a1c1 + a2c1c2m s1c2m \u2212s1s2m \u2212c1 a1s1 + a2s1c2m s2m c2m 0 a2s2m 0 0 0 1 \u23a4 \u23a5 \u23a5 \u23a5 \u23a6 \u23a7 \u23aa \u23a8 \u23aa \u23a9 1 = atan2 \ufffd 2a1pmY p2 mX +p2 mY +p2 mZ +a2 1 \u2212a2 2 , 2a1pmX p2 mX +p2 mY +p2 mZ +a2 1 \u2212a2 2 \ufffd , 1 \u2208 \ufffd \u22124 \u22159, 4 \u22159 \ufffd rad 2m = atan2 \ufffd pmZ a2 , p2 mX +p2 mY +p2 mZ \u2212a2 1 \u2212a2 2 2a1a2 \ufffd , 2m \u2208 \ufffd \u22124 \u22159, 4 \u22159 \ufffd rad pitch joint is rotated \u03b81, where the position is o0o2, the length of the cable is l\ufffd = | | | | \u2322 AB\ufffd | | | | + |B\ufffdC\ufffd | + | | | | \u2322 C\ufffdD\ufffd | | | | , the change in the length of the cable is \u0394l = l\u02b9\u00a0\u2212\u00a0l according to the geometric relationship, |BC| = |B\ufffdC\ufffd | , | | | | \u2322 CD | | | | = | | | | \u2322 C\ufffdD\ufffd | | | | , and \u2220Bo0B\ufffd = 1 , so the change in the length of the cable is \u0394l = | | | | \u2322 AB\ufffd | | | | \u2212 | | | | \u2322 AB | | | | = | | | | \u2322 BB\ufffd | | | | = \ud835\udf0br0\ud835\udf031 180 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001497_1059712312462904-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001497_1059712312462904-Figure1-1.png", + "caption": "Figure 1. The 2D seven-link and 3D nine-link biped models used in this study. (a) The 2D model incorporates a total of seven degrees of freedom. (b) The 3D walking model is a straightforward extension of the 2D model. In addition to the features of the 2D model, the ankles have a \u2018\u2018passively\u2019\u2019 actuated degree of freedom (simulated springs, including damping), and two links were added to allow symmetric hip adduction.", + "texts": [ + " We used anthropomorphic dynamic parameters for the 2D model (Winter, 1991) to allow direct comparison with the kinematics and actuation torques measured in human locomotion, and the 3D model\u2019s parameters were based on the Delft University of Technology\u2019s prototype \u2018\u2018Flame\u2019\u2019 to test the controller\u2019s potential performance on a plausible real-world testbed (Locascio, Solomon & Hartmann, in press). The 2D model consisted of seven links: an upper body, two thighs, two shanks, and two feet, as shown in Figure 1a. Parameter values for the 2D model are provided in (Wisse, 2008; Locascio et al. in press; Table 1), and are based on typical values for a 70-kg, 1.75-m tall human (Winter, 1991). Note that the center of mass distance for each segment is measured relative to its \u2018\u2018parent\u2019\u2019 joint (e.g., the hip joint for the thigh segment). Also, three length values are provided for the foot; these correspond to heel-to-toe length, heel-to-instep length, and instep-to-ankle length, where \u2018\u2018instep\u2019\u2019 refers to the point on the bottom of the foot directly below the ankle when the foot is flat on the ground. The foot center of mass resides at the geometric center of the foot. Geometric and inertial parameter values for the 3D model are provided in the FLAME documentation (Wisse, 2008; Locascio et al., in press). The 3D model, shown in Figure 1b, includes the same seven links as the 2D model, plus two hip links, used to introduce a pelvic width into the model and to allow independent hip adduction/abduction (hereafter simply \u2018\u2018adduction\u2019\u2019 for brevity) and hip flexion/extension. These two links are mechanically constrained (consistent with the real Flame) such that the left and right legs cannot adduct independently, but must always have the same angle relative to the torso. The imposition of this symmetry simplifies the model, while still allowing sufficient step-width control. There is an additional degree of freedom in the ankles, allowing them to \u2018\u2018roll\u2019\u2019 about the axis indicated in Figure 1b. This degree of freedom is un-actuated, but is under the influence of passive torsional springs. Thus, only a single additional actuator (hip adduction) was added to the 3D model. Both the 2D and the 3D models include encoder sensors that measure the angle and velocity of each joint, an inertial sensor to measure the upper body angle and velocity with respect to gravity, and a binary contact sensor at the heel and toe of each foot. Each of the six joints in the 2D model has a torque-controlled actuator, while the 3D model also has one additional actuator for hip abduction/adduction, and corresponding position and velocity sensors at that joint" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002078_978-3-319-63537-8_2-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002078_978-3-319-63537-8_2-Figure2-1.png", + "caption": "Fig. 2. The Basiliskbot", + "texts": [ + " Each trial consisted of observing the lizard running along a straight path without any interruptions. Three trials were completed per basilisk lizard per saturation level. 40 min in the incubator was allocated for their rest between trials. The order of experimentation with the lizards were randomized for each saturation level evaluated. The dimensions of the 3 basilisk lizards were taken from the Motive software and averaged. In particular, the length of fore limbs (3.05 \u00b1 0.25 cm), hind limbs (6.63 \u00b1 0.44 cm), and tail (17 \u00b1 2.89 cm) were measured. The basilisk-inspired robot, Basiliskbot (Fig. 2) was designed based on the averaged dimensions of the animal. The four-spoked legs were 3D printed with Acrylonitrile-Butadiene Styrene (ABS) with a 30% fill and thickness of 8.5 mm, while the body was laser cut using 4.3 mm acrylic. Mass of the robot (without the tail) was 273 g with the following components on board: four 50:1 micro DC geared motors, four magnetic encoders, four aluminum motor mounts, two dual motor drivers, a micro servo, a Teensy 3.2 micro-controller, and an inertial measurement unit (IMU)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000105_amm.37-38.1462-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000105_amm.37-38.1462-Figure3-1.png", + "caption": "Fig. 3 Scheme of cooling gallery moving along axial direction (a) and radial direction (b) of piston", + "texts": [ + " In the study, for convenience, the shape and section aera of cooling gallery are kept unchanged. The investigation is focused on the relationship between cooling gallery position and temperature as well as thermal stress. Both axial and radial direction positions are studied. For axial direction, 11 positions of cooling gallery are considered. The positions are obtained by moving cooling gallery from the position A to position B in increment of 1 mm, while the distance 12 mm between cooling gallery and piston fringe is unchanged, as shown in Fig. 3(a). For radial direction, 10 positions of cooling gallery are taken into account. The positions are obtained by moving cooling gallery from the position C to position D in increment of 1 mm, while the distance 20 mm between cooling gallery and piston crown is unchanged, as shown in Fig. 3(b). Effects of Axial Direction Position of Cooling Gallery. The numerical results of temperature on checked points are presented in Fig. 4(a). It can be seen in Fig. 4(a) that the change in axial direction of cooling gallery strongly affects the temperature on first ring groove, while the effects on other positions are relatively small. Particularly, the temperature decreases 30 when the cooling gallery is moved from A to B position. The relation between the change in temperature and the change in position is nonlinear" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000048_bf00384071-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000048_bf00384071-Figure3-1.png", + "caption": "Fig. 3. Comparison of the shapes and locations ot the plastic regions for free hole and rigid inclusion 7t = 0.875, /3 ~ 0.1.", + "texts": [ + " Tests carried out on the free hole case indicated good agreement with theoretical predictions. Fig. 2 indicates that the plastic modulus desired is one tenth of the elastic modulus, fl = 0.1. Obviously the first estimate, that is the elastic solution corresponds to 15 = 1.0 (p = 0). After changing the right side of equations (15), fourteen times all points of the plastic region fall on a line corresponding to an actual value of = 0.107. Further cycles make the agreement even closer. In the example and comparisons the loading factor is taken as ,~ = aoo/ao = 0.875. In fig. 3 the shapes and locations of the plastic regions are compared for the free hole and the rigid inclusion. It should be noted that at some other values of the loading parameter ~ the shapes can be appreciably different from those presented in this figure. For the free hole the most highly strained points are at 0 = \u00b1~/2, r = a, but for the rigid inclusion the plastic regions develop at 0 = 0 and 0 = ~ and away from the interface with the rigid inclusion. On the basis of fig. 3 it is possible to develop the following argument : Consider a circular inclusion wi th a rb i t r a ry modulus of elasticity. Le t e denote the rat io of the modulus for the insert and the modulus for the plate. Elas t ic-plas t ic solutions are known for three l imit ing values of ~. The smallest possible value is ~ = 0, which corresponds to the free hole. F o r high loads appreciable plast ic flow takes place a t the ne ighborhood of 0 = ~ / 2 , r = a. In a cont inuous p la te it is possible to imagine an insert of the same mate r i a l as the p la te itself" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000712_20110828-6-it-1002.00885-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000712_20110828-6-it-1002.00885-Figure1-1.png", + "caption": "Fig. 1. ADIP (Pendubot)", + "texts": [ + " The paper is organized as follows: Section 2 starts with a modeling of the ADIP and derives a LPV model. Section 3 gives a structure of controller, design specifications and conditions by PDLMIs. Section 4 illustrates simulation and experiment results including comparisons with the PDRE method. Section 5 concludes with remarks. Through out the paper, for a square matrix M , He {M} represents M +M \u2032. 978-3-902661-93-7/11/$20.00 \u00a9 2011 IFAC 9613 10.3182/20110828-6-IT-1002.00885 The experimental apparatus of the ADIP is shown in Fig. 1. The ADIP has two links with joints. Each joint has an encoder to measure each relative angle. The joint 1 is active, which equips with a geared DC motor, while the joint 2 is passive. The control purpose of the ADIP here is to keep the link 2 upright position for a reference angle of the joint 1. The behavior of the ADIP is described as the following motion equation:[ \u03b11 \u03b13cos\u03b812 \u03b13cos\u03b812 \u03b12 ][ \u03b8\u03081 \u03b8\u03082 ] + [ \u03b13\u03b8\u0307 2 2sin\u03b812 \u2212\u03b13\u03b8\u0307 2 1sin\u03b812 ] + [ \u2212\u03b14sin\u03b81 \u2212\u03b15sin\u03b82 ] + [ c1 + c2 \u2212c2 \u2212c2 c2 ][ \u03b8\u03071 \u03b8\u03072 ] = [ \u03c4 0 ] (1) \u03b812 = \u03b81 \u2212 \u03b82, \u03b11 = Jg1 +m1` 2 1 +m2L 2 1 \u03b12 = Jg2 +m2` 2 2, \u03b13 = m2L1`2 \u03b14 = m1g`1 +m2gL1, \u03b15 = m2g`2 where mi is mass of the link i, ci is coefficient of viscous friction of the joint i, Jgi is moment of inertia about the center of gravity of the link i, `i is length between the joint i and the center of gravity of the link i, L1 is length between the joints 1 and 2 and g is the gravity acceleration" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001232_tmag.2012.2198203-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001232_tmag.2012.2198203-Figure2-1.png", + "caption": "Fig. 2. Structure of mover support mechanism.", + "texts": [ + " This actuator mainly consists of a mover, a stator and resonance springs in the and directions that support the mover. The mover is composed of permanent magnets (NbFeB, T) and a back yoke. The stator is composed of an E-type laminated yoke with three phase excitation coils (45 turns). This actuator is assumed to move with a range of mm in the -direction and mm in the -direction, respectively. This movable range assumes use with a compact appliance such as electric shaver, electric toothbrush and etc. The structure of the support mechanism is shown in Fig. 2. To avoid the influence of friction, the mover of -axis is supported by flat springs. The mover of the -axis is supported by linear bearings and coil springs in the -axis mover. The mover can be driven independently in each axis, however, if it moves in the Manuscript received March 02, 2012; revised April 19, 2012; accepted April 30, 2012. Date of current version October 19, 2012. Corresponding author: K. Hirata (e-mail: k-hirata@ams.eng.osaka-u.ac.jp). Digital Object Identifier 10.1109/TMAG.2012" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002168_ifsa-scis.2017.8023339-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002168_ifsa-scis.2017.8023339-Figure4-1.png", + "caption": "Fig. 4. Membership Functions (E1) Fig. 5. Membership Functions (E3)", + "texts": [ + " The four muscles of E1 to E4 mainly generate the movement of dorsal fl xion, ulnar fl xion, palmar fl xion and radial fl xion of a wrist respectively. We paid attention to the characteristic in the movement of these muscles and constructed a fuzzy reasoning rule (Fig. 6 and 10) to estimate the ratio of power of up/down and right/left from EMG data of extensor digitorum muscle (E1) / f exor carpi radialis muscle (E3) and extensor carpi ulnaris muscle (E2) / extensor carpi radialis brevis muscle (E4) respectively. Fig. 4, 5, 8, 9 show the membership functions E1\u223cE4 of the antecedent part, and Fig. 7 and 11 show the singletons of consequent part. The former fuzzy reasoning 1 shows an estimate of the movement of up/down motion of hand, and the latter fuzzy reasoning 2 shows an estimate of the movement of right/left motion of hand. Fuzzy rules were tuned by trial and error. By using these two outputs of the fuzzy reasoning, we are able to estimate the direction where we was going to move by a wrist and use it for the moving instructions of a wheelchair operator" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003361_s0263574719000547-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003361_s0263574719000547-Figure4-1.png", + "caption": "Fig. 4. A 3-DoF manipulator used in experiments. The lengths are l1 = 0.305 m, l2 = 0.225 m and l3 = 0.305 m.", + "texts": [ + " The PD controller will perform linearly only if the mismatch is removed. For that reason, the pole location for the convergence of mismatch to zero (a) must be faster than the PD gains (Kp and Kd). Finally, to compensate for the peaking effect during transient phase, a is selected as a variable with upper and lower bounds (see Section 3.2). A three-degree-of-freedom (3-DoF) planar robot is used to validate the proposed controller in both simulations and experiments. For experiments, 10-ampere Maxon drives along with 150-watt motors are used (Fig. 4). The desired end-effector trajectory qd is a circle in the first quadrant of the x\u2013y plane with radius 0.24 m and center 0.3x\u0302 m + 0.3y\u0302 m. The values for the controller gains are Kp = 40I, Kd = 400I, a = 80 and |eperf|\u221e = 0.01 rad. The end effector starts at x\u0302 0.83 m. The dynamical equations of the 3-DoF robot manipulator used in experiments are calculated from the real robot; see Table I. Since, the adaptive controller is model-free, the dynamics are only used for simulations. The simulation results are divided into three parts" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000108_iecon.2010.5675168-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000108_iecon.2010.5675168-Figure4-1.png", + "caption": "Fig. 4. Analysis spot. Temporal evolution of the pressures along a tooth. The side the more stressed is the one that canalize the flux with the torque creation.", + "texts": [ + "3 the result is a 3D mapping that has in abscises the temporal evolution (here 1/p rotation) and in ordinates the spatial evolution (the value 0 matches with an angle of 0\u00b0 and the last one with a value of 90\u00b0). The diagonal represents the poles that go along the stator. A. Influence of the load angle on the temporal harmonics 1. Evolution of the magnetic pressures The frequency analysis fig.5 presents the evolution of the pressures for a whole rotation of the rotor at the middle of a tooth. The pressures are quite the same all along the teeth aside the extremities where the permeance and the flux increase locally tending to increase the value of the pressure fig.4. In fig.5 we can find the influence of the four poles passing in front of a tooth whose spectrum contains frequencies which all are multiples of 2f, with f the frequency of the current that is to say 2pfmeca with fmeca the mechanical frequency of rotation. In our case the rotor turns at fmeca=43.33Hz which means that the fundamental frequency of the pressure are at 2pfmeca=173.33Hz. As considering that the machine is symmetrical, pressures being proportional to the square of the induction it is enough to know what is going on 1/2p turn of the rotor to know all" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002154_978-3-319-54446-5_16-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002154_978-3-319-54446-5_16-Figure6-1.png", + "caption": "Fig. 6 GuideCane: operational space [30]", + "texts": [ + " 5 Doorplate detection using the ORB feature detector running on a Raspberry Pi B board The NavBelt is based on a belt equipped with eight ultrasonic sensors subtending a 120\u00b0 solid angle of the space, the end user being the apex of this angle. The NavBelt is able to detect the presence of obstacles, and stereophonic headphones inform the VIP about the status of the environment. After 10\u201320 h of self-training, users have been able to travel at 0.6\u20130.8 m/s while avoiding obstacles as small as 10 cm in diameter [29]. The GuideCane is a motorized wheeled cane (Fig. 6). A steering servomotor, operating under the control of the GuideCane\u2019s built-in computer, can steer the guide wheels left and right, relatively to the cane. Attached to each guide wheel is an incremental encoder, and an array of ultrasonic sensors is mounted in a semi-circular fashion above the guide wheels. A digitally controlled fluxgate compass is also mounted above the guide wheels. The embedded computer uses the data from encoders and from the fluxgate compass to compute the relative motion of the traveler\u2014by means of odometry, as well as the instantaneous travel speed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000345_6.2009-5890-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000345_6.2009-5890-Figure9-1.png", + "caption": "Figure 9. Forces acting on chi A-frame mechanism.", + "texts": [ + " The flexible skin/rib structure is suspended between the spars. By design, the acute angle of the members is limited to a range of 29 to 76 degrees. The opening and closing of this angle represents the chi degree-of-freedom of the wing. The forces encountered by the chi mechanism can be attributed to two primary sources: flight loads and skin forces. Flight loads are expected to be inconsequential given the geometry of the wing and chi mechanism. Skin forces, especially those caused by friction, are significant. Figure 9 shows the representative geometry and forces of the chi mechanism. Fk is a skin force derived from the skin tile tests; Fscrew is the actuation force generated by the motor and ball screw assembly; the rest of the forces are internal reaction forces. The final model of the chi mechanism includes the ball screw thread pitch that converts the gearbox shaft angle to linear carriage position, a linear estimate of carriage position (inches) to chi angle (degrees), and a polynomial representation of skin loads as a function of chi angle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001616_ut.2011.5774088-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001616_ut.2011.5774088-Figure9-1.png", + "caption": "Figure 9 Test result, inner pressure changes", + "texts": [], + "surrounding_texts": [ + "The vehicle should run on-land such as a desert as well as in underwater by its tracks. Their axis seal should operate these different conditions with high reliability. Conventional underwater-purpose axis sealing such as mechanical sealing, oil-filled sealing are not proper for onland purpose for a long time. Usual air-purpose watertight axis sealing methods are not proper for the professional underwater use. As shown in Figs.9-10, we proposed the MFS(MagneticFluid Seal)based axis sealing mechanism[12]. The seal blocks between the chamber A and B. The chamber A\u2019s air pressure is controlled depending on the water pressure in underwater. This air barrier prevents the sea water contaminates the MFS. In theory, this MFS has no limit to last the pressure. In proportion to the numbers of the magnetic fluid\u2019s layers, the maximum pressure\u2019s limit is increased. Usually, single axis layer could last 30-50 m depth. We carried out the MFS modeling for the simulation using the magetostatic analysis. This approach concerns the ferrofluid that is under the influence of a magnetic field that arises from a permanent magnet. TABLE I PARAMETER FOR SIMULATION Magnet relative permeability 1.05 Magnetic fluid density 1400 [kg/m3] Magnetic fluid viscosity 0.04 [Pa s] Magnetization of magnetic fluid 48000 [A/m] Magnetic susceptibility of magnetic fluid 0.03 Remanent magnetic flux density 0.8 [T] Figure 12 Test result, inner pressure changes" + ] + }, + { + "image_filename": "designv11_33_0002525_28465-ms-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002525_28465-ms-Figure3-1.png", + "caption": "Figure 3\u2014Schematic of a typical Wire Arc Additive Manufacturing process. The type of welding head can utilize different methods of wire feed.", + "texts": [ + " Additive manufacturing (AM) is the manufacture of an item by the layer upon layer addition of material. There are a wide range of materials that can be produced in this way, with polymers and metallic systems dominating a market which also includes ceramic and biological systems. This paper will discuss the challenges facing metallic additive manufacturing as this is the most relevant to offshore industry applications. The most favored methods are Powder Bed Fusion (PBF), Figure 1, Direct Energy Deposition (DED), Figure 2, and Wire Arc Additive Manufacturing (WAAM), Figure 3. Within these are subsets of technologies that utilize different energy sources, such as lasers, electron beams or electric arc, to fuse the material. They all follow the same principle of a directed energy source and a feedstock, generally a powder or wire, which is fused on the previous layer. Since the commercialization of additive manufacturing in 1987 (Wohlers, T.T., 2014), the development has been rapid compared to that of more traditional subtractive manufacturing methods such as castings and forgings" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000572_0016-0032(65)90271-1-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000572_0016-0032(65)90271-1-Figure7-1.png", + "caption": "FIG. 7. Expanded block diagram of the linear part transfer function Gl(s).", + "texts": [ + ") (3) If r2( t )_ 0; [\u22121, 1] if x = 0; \u22121 if x < 0. \u2013 The symbol sec(x) is used to refer the trigonometric secant function of an angle. That is: sec(x) = 1 cos(x) . \u2013 The symbol \u2016 \u00b7 \u2016 denotes euclidian norm. The dynamics of the PVTOL aircraft, shown in Fig. 1, are described by the following model [21]: mX\u0308 = \u2212 sin \u03b8(f1 + f2) + \u03b5 cos \u03b8(f1 \u2212 f2)L; mY\u0308 = cos \u03b8(f1 + f2) + \u03b5 sin \u03b8(f1 \u2212 f2)L \u2212 mg; In\u03b8\u0308 = (f1 \u2212 f2)L; (1) where X and Y are, respectively, the horizontal and vertical positions; mg is the gravity force exerted on the aircraft mass center, In is the inertia moment, and L is the length from the aircraft mass center to either of its rotors. The angle \u03b8 is formed by the aircraft and the imaginary horizontal plane, and f1 and f2 are the total forces produced by the rotors\u2019 thrust" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002131_metroaerospace.2017.7999605-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002131_metroaerospace.2017.7999605-Figure1-1.png", + "caption": "Fig. 1. Architecture of the measurement system based on electronic balance.", + "texts": [ + " For each subsystem, in literature, several measurement systems are proposed. In this section, a review of them is reported. In case of ESC-motor-propeller subsystem, the most important parameters to be measured are: (i) the thrust forces referred to the motor speeds, (ii) the motor speed response time, and (iii) the power efficiency in terms of thrust for Watt. A common test bench used for measuring the static thrust force is implemented using an electronic weight balance [6]. This system consists of a pivot with one degree of freedom on the middle of a rod (see Fig.1). An ESC-motor-propeller subsystem is placed on one of the rod extremes, with a speed measurement system attached to the drive shaft. On the other shaft extreme, an electronic balance is attached. By means of the shaft, the thrust force exerted by propeller, due to the motor rotation, is transmitted on the electronic balance. By using this test bench, it is possible to measure the thrust force exerted by propeller for each motor rotation speed and to determine the relationship between them. This relationship is used for optimizing the mechanical and the electrical RPAS design and the control method" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000215_cefc.2010.5481322-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000215_cefc.2010.5481322-Figure4-1.png", + "caption": "Fig. 4. Contours of the density of eddy current density (Rotation speed = 50 rpm)", + "texts": [], + "surrounding_texts": [ + "Fig. 1 shows an axial-type magnetic gear used in this study, which mainly consists of a high-speed rotor, a low-speed rotor, and stationary pole pieces. A high-speed rotor generates magnetic harmonics in the air gap between stationary pole pieces and the low-speed rotor. While a high-speed rotor rotates, a low-speed rotor rotates in accordance with the gear ratio." + ] + }, + { + "image_filename": "designv11_33_0001138_j.mechmachtheory.2011.07.015-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001138_j.mechmachtheory.2011.07.015-Figure4-1.png", + "caption": "Fig. 4. Coordinate systems for (a) hob cutter setting with an inclined lead angle and (b) screw surface generation of ZN worm-type hob cutter.", + "texts": [ + " 3(b)) Mij homogeneous coordinate transformation matrix transforming from coordinate system Sj to Si (Eqs. (8), (9), (18), and (19)) mn normal modulus (Fig. 2) mx axial modulus (Figs. 2 and 7), and mx=mn/cos \u03bb1 m21 angular velocity ratio of hob cutter to RA worm gear (Eqs. (28), (33) and (34)) N1,N1 (c) normal vectors of hob cutter tooth surface (Eqs. (12) and (15)) Nx1, Ny1, Nz1 components of the normal vector expressed in coordinate system S1 (Eqs. (34)) p1 lead-per-radian revolution of hob cutter blade surface (Fig. 4) R1, R2 position vectors of hob cutter and RA worm gear, respectively ro, r1, rf outside radius, pitch radius and root radius of hob cutter, respectively (Fig. 3(c)) r2 pitch radius of RA worm gear rc circular tip radius of hob cutter (Fig. 3(b)) rt design parameter of hob cutter (Fig. 3) Si, Sj reference and rotational coordinate systems (i= f, g, p and j=c, 1, 2, 3 ) T1, T2 number of teeth of hob cutter and RA worm gear, respectively tc, tt transverse chordal thicknesses at pitch circle and throat circle of RA worm gear, respectively (Fig. 11) V12 (1) relative velocity vector of hob cutter and RA worm gear expressed in coordinate system S1 (Eqs. (22) and (33)) Vi (1) velocity vectors of hob cutter and RA worm gear (i =1, 2) expressed in coordinate system S1 (Eqs. (22)) \u03b11 pressure angle of hob cutter (Fig. 3(b)) \u03b31 cross angle of hob cutter in generating RA worm gear (Fig. 5) \u03b81 rotation angle of hob cutter in screw surface generation (Fig. 4) \u03bb1 lead angle of hob cutter (Fig. 4) \u03d51, \u03d52 rotational angles of hob cutter and RA worm gear, respectively (Fig. 5) \u03c91, \u03c92 angular velocities of hob cutter and RA worm gear, respectively (Eqs. (24), (25), (28) and (33)) \u03c9i (j) angular velocity vectors, expressed in coordinate system Sj (j=1, p) of hob cutter and RA worm gear (i=1, g) (Eqs. (23)\u2013(26) and (28)) meshing system. As contact progresses from point A (starting point of tooth engagement) to point P (pitch point), there is an approach action, while recess action occurs from point P to point C (final point of tooth engagement)", + " A special case is dx=0 mm for generating the standard proportional tooth worm gear. Similarly, the circular tip of cutting blade can also be expressed in coordinate system Sc as follows: R c\u00f0 \u00de c = rt + bn 2 tan\u03b11 + h 2 + rc cos \u03b1c\u2212 sin \u03b11\u00f0 \u00de 0 bn 2 + h tan\u03b11 2 + rc cos \u03b11\u2212 sin \u03b1 c\u00f0 \u00de 1 2 666666664 3 777777775 ; \u00f07\u00de where 0\u2264\u03b1c\u226490\u2218\u2212\u03b11, and \u03b1c denotes the angular design parameter of hob cutter circular tip. The moving point M2 represents any point on the cutting blade circular tip surface moving from the initial point Mc to the end point Mb. Fig. 4 shows the relationship among the coordinate systems Sc(Xc, Yc, Zc), S1(X1, Y1, Z1) and Sf(Xf, Yf, Zf), where Sc is the blade coordinate system, coordinate system S1 is rigidly connected to the hob cutter tooth surface, and Sf is the reference coordinate system. The inclined angle \u03bb1, formed by axes Zc and Zf, is the lead angle of hob cutter. The tooth surface equation of the ZN worm-type hob cutter can be obtained by considering the blade coordinate system (i.e.Sc) performs a screw motion with respect to the fixed coordinate system Sf" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001063_robio.2011.6181298-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001063_robio.2011.6181298-Figure4-1.png", + "caption": "Fig. 4. Mechanism of right foot of humanoid robot", + "texts": [ + " (1) In this research, we consider the proportional gain (KP ) as 0.0125, the integral gain (KI ) as 0.00025, and the sampling time as 0.001 [s]. III. INVERSE KINEMATICS OF LEG MECHANISM Within the present research, we implemented a forward, linear, and parallel\u2013to\u2013the\u2013floor sole movement for a humanoid robot. To move the foot forward, it is necessary to move the center of gravity; after that, we must rotate the leg over both the Y axis and the X axis, performing a three\u2013 dimensional movement. A model of the right leg mechanism of a humanoid robot is shown in Fig. 4. A front view of the humanoid robot right leg is shown in Fig. 4(a), the right side view is shown in Fig. 4(b). In these leg mechanism models, we consider the hip joint to have three degrees of freedom (DOF) (x,y,z), the knee to have one DOF (x), and the ankle joint to have two DOF (x,y). The \u03a3FR coordinate system is defined using the hip joint as the origin, and considers the forward direction to be the X axis, the left direction to be the Y axis, and the vertical line directed at the floor to be the Z axis, as shown in Fig. 4. We consider the position of the humanoid robot\u2019s hip joint as p1, the position of the knee as p2, the position of the ankle as p3, and the position of the foot as p4. The length of the humanoid robot\u2019s thigh is l1, the length of the calf is l2, and the distance from the ankle joint to the sole is l3. As shown in Fig. 2, the rotation of the hip joint and the ankle over the X axis are given by \u03c81 and \u03c83, respectively, and observing this from the XZ plane it is possible to obtain the line segments as the product of cos\u03c81, l1, and l2, respectively", + " (2) In this equation, the following abbreviations are used: c\u03c81 = cos\u03c81, s\u03b81\u03b82 = sin(\u03b81 + \u03b82), t\u03c82 = tan\u03c82. The coordinates of the current position of the foot can be calculated using Eq. (2), where angles \u03c81, \u03c83, \u03b81, \u03b82, \u03b83 can be retrieved from the angle sensors of each joint. As the foot coordinates obtained from Eq. (2) are added to the variation caused by one step of the sliding walk, it is possible to use inverse kinematics to calculate the angle values after each step. According to the Fig. 4(a), we have \u03c81 = atan2(y4, z4 \u2212 l3), (3) \u03c83 = \u2212\u03c81, (4) therefore with lr = \u221a x2 4 + (z4 \u2212 l3)2, as shown in Fig. 4(b), it is possible to obtain \u03b81 = cos\u22121 (l1 cos\u03c81) 2 + l 2 r \u2212 (l2 cos\u03c81) 2 2l1 cos\u03c81lr +atan2(x4, z4 \u2212 l3), (5) \u03b82 = \u2212\u03c0 + cos\u22121 (l1 cos\u03c81) 2 + (l2 cos\u03c81) 2 \u2212 l 2 r 2l1 cos\u03c81l2 cos\u03c81 , (6) \u03b83 = \u2212(\u03b81 + \u03b82). (7) If we divide the distance between the current position and the subsequent position into 500 equal parts, and solve Eqs. (3), (4), (5), (6), and (7) for each part, then, using the variations, it is possible to perform a sliding walk. Fig. 5(a) shows the initial state of the robot, and Fig. 5(b) shows the state of the robot after a displacement of its center of gravity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003893_0954406219888240-Figure21-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003893_0954406219888240-Figure21-1.png", + "caption": "Figure 21. 3D model of the revolute joint.", + "texts": [ + " Therefore, the maximum clearance model of the ideal spherical joint can be simply defined as Pcle S \u00bc \u00bdc1, c1, c1; 0, 0, 0 T \u00f019\u00de where the first three components denote the translations, the remaining three components denote the rotations, and this clearance model is also applicable in the following discussions. Since the spherical joint has three rotation motions, the joint clearances around the directions of three axes are null. Approximate clearance model of the combined spherical joint As shown in Figure 21, since the combined spherical joint is assembled by a revolute joint and a universal joint, the clearance model of the combined spherical joint can be defined as Pcle S \u00bc Pcle R \u00fe Pcle U \u00f020\u00de where Pcle R and Pcle U denote the clearance models of the revolute joint and the universal joint, respectively. Therefore, the clearance models of the revolute joint and the universal joint are discussed as follows. As shown in Figure 21, the coordinate system O-u0v0w0 is obtained by rotating the O-UVW coordinate system along its V-axis by , and the coordinate system O-uvw is obtained by rotating the O-u0v0w0 coordinate system along its u0-axis by . Comparing to the workspace analysis, it is worth noting that the rotation angle has been newly defined for the convenience of description. When the constraint is pure force or pure moment, the joint clearance can reach the maximum value. According to the first contact model and their hypotheses of revolute joints,21,22 if the constraint is a pure force and along the u-axis, v-axis and w-axis, respectively, the corresponding clearances can be obtained easily and they are c2 and c3, respectively, where c2 and c3 are the gaps shown in Figure 21. As shown in Figure 22, if the constraint is a pure moment and around the u-axis, the clearance can be solved by dR dR 2c2 cos\u00f0 \"\u00de \u00bc LR tan\u00f0 \"\u00de \u00f021\u00de where D is the contact deformation around the u-axis, LR is the equivalent length, which depends on the actual structural parameters, and dR is the inner diameter of the housing. Thus \" LR \u00fe ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi L2 R \u00fe 4dRc2 p 2dR \u00f022\u00de Similarly, when the constraint is a pure moment and around the v-axis, it has the same clearance", + " If the contact force is given, the contact deformation can be expressed by \u00bc 9F2c1 4\u00f0E \u00de2dS\u00f0dS 2c1\u00de 1=3 \u00f029\u00de Therefore, the maximum contact deformation model of the ideal spherical joint can be expressed as Pcon S \u00bc 9F2 uc1 4\u00f0E \u00de2dS\u00f0dS 2c1\u00de 1=3 \" , 9F2 vc1 4\u00f0E \u00de2dS\u00f0dS 2c1\u00de 1=3 , 9F2 wc1 4\u00f0E \u00de2dS\u00f0dS 2c1\u00de 1=3 , 0, 0, 0 #T \u00f030\u00de Approximate contact deformation model of the combined spherical joint As the same as the clearance model of the combined spherical joint, the contact deformation can be defined as Pcon S \u00bc Pcon R \u00fe Pcon U \u00f031\u00de where Pcon R and Pcon U denote the contact deformation models of the revolute joint and the universal joint, respectively. For the revolute joint shown in Figure 21, if the contact deformation is caused by a pure force and along the u-axis or v-axis, the contact model in Hertzian Contact theory is that two cylinders are in parallel contact,25 so the contact deformation can be solved by F \u00bc 4 E LR \u00f032\u00de Thus, the contact deformation is \u00bc 4F E LR \u00f033\u00de If the contact force along the w-axis, the contact model is a parallel contact of two planes, so the contact force is expressed by another expression as F \u00bc 2E ffiffiffiffiffiffi AR r \u00f034\u00de where AR denotes the area of the contact surface shown in Figure 24" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002089_0954406217718857-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002089_0954406217718857-Figure1-1.png", + "caption": "Figure 1. Main components of WHIGPs \u2013 exploded view. WHIGP: water hydraulic internal gear pump.", + "texts": [ + " It is becoming more and more popular in many applications, especially in pollution free applications such as food and drug processing facilities and nuclear industries.1,2 Water hydraulic internal gear pumps (WHIGPs), a kind of water hydraulic power unit, show a promising future in water hydraulics due to their advantages of compactness, low noise level and low flow ripple.3 The gear shaft/journal bearing interface represents one of the key design elements of WHIGPs, and a simulation model is proposed in this paper to investigate the lubricating characteristics of the interface. Figure 1 depicts the main components of the WHIGP: the internal gear pair for fluid delivery, the fillers and the floating plates for the sealing function, and the journal bearings for the bearing function. As shown in Figure 2, by meshing the gears, the fluid is sucked in due to the increase of the tooth space volumes (TSVs) in the suction area (low pressure (LP) in Figure 2) and discharged due to the decrease of the TSVs in the discharge area (high pressure (HP) in Figure 2). Consequently, the gear shaft is subjected to a radial force that needs to be balanced by the water film in the gear shaft/journal bearing interface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000222_j.rcim.2009.09.003-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000222_j.rcim.2009.09.003-Figure4-1.png", + "caption": "Fig. 4. The simulation mechanism of 3-SPR parallel machine tool and a tool path of letters a Carved letters w on a 3D free-form surface S; b Carved letters w on a", + "texts": [ + " The simulation parallel machine tool for letters machining When the normal line z of m of the 3-SPR simulation mechanism is replaced by a tool (such as a mill cutter) and its driving motor, a novel 3-SPR parallel machine tool can be constituted. How to use a 3-SPR parallel machine tool to carve complex letters on the plane P0 or the 3D free-form surface S is a key problem to solve. In the light of the 3-SPR parallel simulation mechanism, a simulation 3-SPR parallel machine tool can be created, see Fig. 4. When some letters are requested to be carved on a 3D free form surface S, both S and a plane P0 are requested to be created by adopting the 3D modeling technique [3]. When some letters are carved on P0, only P0 is request to create. The creating processes are explained below. 1. From the managing tree, return to 2D sketching to modify B of the 3-SPR simulation mechanism, construct several datum planes, and set all datum planes parallel to each other and perpendicular to B by adopting the reference plane command. 2. Based on the given prescript curve data or curve equation, construct each spline curve sk (k=1, 2, y, j) on each datum plane by adopting spline command or data table, and set each spline curve above m. plane P0. 3. Construct S from all spline curves sj (j=1, 2, y, k) by adopting some special 3D surface modeling techniques, such as loft, swept, extrude, rotation commands, etc. Here, S over m is constructed by adopting a loft modeling technique, see Fig. 4a. 3.3. The guiding plane for tool path of letters In order to create a reasonable tool path of letters, a guiding plane P0 for the tool path should be constructed. Its creation processes are explained follows: 1. Modify fixed dimension c=90 cm, C=140 cm. 2. Constitute P0 in 2D sketching of B by using the reference plane command, and set P0||B. 3. Give the distance from P0 to B an initial driving dimension h=530 cm to keep P0 close to and without intersecting into S and m. 4. Copy and paste the spline w of letters (Y S U) onto P0, and set it at the central position and modify their suitable size in workspace of the 3-SPR simulation mechanism by driving dimensions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001473_aero.2010.5446818-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001473_aero.2010.5446818-Figure5-1.png", + "caption": "Figure 5 \u2013 A diagram of spin testing", + "texts": [ + "7 as the alarming point to call for stopping the spin test in order to preserve the test article from being destroyed. In 2008, as part of an ongoing effort, the US Navy was tasked to carry out a rotor spin test for turbine engine disk prognosis. The purpose of this particular test was to evaluate the capability of engine health monitoring systems using the passive eddy current blade tip timing measurements. The tests were conducted at the Rotor Spin Test Facility, Naval Air Warfare Center Aircraft Division (NAWCAD), Patuxent River, Maryland. A simple diagram of the spin test rig is shown in Figure 5. The test article was the first stage fan disk (23 blades) of the Rolls-Royce Spey engine. The disk has a tip-to-tip diameter of 32.4\u201d, the mean rotor diameters are 11.2\u201d and 6\u201d for the outer rim and inner bore respectively. The test was a seeded fault test under LCF cyclic loading with the speed range of 5500rpm to 9000rpm. Two large electron dischargemachined (EDM) notches of one inch (25.4mm) were introduced to the bottom of dovetail slots of blade #1 and blade #9 as crack initiation sites. The blade numbers were marked in reference to the 1/rev tachometer signal" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003894_ecce.2019.8913269-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003894_ecce.2019.8913269-Figure4-1.png", + "caption": "Fig. 4: Magnetic flux density distribution at time zero under 68.75% dynamic eccentricity (a), conventional TSFE analysis(b) VBR method.", + "texts": [ + "6 mm displacement of stator axis in the positive \u201cX axis\u201d direction in the FE grid) was applied. This static eccentricity was applied both to the eddy current frequency domain solver as well as the TSFE model. Through the procedure explained in the previous section, the permeabilities in the entire machine geometry were extracted and imported into the TSFE simulation. The magnetic flux density distribution of the case-study induction motor at time zero, simulated by means of conventional TSFE and VBR methods are given in Fig. 4a and Fig. 4b, respectively. In Fig. 4a, all initial Magnetic Vector Potentials (MVPs), and hence flux densities are assumed to be zero in the conventional TSFE analysis. This leads to a lengthy numerical transient response. It can be concluded from Fig. 4b that the permeabilities of the induction machine are successfully imported into the TSFE simulation through the VBR method described above. It can also be seen that static eccentricity has been taken into account while calculating the permeabilities in the frequency domain solver, as is reflected in the asymmetry in the magnetic flux density distribution in areas 1 through 4. If the permeabilities calculated in the frequency domain are accurate and they reflect both steady state operating frequency, slip, load and eccentricity level, the numerical transient response of the TSFE simulation would be expected to be effectively reduced" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003002_ecc.2018.8550111-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003002_ecc.2018.8550111-Figure1-1.png", + "caption": "Fig. 1: The sensing and communication protocol between agent i and its neighbor j.", + "texts": [ + " The following result is borrowed from [16]: Theorem 1: The equilibrium set Ex = {x \u2208 Rnd : xi = xj , \u2200i, j \u2208 V} of (1) is globally exponentially stable if and only if G has a rooted-out branch. Moreover, there exists x\u221e \u2208 Rd such that x(t) exponentially convergences to (1n\u2297 Id)x \u221e as t \u2192 \u221e. Consider a group of n single-integrator modeled agents: p\u0307i i = ui i (2) where pi i \u2208 R2 and ui i \u2208 R2 denote the position and control input of agent i, respectively, expressed in its local coordinate frame i\u03a3. The orientation misalignment between the local coordinate system i\u03a3 and the global coordinate system g\u03a3 is denoted by an angle \u03b8i, 0 \u2264 \u03b8i \u2264 2\u03c0, as illustrated in Fig. 1. Each agent maintains its local coordinate frame and does not have common sense of the global reference frame g\u03a3. There is a stationary control station acting as a beacon x positioned at px \u2208 R2 such that all agents can sense the direction with respect to it. The direction measurement of agent i is given by an unitary bearing vector, expressed in its local coordinate frame bi i = pi x \u2212 pi i \u2225pi x \u2212 pi i\u2225 . The control station can be a dynamic agent. In this case, it is assumed that the system of n agents- and a control station satisfies the velocity matching condition by additional flocking control scheme [17], [18]", + " For the star framework in Fig. 2, the surrounding formation can be defined by set of desired relative angles \u03b1\u2217 = {\u03b1\u2217 12, \u03b1 \u2217 23, \u03b1 \u2217 34, \u03b1 \u2217 45, \u03b1 \u2217 56}. Note that the group objective is equivalent to achieving b = Q\u0304b\u2217, where Q\u0304(\u03b2) = In\u2297Q(\u03b2) and Q \u2208 R2\u00d72 is a rotation matrix. Note that we use Qi as a shorthand for Q(\u03b8i)-the rotation matrix of agent i. In order to achieve the goal, we assume that each agent i (i \u2208 V) can sense and exchange information about relative orientation with its neighboring agents. Figure 1 describes such sensing and communication topology, the sensing is bidirectional in which agent i senses \u03b4ij and its neighbor j senses \u03b4ji. However, the communication is unidirectional; the neighbor j sends \u03b4ji back to i. From the sensing and communication information, agent i can obtain the relative orientation between i and j as \u03b8ij = PV (\u03b8j \u2212 \u03b8i) = PV (\u03b4ji \u2212 \u03b4ij + \u03c0), (3) where PV (\u03b8j \u2212 \u03b8i) := [\u03b8j \u2212 \u03b8i mod 2\u03c0]\u2212 \u03c0 [2]. Assumption 1: In surrounding formation of star frameworks, the control station assigns desired relative angles and commands desired bearings to agents via communication", + " Define an auxiliary variable z\u0302 = [z\u0302T1 , . . . , z\u0302 T n ] T \u2208 R2n; then the update law and bearing-only formation control law are proposed as \u02d9\u0302zi(t) = \u2211 j\u2208Ni aij(Q T ij z\u0302j(t)\u2212 z\u0302i)(t), (5a) ui i(t) = p\u0307i i(t) = Pbi i [z\u0302i] T \u00d7b \u2217 i , (5b) 1See also the bearing rigidity theory in SE(2) [5]. where, Pbi i = I2 \u2212 bi ib i i T is the projection matrix, Qij = QT i Qj denotes the rotation of \u03b8ij , aij is the element with index (i, j) of adjacency matrix A[aij ] \u2208 Rn\u00d7n of the orientation sensing and communication graph (see Fig. 1), and the skew matrix form of z\u0302 = [x y]T is denoted as [z\u0302i]\u00d7 := [ x \u2212y y x ] . In the estimation law (5a), the relative orientations \u03b8ij , j \u2208 Ni, are measured and communicated between neighboring agents, while the states z\u0302j , j \u2208 Ni, are sent to agent i by communication. The dynamics (5a) can be rewritten as \u02d9\u0302zi = \u2211 j\u2208Ni aij(Qij z\u0302j \u2212 z\u0302i) QT i \u02d9\u0302zi = \u2211 j\u2208Ni aij(Q T j z\u0302j \u2212QT i z\u0302i) blkdiag{QT i } \u02d9\u0302z = \u2212(L\u2297 I2)blkdiag{QT i }z\u0302, (6) where blkdiag{.} denotes the block diagonal matrix, z\u0302 = [z\u0302T1 , " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001301_1.52410-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001301_1.52410-Figure5-1.png", + "caption": "Fig. 5 Geometrical relationship of the record disk orbit.", + "texts": [ + " (23) and minor and major radius of a\u0302 x2 0 _x2 0 =!2 p (24) b\u0302 2 x2 0 _x2 0 =!2 p (25) in the x-axis, and y-axis, respectively. In addition, if the initial position and velocity in the z-axis are selected as z 0 x 0 tan (26) _z 0 _x 0 tan (27) and is selected as =3, then the motion in the z-axis has the same phase as that in the x-axis, and the trajectory in the Hill\u2019s coordinate becomes a circle of radius 2 x2 0 _x2 0 =!2 p (28) This trajectory, shown in Fig. 4, is well known as a cart orbit, or a record disk orbit. Figure 5 shows the geometrical relation between the inclination i and eccentricity e in the record disk orbit, which is drawn from a side view. If the relative inclination 2i and eccentricity e are small enough to treat the geometrical relation in term of linearity, the parameter can be represented as arctan i e (29) whichmust be =3 to obtain a record disk orbit, therefore, the relative inclination i must be related to the eccentricity e, by i 3 p e (30) D ow nl oa de d by U N IV E R SI T Y O F O K L A H O M A o n Ja nu ar y 24 , 2 01 5 | h ttp :// ar c" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001135_1077546309353917-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001135_1077546309353917-Figure1-1.png", + "caption": "Figure 1. A rotating and translating beam element.", + "texts": [ + " The complete derivation of the governing equations will be presented in Section 4, since the equations are associated with shape functions, which are shown in Section 3. The Lagrangian takes the form (Shabana, 1997) Le \u00bc Te Ve Ye \u00f01\u00de where Te, Ve, and Ye are the kinetic energy, the flexural strain energy, and the work done by a tensile longitudinal load for a translating and rotating beam element, respectively. The three terms are formulated as follows: Based on the modeling of a translating and rotating beam element shown in Figure 1, the velocity of an arbitrary point on the beam element is given as V * \u00bc Vox _ v\u00fe _u\u00bd i\u00fe Voy \u00fe _ x\u00fe u\u00f0 \u00de \u00fe _v j \u00f02\u00de where (Vox, Voy) is the absolute velocity of the left end point of the beam element; _ is the angular velocity of the beam element; u and v are the longitudinal and transverse displacements of an arbitrary point on the beam element, respectively; u and v are longitudinal and transverse deflections, respectively; x is a longitudinal position on the beam element shown in Figure 1. If we let be the mass per unit volume of element material; A, the element cross-sectional area, and l the element length, then the kinetic energy of an element is expressed as Te \u00bc 1 2 A Zl 0 Vox _ v\u00fe _u\u00bd 2 \u00fe Voy \u00fe _ x\u00fe u\u00f0 \u00de \u00fe _v 2n o dx \u00f03\u00de The flexural strain energy of uniform axially rigid element with modulus of elasticity, E and second moment of area, I is given as Ve \u00bc 1 2 EI Zl 0 v2xxdx \u00f04\u00de For high-speed rotation in the flexible mechanism, each moving link has inertial axial load, which might contribute transverse deflection (Meirovitch, 1967)", + " Thus, the work done by a tensile longitudinal load P, in an element that undergoes an elastic transverse deflection is given by (Meirovitch, 1967) Ye \u00bc 1 2 Zl 0 Pv2xdx \u00f05\u00de Longitudinal loads in a moving mechanism element are not constant, and depend both on the position in the element and on time. With the longitudinal elastic motions neglected, the longitudinal loads may be derived from the rigid body inertia forces, and can be expressed as P \u00bc PR Aaox l x\u00f0 \u00de \u00fe 1 2 A _ 2 l 2 x2 \u00f06\u00de where PR is an external longitudinal load acting at the right hand end of an element, and aox is the absolute acceleration of the point O in the x direction shown in Figure 1. In view of the high axial stiffness of a beam, it is reasonable to consider the beam as being rigid in its longitudinal direction. A two-node beam element having a rigid longitudinal motion is shown in Figure 2, where u1 is the longitudinal displacement; vi (i\u00bc 1, 2) are the transverse displacements; i (i\u00bc 1, 2) are the rotations; mi (i\u00bc 1, 2) are the curvatures. Because the DFE method is to approximate the displacements as polynomials, the longitudinal and transverse deflections are given as Longitudinal deflection: u \u00bc u1 \u00f07\u00de Transverse deflection: v \u00bc \u00bdN f eg \u00f08\u00de at The University of Iowa Libraries on June 21, 2015jvc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002428_hpd.2017.8261076-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002428_hpd.2017.8261076-Figure1-1.png", + "caption": "Fig. 1. Fiber coupled T-bar package, Coherent|Dilas.", + "texts": [ + " An emerging application space is laser additive manufacturing. High power diode lasers for material processing use fast and slow axis cylindrical optics to collimate the laser diode light and focus that output directly onto the work piece or an optical delivery fiber. For fiber delivered systems each emitter or group of emitters is focused into a fiber. Those emitters or group of emitters can be optically stacked to increase the output power. Those packages typically have an output power of 50-400 W (figure 1). Diode laser bars are established as the solution for high power diode laser systems because of their high average power capability. Geometrically stacking those diode in combination with water cooled plate heat sinks enabled even high power out of a single package. Those stacks typically have output powers exceeding 1000 W. To further scale the power the stacks are combined by polarization and wavelength. Figure 2 below illustrates an example for an 8 kW fiber coupled diode laser using diode stack combinations based on the previously mentioned techniques" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000577_s12206-010-0606-y-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000577_s12206-010-0606-y-Figure1-1.png", + "caption": "Fig. 1. The schematic view for flight vehicle.", + "texts": [ + " This paper develops an analytical model for studying the aeroelasticity of a two-section flight vehicle with freeplay in rotation. For this purpose, component mode synthesis is used for elastic deflection, and slender body theorem is assumed for aerodynamic modeling. The solution of the equation of motion in time-domain is analyzed to study the effects of the free fittings under maximum altitude, range, and instability of the flight vehicle. 2. Formulation 2.1 Model definition A two-section flexible flight vehicle with a torsional spring at joint, in a 2-D space is illustrated in Fig. 1. Each section of the flight vehicle has a body coordinate system originated at the joint and rotates with the appropriate section. The inertia frame is assumed to be at the flight starting point by which the \u2020 This paper was recommended for publication in revised form by Associate Editor Eung-Soo Shin *Corresponding author. Tel.: +98 21 6616 4617, Fax.: +98 21 6602 2731 E-mail address: m_ehramian@yahoo.com \u00a9 KSME & Springer 2010 relative displacement of the origin of the body coordinates is evaluated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000583_09596518jsce969-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000583_09596518jsce969-Figure1-1.png", + "caption": "Fig. 1 Schematic diagram of the strip casting process", + "texts": [ + " Section 3 presents the model-free adaptive fuzzy sliding-mode control strategy. Section 4 describes the numerical results of the proposed controller. Final conclusions are presented in section 5. A process mathematical model, which describes the relationship between the command inputs and the measured outputs, is required for the numerical simulation to evaluate the dynamic performance of a model-free controller. The mathematical model for the molten steel levelling dynamics developed in reference [7] is adopted and described in this section. Figure 1 shows a schematic drawing of the strip casting process. For developing the mathematical model, it is assumed that the molten steel is incompressible and the two rolls are identical. The continuity equation of the liquid steel can be described as dV dt ~Qin{Qout \u00f01\u00de where V is the volume of molten steel stored between the twin-roll cylinders, Qin is the input flowrate into the space between the roll cylinders, and Qout is the output flowrate from the roll cylinders. The volume V can be calculated as V ~ALr \u00f02\u00de where Lr is the length of the roll cylinders and A is the oblique area shown in Fig. 1, which is given by A~2 \u00f0y 0 xg(t) 2 zR{ ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi R2{y2 p dy \u00f03\u00de where xg(t) is the roll gap, R is the radius of the roll cylinder, and y(t) is the height of molten metal above the axis of the rollers. Substituting equations (2) and (3) into equation (1) yields Proc. IMechE Vol. 224 Part I: J. Systems and Control Engineering JSCE969 at UNIVERSITE DE MONTREAL on June 24, 2015pii.sagepub.comDownloaded from dV dt ~Lr dA dt ~Lr y dxg dt z xgz2R{2 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi R2{y2 p dy dt \u00f04\u00de If xgz2R{2 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi R2{y2 p is defined as Br(xg, y), the following form can be derived from equation (1) dy dt ~ 1 Br(xg, y)Lr Qin{Qout{Lry dxg dt \u00f05\u00de Here, the input flowrate Qin can be derived from the stopper opening height h(t) and a non-linear, time-varying, input flowrate parameter a(t) that depends on the shape of the nozzle and the stopper, the clogging/unclogging dynamics, and the height and viscosity of the molten metal in the tundish Qin~a(t)h(t) \u00f06\u00de The orifice opening h(t), equal to the height of the stopper, is controlled by an electric servo motor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000993_055508-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000993_055508-Figure1-1.png", + "caption": "Figure 1. Sketch of the geometry of the problem as well as the two Cartesian coordinates, (X, Y, Z) and (x, y, z), used in our theoretical analysis. (a) A magnetotactic bacterium has the shape of an elongated prolate spheroid with arbitrary eccentricity E with the semi-axis a and c = a/ \u221a (1 \u2212 E2). The Cartesian coordinates (X, Y, Z) represent a reference frame fixed in the laboratory, while (x, y, z) denote a reference frame fixed in the body of the spheroidal bacterium. (b) The three Euler angles (\u03b8, \u03c8, \u03c6) connect the two Cartesian coordinates (X, Y, Z) and (x, y, z).", + "texts": [ + " The third assumption is that the motion of a magnetotactic bacterium is propelled by a force acting at a fixed point on its membrane. It is spherical geometry that allows Nogueira and Lins de Barros (1995) to derive a system of six ordinary differential equations that govern the motion of a swimming spherical magnetotactic bacterium. The primary objective of this paper is to study, via both theoretical and experimental methods, the swimming motion of magnetotactic bacteria that have the shape of an elongated prolate spheroid with arbitrary eccentricity E which, as depicted in figure 1(a), may be mathematically described as x2 a2 + y2 a2 + z2 c2 = 1, (3) where c2 = a2/(1 \u2212 E2) with 0< E < 1. A microscope image of a magnetotactic bacterium that has a spheroidal shape and is approximately described by (3) is displayed in figure 7. It is of primary importance to note that, in comparison with spherical geometry, elongated spheroidal geometry introduces not only complicated mathematics but also important new physics/dynamics for a better understanding of the swimming motion of nonspherical magnetotactic bacteria. Describing a swimming spheroidal bacterium requires threedimensional solutions of the spheroidal Stokes flow in a viscous and incompressible fluid. The following two solutions of the Stokes flow are needed: (i) a three-dimensional Stokes flow driven by a translating prolate spheroid of arbitrary eccentricity E at an arbitrary angle of attack \u03b3 , the angle between the direction of translation v and the symmetry axis z in figure 1(a), and (ii) a three-dimensional Stokes flow driven by a rotating spheroid of arbitrary eccentricity E with an arbitrary angle \u03b1, the angle between the angular velocity \u2126 and the symmetry axis z in figure 1(a). If the above two solutions are available, one can then derive an expression for the corresponding drag D\u00b5 and torque T\u00b5, which would be much more complicated than those given by (1) and (2). Motivated by the desire to understand the dynamics of a slowly swimming magnetotactic bacterium that has the shape of an elongated prolate spheroid, we have made the first application of the Papkovich\u2013Neuber formulation to the spheroidal Stokes problem and obtained the Papkovich\u2013Neuber-type analytical solutions in prolate spheroidal coordinates for the three-dimensional spheroidal Stokes flow that is driven by either a translating spheroid at an arbitrary angle \u03b3 or a rotating spheroid with an arbitrary angle \u03b1 (Kong et al 2012)", + " This is followed in section 3 by discussing the swimming motion of spheroidal magnetotactic bacteria with different eccentricities and magnetic moments and by comparing the theoretical swimming patterns to the trajectories of swimming motion observed in the laboratory experiments. The paper ends in section 4 with a summary and some remarks. Consider the swimming motion of a spheroidal magnetotactic bacterium having the magnetic moment vector m with mass M in a uniform viscous fluid. The geometry of the magnetotactic bacterium, together with the two Cartesian coordinates (x, y, z) and (X, Y, Z), is sketched in figure 1(a). Cartesian coordinates (x, y, z) along with the corresponding unit vectors (x\u0302, y\u0302, z\u0302) represent a reference frame fixed in the bacterium\u2019s body with z at its symmetry axis; this reference will be referred to as the body frame. The position of its center of mass is described by the position vector R = XX\u0302 + Y Y\u0302 + Z Z\u0302, in the Cartesian coordinates (X, Y, Z) with the corresponding unit vectors (X\u0302, Y\u0302, Z\u0302) which is fixed in the laboratory; this reference will be referred to as the laboratory frame", + " The velocity vector vB, when \u03b3 6= 0, is marked by the three quantities |vB|, \u03b3 = cos\u22121 [ z\u0302 \u00b7 vB |vB| ] , \u03c4 = cos\u22121 [ x\u0302 \u00b7 vB sin \u03b3 |vB| ] , (4) while there are also three quantities characterizing the vector \u2126: | |, \u03b1 = cos\u22121 [ z\u0302 \u00b7 | | ] , \u03b2 = cos\u22121 [ x\u0302 \u00b7 sin\u03b1| | ] (5) for \u03b1 6= 0. The axisymmetric case \u03b3 = 0 (or \u03b1 = 0) is specially treated. Moreover, the relative position between (x, y, z) and (X, Y, Z) is determined by the position vector R together with the three Euler angles (\u03b8, \u03c6,\u03c8) illustrated in figure 1(b). There exist 12 degrees of freedom that completely determine the swimming motion of a spheroidal magnetotactic bacterium: its position in the laboratory R = (X, Y, Z), the three components of its velocity vector vL, its angular rotating velocity \u2126= ( x , y, z) and the three Euler angles \u03b8, \u03c6 and \u03c8 . We shall also make, similarly to the study by Nogueira and Lins de Barros (1995), the following three assumptions: (i) the body of the magnetotactic bacterium is non-deformable and, hence, the equations of rigid-body dynamics are applicable; (ii) the interaction between different magnetotactic bacteria (see, for example, Ishikawa et al 2007) on their swimming motion is weak and hence negligible; and (iii) the translation and rotation of the bacterium are powered by the rapid rotation of helical flagellar filaments at a fixed point P , as shown in figure 1(a), in connection with a force FB in the body frame: FB = F12 [ cos(\u03c90t)x\u0302 + sin(\u03c90t)y\u0302 ] + F3z\u0302, (6) where \u03c90 is the frequency of flagellum rotation and F12, F3 and \u03c90 may be regarded as parameters of the problem. There exist, however, three major differences between our dynamical model and that of Nogueira and Lins de Barros (1995). Firstly, our magnetotactic bacterium has the shape of an elongated prolate spheroid, while theirs is spherical. It will be seen that the swimming motion of a magnetotactic bacterium is highly sensitive to its shape marked by the size of eccentricity E " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000924_s1064230711040083-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000924_s1064230711040083-Figure9-1.png", + "caption": "Fig. 9. Specific features of motion of the underwater vehicle along a curvilinear trajectory.", + "texts": [ + " In other words, the application of the proposed system for \u03b5\u0303 \u03b5\u0303n \u03b5\u0303 X\u0303* of the underwater vehicle: v(t) = scale (m/s), z(t) = scale \u00b7 5 (m), || (t)||= scale \u00b7 0.2 (m), || (t)|| = 2 \u00b7 scale (m). \u03b5\u0303n \u03b5\u0303 \u03b5\u0303n \u03b5\u0303 JOURNAL OF COMPUTER AND SYSTEMS SCIENCES INTERNATIONAL Vol. 50 No. 4 2011 A METHOD FOR FORMING PROGRAM CONTROL FOR VELOCITY REGIME 681 (t) formation makes it possible to increase the accuracy of motion of the underwater vehicle along the assigned trajectory by more than a factor of 7 and increase the velocity by a factor of 1.35 without compli cating the control system of this vehicle. Figure 9 shows the given trajectory of the underwater vehicle (curve 1), the trajectory formed by the signal (t) of the synthesized system (curve 2), and the real trajectory of the underwater vehicle (curve 3). It can be seen from this figure that the real position of the underwater vehicle and the objective point (t) are always at different sides of the assigned trajectory. Thus, the results of simulation completely proved the performance and high efficiency of the new con trol principle for complex dynamic objects used to synthesize systems for automatic formation of program signals (t) controlling the motion of an underwater vehicle along complex spatial trajectories with lim iting high velocity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000476_9783527650002.ch14-Figure14.38-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000476_9783527650002.ch14-Figure14.38-1.png", + "caption": "Figure 14.38 Examples of optical confi gurations for passive optical sensors.", + "texts": [ + " Multiplex capabilities represent the main advantage of such sensors compared to electrochemical devices. Very different techniques are used in the design of fl uorescence - based optical sensors. Passive and active modes of operation should be distinguished. \u2022 In the passive mode, the optical device measures the variation in fl uorescence characteristics (intensity, lifetime, and polarization) of an intrinsically fl uorescent analyte. The optical device can have different optical confi gurations involving in most cases an optical fi ber (passive optode) (Figure 14.38 ). \u2022 In the active mode, the optical device uses for optical transduction the changes in the fl uorescence characteristics of a fl uorescent molecular sensor (as described in the preceding sections) resulting from the interaction with an analyte. 4) The two main optical confi gurations are as following: 4) Alternatively, in some cases, the fl uorophore chemically reacts with the analyte \u2013 these are outside the scope of this chapter. 1) In most cases, the fl uorescent molecular sensor is immobilized at the tip of an optical fi ber (active optode) in a matrix that permits diffusion of the analyte (Figure 14" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002938_icelmach.2018.8507253-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002938_icelmach.2018.8507253-Figure5-1.png", + "caption": "Fig. 5. Thermal analysis for nominal operating point 100 Nm \u2013 60,000 rpm: (a) FEA + correlations, and (b, c) MotorCad.", + "texts": [ + " confirm the obtained results. The maximum temperature is located in the winding (hot spot in the endwinding at 105 \u00b0C). The maximum temperature in the slots is 100\u00b0C. The difference is mainly due to the conservative equivalent slot thermal conductivity that has been considered. ( ) ( ) ( ) ( ) ( ) 2 3 2 3 Nu 7.49 17.02 H W 22.43 H W 9.94 H W 0.065 D L Re Pr 1 0.04 D L Re Pr = \u2212 \u22c5 + \u22c5 \u2212 \u22c5 \u22c5 \u22c5 \u22c5 + + \u22c5 \u22c5 \u22c5 (5) For the nominal operating point, the same approach has been used. The results obtained by thermal FEA (see Fig. 5.a) shows that the maximum temperature is located in the stator slots (117\u00b0C). The rotor temperature is lower than 95\u00b0C, which is not critical for PM demagnetization. When using the software MotorCad, the maximum temperature inside the slots is 119\u00b0C. However, the endwinding are hotter and can reach 136\u00b0C (see Fig. 5.b-c). This is due to lower heat convection coefficient in the endwinding, and this has not been considered in the FEA model. However, this temperature remains acceptable for the considered enameled wires. The magnet temperature is comprised between 92 and 107 \u00b0C depending on the axial position. This confirms the accuracy of the proposed approach combining FEA and analytical correlations. In this paper, four high-speed PM machines have been studied for turbogenerator application. The electromagnetic and structural analyzes have been performed in order to study the influence of the rotor structure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001622_icsens.2010.5690737-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001622_icsens.2010.5690737-Figure5-1.png", + "caption": "Fig. 5. Illustrative 3D picture of the standard and modified lift-off processes.", + "texts": [ + " To achieve a reliable and robust electrochemical measurement system, such delamination cannot be tolerated since it gives misleading measurement results by affecting the total electrode area. Additionally, it is also observed that the layer delamination propagates during the electrochemical tests, resulting in random drift in the output signal. C. Improved Design and Fabrication for Microelectrodes In the improved version of the microelectrodes, the lift-off process and the electrode design are modified to solve the oxide delamination problem. Fig. 5 illustrates the difference between the lift-off processes employed in the two versions of microelectrodes. In the modified version, the LOR undercut is reduced to have notches at the metal edges in order to enhance the oxide adhesion. However, reducing the undercut too much may result in deposition of almost or completely conformal metal film, which would lengthen the lift-off process considerably and possibly require ultrasonic agitation to break the metal. Thus, the resist development time has to be adjusted to optimize the LOR undercut, providing notches at the edges, while maintaining easy and precise metal patterning" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002481_ibcast.2018.8312233-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002481_ibcast.2018.8312233-Figure1-1.png", + "caption": "Fig. 1.Quadrotor configuration with reference frames", + "texts": [], + "surrounding_texts": [ + "\u03b6B and \u03c9B characterize position vector and body angular rate vector of quadrotor model in body reference frame respectively. Both can be translated into inertial reference frame by means of R and T respectively [19]. The symbol R corresponds to rotation matrix and T denotes transfer matrix. Equations (1) and (2) express the quadrotor kinematics as follows: \u03b6E = R\u03b6B (1) \u0398\u0307 = (2) \u03b6E and \u03b6B represent quadrotor position vectors in earth and body reference frames respectively i.e. \u03b6 = [ ] . While = [ ] and \u0398\u0307 = [ \u0307 \u0307 \u0307 ] denote body angular rates and Euler angular rates respectively. The rotation matrix R and transfer matrix T can be written: R= \u03c8 \u03c8 \u2212 ++ \u2212\u2212 T = 10 \u22120 Matrix T depends on Euler angles and it is invertible if following conditions exist: for roll (\u2212 2 < < 2), for pitch (\u2212 2 < < 2) and for yaw (\u2212 < < ), otherwise, singularity will occur. In other words, maneuvering beyond these constraints using this notion can\u2019t be performed. Quaternion approach can be used to avoid these singularity issues. Recently, some studies have designed flight controller on special Euclidean group SE(3) and the adopted geometric control approaches solve the singularity issues as well [20, 21]." + ] + }, + { + "image_filename": "designv11_33_0001164_sii.2011.6147547-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001164_sii.2011.6147547-Figure3-1.png", + "caption": "Fig. 3. A cart-table model [4]", + "texts": [ + " Applying preview control theory to generate the trajectory of center of mass (CoM) that follows desired ZMP is proposed by Kajita et al. [4]. This method can generate a trajectory of CoM that corresponds desired zero moment point (ZMP) at the starting point of the trajectory, so the method is suitable for highly frequent repetitive pattern generation. Since the pattern generation cycle time is 5 [ms] in this research, it is proper to use the method. The method will be overviewed for further explanation. A humanoid robot is modeled by a cart whose mass is M, and a massless table (Fig. 3). p and x in Fig. 3 are x coordinates of the ZMP and CoM, respectively. ZMP equation can be led as follows: p = x\u2212 zc g x\u0308 , (1) where g is gravity acceleration. Then (1) can be translated into a strictly proper dynamical system as: d dt x = \u23a1 \u23a30 1 0 0 0 1 0 0 0 \u23a4 \u23a6x+ \u23a1 \u23a30 0 1 \u23a4 \u23a6u , p = [ 1 0 \u2212zc/g ] x, u\u2261 ... x , (2) where x \u2261 [ x x\u0307 x\u0308 ]T . Let input u of (2) be the time derivative of x\u0308, then (2) is discretized with sampling time of \u0394 t (5 [ms]) as follows: xk+1 = Axk +buk , p = cxk , (3) where A\u2261 \u23a1 \u23a31 \u0394 t \u0394 t2 2 0 1 \u0394 t 0 0 1 \u23a4 \u23a6 , b\u2261 \u23a1 \u23a2\u23a3 \u0394 t3 6 \u0394 t2 2 \u0394 t \u23a4 \u23a5\u23a6 , c\u2261 [ 1 0 \u2212zc g ] , xk \u2261 [ x(k\u0394 t) x\u0307(k\u0394 t) x\u0308(k\u0394 t) ]T " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003466_j.mechmachtheory.2019.06.017-Figure31-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003466_j.mechmachtheory.2019.06.017-Figure31-1.png", + "caption": "Fig. 31. Illustration of finite element models.", + "texts": [ + " This kind of predesigned transmission errors can absorb the non-continuous linear function caused by misalignments, which has been proved in [30] . Contact stress and bending stress of case (I)\u2013(IV) are calculated by finite element method (FEM) in this section. The bearing contact and the edge contact during the meshing cycle are investigated. Both the pinion surfaces and the gear surfaces of the finite element models are generated based on the surface equation directly, so as to improve the calculating accuracy of the contact and bending stresses. Fig. 31 shows the applied models with five pairs of teeth, which is automatically generated by a developed computer program. The stress analysis is performed by applying a general-purpose finite element analysis (FEA) application ABAQUS?c? c . The boundary conditions of the face-gear drive finite element models are set as the same in [11] . The FEM model is composed of 296,400 nodes and 232,288 elements. The Poisson\u2019s ratio and Young\u2019s modulus of the material are set to be 0.29 and 2.068 \u00d7 10 5 MPa, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001452_1464419311408949-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001452_1464419311408949-Figure2-1.png", + "caption": "Fig. 2 (a) The ball bearing model and the defined paths and (b) the ADAMS model of the ball bearing", + "texts": [ + " The balls are positioned equi-pitched around the inner ring and there is no direct interaction between them. The rigidity assumption of some of the ball bearing components can limit the observation of high-frequency vibrations. On the other hand, a flexibility assumption of these components leads to a severe increase in central processing unit (CPU) requirements. After the three-dimensional (3D) models of the ball bearing components (rings and balls) are imported to the ADAMS environment, the paths on which the components get into contact are defined. As shown in Fig. 2(a), these paths are defined as curves. The balls and the races get into contact on these paths. In ADAMS software, this kind of contacts can be defined either as solid-to-solid or curve-to-curve. The main idea behind solid-to-solid and curve-tocurve contacts is to impose constraints between certain DOF in the ADAMS model. Although solid-tosolid contact option is attractive, this choice increases the CPU times significantly and, when the structure becomes more complex after the rotor part is added, the CPU requirements can easily exceed the acceptable limit", + " After the paths and contacts are defined, the kinematical relationships between the parts have to be specified. Initially, for the model of the ball bearing itself, the outer ring is supposed to be fixed to the ground. After the housings are assembled, this constrain is changed. All the balls, the outer and inner rings are constrained to be in the same plane using planar joints. Inner ring is fastened on balls. In other words, the supports that hold inner ring in radial direction, are the balls. The final state of the ball bearing model is shown in Fig. 2(b). The angular velocities of the balls, cage, and the inner ring are shown in Fig. 3. Motion is defined on the inner ring. It is accelerated to 1200 r/min in 10 s and continued its rotation at that speed for 3 s more. It can be seen that the relationship between the angular velocities of the inner ring, cage, and a ball is as described in Fig. 3. This confirms that the ball bearing model is functioning properly and satisfies the kinematic behaviour of the real ball bearing. After the ball bearing model is created, the rotor part is built as an assembly of a shaft and a disc at one end of the shaft" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003515_bmz047-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003515_bmz047-Figure6-1.png", + "caption": "Figure 6. Schematic separation mechanism of monolithic [PMAPA] monolithic column in the CEC system at pH 7.0.", + "texts": [ + " However, separation of molecules is mainly driven due to the strong interactions between hydrophobic nonpolar groups D ow nloaded from https://academ ic.oup.com /chrom sci/advance-article-abstract/doi/10.1093/chrom sci/bm z047/5533476 by N ottingham Trent U niversity user on 20 July 2019 (aromatic rings) of the capillary [PMAPA] monolithic column and the hydrophobic phenyl rings of DA and NE molecules. Additionally, in [PMAPA] monolithic column used in CEC system, higher elution times was observed for DA than NE molecule as shown Figure 6. To control if the last eluted compound is DA or not, the separation performance was assessed by comparing two chromatograms of DA and NE samples loaded separately. DA elution with same retention time was shown in Figure 7. Problems related to the strong interactions and variability in accessibility to the binding sites impose limitations for the use of CEC-based chromatography. These problems lead to poor chromatographic performance, manifested in broad and tailing peaks. Inevitable peak tailings were observed when the separation of molecules is mainly governed by the strong interactions between hydrophobic nonpolar groups (aromatic rings) of the capillary [PMAPA] monolithic column and the hydrophobic phenyl rings of DA and NE molecules" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001622_icsens.2010.5690737-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001622_icsens.2010.5690737-Figure2-1.png", + "caption": "Fig. 2. Simplified process steps used in the fabrication of microelectrodes composed of Au WE and Pt CE.", + "texts": [ + " In the control part of the potentiostat circuit, a feedback loop ensures that the potential difference between the onchip WE and the external RE is kept at the applied level, while preventing any current to flow through the RE [10]. In the current measurement part, a transimpedance amplifier keeps the WE potential at the ground (virtual) potential and converts the current flowing between the CE and the WE to a voltage through a feedback resistor. The output voltage is then digitized and processed to be displayed in the CV plots. The first version of the electrochemical sensor chip is fabricated through a series of standard microfabrication steps, as illustrated in Fig. 2. Au and Pt microelectrodes are fabricated on a 100 mm diameter silicon wafer with 500 nm of top silicon dioxide (SiO2, hereafter shortened as \u201coxide\u201d) film for isolation. First, a standard double-layer lift-off process is used to realize Au WE electrodes. In this step, spin-coated lift-off resist (LOR) and positive-tone photoresist are patterned by photolithography, then 20 nm of Ti and 200 nm of Au are deposited by e-beam evaporation. Metals on the unexposed regions are removed by stripping the resists, forming the electrodes and the interconnections" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001540_j.1747-1567.2011.00755.x-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001540_j.1747-1567.2011.00755.x-Figure1-1.png", + "caption": "Figure 1 Drop tower and mounting PCB assembly.", + "texts": [ + " The JESD22-B111 is a standard for board-level drop test methods of components for handheld electronic Experimental Techniques 36 (2012) 60\u201369 \u00a9 2011, Society for Experimental Mechanics 61 products. The standard entails the specification of test components, test board assembly, test apparatus, and test procedure. Drop towers are available on the market at an expensive cost; therefore, small-scale laboratories cannot afford them. Furthermore, when the test board or test procedure is designed in-house for certain products, the demand of a versatile drop tower is inevitable. The detailed infrastructure drop tower and mounting PCB assembly are shown in Fig. 1. The drop tester is positioned on a rigid base. Means are provided in the equipment to exclude rebound and to prevent shock to the test board. A base plate with standoffs is stiff mounted on the drop table. The thickness and mounting locations of the base plate are selected such that there is no relative movement between the drop table and any part of the base plate during drop testing. The PCB assembly is mounted to the base plate standoffs using four screws, one at each corner of the board. The experiment adopted horizontal orientation with components facing down in the cause of PCB flexure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000770_lindi.2012.6319490-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000770_lindi.2012.6319490-Figure3-1.png", + "caption": "Figure 3. Presentation of angular deviation (\u03d5 )", + "texts": [ + " It compares x and y outputs of the bicycle\u2019s kinematic model with the xD and yD (desired) target coordinates. The results of subtraction are xE and yE (error) which shows how far away the robot is from the goal (3) yyyxxx DEDE \u2212=\u2212= , (3) The controller needs to know what the distance is between the currently located vehicle and the target, therefore the distance in (4), should be calculated. We also need to know the deviation between the angle of the robot\u2019s direction and the proper angle which leads to the goal. Formulae (5) and (6) can be written with the help of diagram in Fig. 3. These formulae are also used for P controller in the [2]. 22 EE yxd += (4) \u239f\u239f \u23a0 \u239e \u239c\u239c \u239d \u239b = \u2212 E E x y1tan\u03b1 (5) \u03b1\u03b8\u03d5 \u2212= (6) III. BUILDING FUZZY ROUTE CONTROLLER The Fuzzy controller includes two inputs as can be seen in Fig. 4: d as distance and phi as the angle of direction error. Outputs are v as velocity and gamma as angle of steering wheel, with which the model is controlled. These outputs will be the inputs of the dynamic and kinematic model. We chose trimf as a membership function for easy use and implementation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001520_pes.2011.6039209-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001520_pes.2011.6039209-Figure3-1.png", + "caption": "Fig. 3. Schematic diagram of the lumped-mass equivalent model", + "texts": [ + " The differential equations of the two-mass model are: ( )\u23aa \u23aa \u23aa \u23a9 \u23aa \u23aa \u23aa \u23a8 \u23a7 \u2212= \u2212\u2212= \u2212\u2212= tgs tg ttmtg t t ggtgg g g dt d DTT dt dH DTT dt d H \u03c9\u03c9\u03c9 \u03b8 \u03c9\u03c9 \u03c9 \u03c9 2 2 (3) ( )tgtgtgtgtg DKT \u03c9\u03c9\u03b8 \u2212+= (4) where tH and gH are the inertia constants of the wind turbine the generator rotor, respectively; tD and gD are the damping coefficients of the wind turbine the generator rotor, respectively; t\u03c9 and g\u03c9 are the angle speeds of the wind turbine the generator rotor, respectively, and 21 ttt HHH += ; tgD , tgK , tg\u03b8 and tgT are the shaft damping coefficient, the stiffness coefficient, the twist angle and torque, respectively. Neglecting the damping coefficient and the stiffness coefficient, the drive train is represented by a lumped -mass model [4], [14] as shown in Figure 3. The differential equations of the lumped-mass model are: \u23aa \u23aa \u23a9 \u23aa\u23aa \u23a8 \u23a7 = \u2212\u2212= m m mmg m dt d DTT dt dH \u03c9\u03b8 \u03c9\u03c92 (5) where H , D , m\u03c9 and m\u03b8 are the inertia constant, the damping coefficient, the angle speed and twist angle of the equivalent model, respectively; and gtt HHHH ++= 21 . Setting the positive direction of winding voltage, current and magnetic flux following motor convention, and threephase the voltage equations in d-q frame are [9,15]: \u23aa \u23aa \u23aa \u23aa \u23aa \u23a9 \u23aa\u23aa \u23aa \u23aa \u23aa \u23a8 \u23a7 ++= \u2212+= ++= \u2212+= drsqrqrrqr qrsdrdrrdr dssqsqssqs qssdsdssds s dt diRu s dt diRu dt diRu dt diRu \u03c8\u03c9\u03c8 \u03c8\u03c9\u03c8 \u03c8\u03c9\u03c8 \u03c8\u03c9\u03c8 (6) The corresponding magnetic flux equations are as follows: \u23aa \u23aa \u23a9 \u23aa \u23aa \u23a8 \u23a7 += += += += qrrrqsmqr drrrdsmdr qrmqsssqs drmdsssds iLiL iLiL iLiL iLiL \u03c8 \u03c8 \u03c8 \u03c8 (7) Electromagnetic torque: ( ) dsqsqsdsqrdsdrqsme iiiiiiLT \u03c8\u03c8 \u2212=\u2212= (8) where sR and rR are the stator and rotor resistance, respectively; lsL and lrL are the stator and rotor leakage, respectively; mL is the mutual inductance; lsmss LLL += is the stator self-inductance; lrmrr LLL += is the rotor selfinductance\uff1b u \u3001 i \u3001\u03c8 each winding of the stator and rotor voltage, current and magnetic flux, respectively; s\u03c9 is the synchronous angle speed; s is the rotor slip" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000105_amm.37-38.1462-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000105_amm.37-38.1462-Figure2-1.png", + "caption": "Fig. 2 Schematic of heat transfer boundary condition", + "texts": [ + " Tetrahedron element with 10 nodes is used in messing the geometry model of piston. Messing density is 5 mm for piston and 4 mm for ring carrier. In order to reduce the influence of cooling gallery change in position, the messing density in vicinity of cooling gallery is set to be 1 mm. The finite element model of piston is shown in Fig. 1(b). The finite element model includes 102,724 elements and 155,200 nodes. Material Parameters and Boundary Condition. Schematic of heat transfer boundary condition is shown in Fig. 2, and the heat transfer coefficient is given in Table 1 [6-7]. The mechanical and physical properties of material are sensitive to its temperature. The studied material is ZL8(12Si-2.5Ni-1Mg-1Cu). This metal are commonly used for diesel engine piston, pulley, strap wheel, and other parts which require high heat-durability, low thermal expansion All rights reserved. No part of contents of this paper may be reproduced or transmitted in any form or by any means without the written permission of Trans Tech Publications, www" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003644_c9ja00264b-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003644_c9ja00264b-Figure2-1.png", + "caption": "Fig. 2 A schematic illustration of the axially viewed ICP-OES setup.", + "texts": [ + " conventional torch for axially viewed ICP-OES, the effects of the injector tube internal diameter (i.d.) and power for both torches, and of the intermediate gas for the Fassel torch only, are analyzed. Also, a mathematical model has been developed with the capability of predicting the effects of Na concentration on ion density and signal intensity of other analytes. The model is used to interpret the experimental results of Na interference. The experimental setup used in the work is shown schematically in Fig. 2. The RF generator, sample introduction system, and optical system (optical lens, ber, monochromator, and CCD detector) are identical to those previously used for radially viewed experiments.25 Only the position of the optical lenses was changed, to observe the plasma axially. To reduce selfabsorption and protect the optical lenses from thermal damage and sample deposition, a ow of air was implemented in front of the torch at a suitable distance to divert the tail plume of the plasma. The geometrical and operating parameters for both torches are shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003261_s0263574719000158-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003261_s0263574719000158-Figure2-1.png", + "caption": "Fig. 2. Schematic of coordinate systems.", + "texts": [ + " First, since in reality the SAA is more necessary and important for objects that flies at same attitude, we focus on two-dimensional SAA study and assume that the host UAV and intrude are flying in the same horizontal plane. Second, the motions of the host UAV are measured accurately by its conventional navigation sensors, such as GPS/INS. Third, the vision sensing of intruder is already finished and the directional angle of the intruder with respect to the host UAV is the visual measurement. Fourth, wind influence is not considered. An inertial frame is defined as I(O, X, Y ) (Fig. 2). We assume the body coordinate frame of UAV is H(Ob, xb, yb) with the origin at the centroidal center of the host UAV and the xb-axis pointing forward. The yaw angle \u03d5(t) denotes the rotation from I to H (Fig. 2). The directional angle of the intruder in H is defined as \u03c6(t) and provided by the monocular camera. The forward velocity of the host UAV is denoted as v f (t). In this section, we present the detailed designs of the monocular vision-based SAA approach in avoidance phase. In order to significantly reduce the computation complexity of optimization problem in NMPC and make the controller more practical, the host UAV maintains a constant forward ground speed in the avoidance phase and the speed is denoted as vc, that is, v f (k) = vc, k \u2208N for the kth step in time", + " In addition, the constant velocity assumption is not needed to be strictly satisfied because in filter design, the state model will include noise terms to compensate for system uncertainties. A target tracking filter is used to estimate the relative motion between the intruder and the host UAV. Polar coordinate system has been shown more robust for relative motion estimation using only angle information.30 We define a polar coordinate system with the origin coinciding with that of H and the polar axis paralleling the X -axis of I (Fig. 2). The relative target range and bearing angle of the intruder with respect to the host UAV are denoted as r(k) and \u03b8(k), respectively. As such, \u03b8(k) = \u03d5(k) + \u03c6(k) can be obtained by combining the measurements of the monocular camera and onboard navigation sensors. We then define the state vector of the intruder in the polar coordinate as z(k) = \u23a1 \u23a2\u23a3 z1(k) z2(k) z3(k) z4(k) \u23a4 \u23a5\u23a6 := \u23a1 \u23a2\u23a2\u23a3 \u03b8\u0307 (k) r\u0307 r (k) \u03b8(k) 1 r (k) \u23a4 \u23a5\u23a5\u23a6 \u2208R 4. (10) For an intruder with a constant velocity (see (9)), the relative motion between the intruder and the host UAV can be modeled as30 z(k + 1) = z(k) + T f (k) + v f (k), (11) and f (k) = \u23a1 \u23a2\u23a3 \u22122z1z2 + z4(\u2212ahx(k) cos(z3) + ahy(k) sin(z3)) z2 1 \u2212 z2 2 + z4(\u2212ahx(k) sin(z3) \u2212 ahy(k) cos(z3)) z1 \u2212z2z4 \u23a4 \u23a5\u23a6, (12) where v f (k) \u2208R 4 represents the zero-mean white Gaussian noise for system uncertainties" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003931_s11012-019-01101-4-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003931_s11012-019-01101-4-Figure1-1.png", + "caption": "Fig. 1 General landing gear system representation. Inputs and outputs", + "texts": [ + " _x\u00f0t\u00de \u00bc Fx\u00f0t\u00de \u00feGu\u00f0t\u00de \u00f01\u00de y\u00f0t\u00de \u00bc Hx\u00f0t\u00de \u00f02\u00de The first step in this pseudo feedback strategy is to define the input and output matrices (G and H) considering an algebraic similarity between the system\u2019s physical parameters and the inputs/outputs. The proposed approach considers a pseudo control force Fn\u00f0t\u00de between two structural nodes (two particular degrees of freedom, i.e., n and n 1) written in terms of an incremental stiffness kn such that Fn\u00f0t\u00de \u00bc knDxn\u00f0t\u00de, where Dxn \u00bc xn xn 1. This pseudo control force can be considered in all parts of the landing gear candidates to be modified increasing or decreasing the structural stiffness (see a conceptual example in Fig. 1). Similarly, it is possible to consider pseudo control forces involving the relative velocities, i.e., incremental values of structural damping. Assuming m candidates parts of changing both local stiffness and damping, the pseudo output and input vectors y\u00f0t\u00de and u\u00f0t\u00de are conveniently defined by y\u00f0t\u00de \u00bcfDx1\u00f0t\u00de D _x1\u00f0t\u00de . . . Dxm\u00f0t\u00de D _xm\u00f0t\u00degT 2 R2m 1 \u00f03\u00de u\u00f0t\u00de \u00bcfF1\u00f0t\u00de . . . Fm\u00f0t\u00degT 2 Rm 1 \u00f04\u00de Denoting K 2 Rm 2m a pseudo control gain for an output feedback approach, it is possible to write the classical linear relation u\u00f0t\u00de \u00bc Ky\u00f0t\u00de \u00f05\u00de However, in order to identify the incremental values of stiffness (and damping) to obtain the desired structural landing gear mode, the pseudo matrix of gain is written by imposing null elements to eliminate those do not related to this physics of local modifications of the structural properties" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001575_robot.2010.5509764-Figure15-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001575_robot.2010.5509764-Figure15-1.png", + "caption": "Fig. 15 Active ankle foot orthesis with QuadHelix-Drive", + "texts": [ + " Energy consumption measurements of the robotic arm during a pick-&-place-scenario with a 1 kg object yield to an average value of less than 80 W. Furthermore, the massrelated torque-density of 8.45 Nm/kg for the first DoF and 1.35 Nm/kg for the second DoF of a 2-DoF-module is high compared to other approaches [14]. These are only examples for the high efficiency of the whole actuation system and the possibilities for robotic applications. Fig.14 ISELLA 2 within Fraunhofer IPA testing field The second project is an active ankle foot orthesis powered by a QuadHelix-Drive, which is still under development. Fig. 15 shows the latest version of the concept, which uses one QuadHelix-Drive to do the dorsiflexion and plantarflexion. The QuadHelix-Drive shows a way, how a compact and powerful rope actuator can be used in robotic applications. The key challenges of the DoHelix-Muscle were addressed and a new drive concept was developed. First applications in a lightweight robotic arm and in an active ankle foot orthesis are developed and next steps are an evaluation and a performance measurement. A detailed system analysis, the optimization of key components and a mobile implementation will follow" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002308_s1995080217060051-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002308_s1995080217060051-Figure2-1.png", + "caption": "Fig. 2. Contact patch.", + "texts": [ + " The main difference is that the normal stress distribution is not static as in Contensou model, but is calculated from the Kelvin\u2013Voight model of visco-elastic plane. The model is the following (for details, see [2]): \u2022 We define the spherical coordinates \u03b2 \u2208 (0, 2\u03c0), \u03b8 \u2208 (0, \u03b80) for the points of the sphere P that are lower than the undeformed support plane level: zP = zO \u2212 r cos \u03b8 = r(cos \u03b80 \u2212 cos \u03b8) < 0. The angle \u03b2 is measured from the direction e1 of horizontal velocity uO = vO \u2212 (vO,\u03b3)\u03b3 of the point O (see Fig. 2); \u2022 We calculate vertical infinitesimal force fz = kr(cos \u03b8 \u2212 cos \u03b80) + \u03bdu sin \u03b8 cos \u03b2 \u2212 \u03bdvz, where u = |uO|, vz = (vO,\u03b3), k and \u03bd are the elastic and viscous coefficients of the plane. It can be treated as a force produced by a vertical spring that is in contact with the sphere. We call the domain I = {(\u03b8, \u03b2) : fz > 0} the contact patch; \u2022 For each point P in the contact patch with the radius vector \u03c1P = \u2212\u2212\u2192 OP = \u2212r cos \u03b8\u03b3 + r sin \u03b8 cos \u03b2e1 + r cos \u03b8 sin \u03b2e2, where e2 = [\u03b3, e1], we calculate its velocity vP = vO + [\u03c9,\u03c1P ] and normal to the sphere vector nP = \u2212 sin \u03b8 cos \u03b2e1 \u2212 sin \u03b8 sin \u03b2e2 + cos \u03b8ez" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003269_s42417-019-00086-4-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003269_s42417-019-00086-4-Figure3-1.png", + "caption": "Fig. 3 FE models of (a) the PGJB rotor system and (b) the plain gas film", + "texts": [ + " 1 3 Detailed boundary conditions of the FFSIM are as follows: (1) the rotor shaft and the bearing sleeve are flexible structures; (2) the wall of the gas film model is treated as a flexible wall; (3) the flow condition is laminar and isothermal; (4) the gas inlet is complete self-absorption and the change of the gas inlet and outlet depends on the rotating state of the rotor; (5) the rotor shaft starts its motion from a concentric location within the bearing under action of the gravitational force; and (6) the bearing eccentric location varies with the gravitational force and the rotating speed. FE models of the PGJB rotor system and the plain gas film are shown in Fig.\u00a03, in which the structural hexahedron grid is used. Compared to the unstructured tetrahedral grid, the structured hexahedron grid has higher calculation accuracy and better quality of grid, especially for the gas film clearance flow problem. Rotating process of the PGJB rotor system within the physical time of 0.5\u00a0s from the beginning of rotation is simulated through the FFSIM (Fig.\u00a01). To assure the convergence of the numerical simulations, the iteration time step should be small enough and it is set to be 0", + "\u00a04 we can see that, obviously, as a whole, the frequency response curve calculated by the FFSIM agrees very well with that of the experiment given in Ref. [50]. This, to some extent, indicates that the FFSIM could accurately simulate the nonlinear dynamics of the gas-bearing rotor system. Accurate Nonlinear Dynamic Characteristics of\u00a0the\u00a0PGJB Rotor System Obtained by\u00a0the\u00a0FFSIM Since in the previous section the accuracy of the FFSIM has been verified by the experimental results of Ref. [50], in this section, accurate nonlinear dynamic characteristics of the single-span single-disc rotor system supported by two PGJB (Fig.\u00a03) are systematically and deeply studied by the FFSIM. Bifurcation analysis, effects of rotor mass and rotating speed on shaft orbits, phase portraits and frequency response curves of this PGJB rotor system are investigated, respectively. To show the exact characteristics of our numerical results, under the same rotor system and boundary conditions, numerical results of the FFSIM are compared with those of the simplified two-dimensional analytical method commonly seen in the existed literature [33, 34, 41, 43, 50], which establishes the nonlinear dynamic equations of the PGJB rotor system by simplifying the pressure gas film of the PGJB as dampers and springs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000523_978-94-007-0020-8_28-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000523_978-94-007-0020-8_28-Figure2-1.png", + "caption": "Fig. 2 Deformation in balls and races", + "texts": [ + " The foregoing procedure using analogous plane treatment simplifies the procedure when it is needed to add further details of the system such as internal damping and bearing stiffness, damping and support flexibility. More details of the formulation are given in Hashish and Sankar [15]. Assuming perfect rolling of balls on the races, the varying compliance frequency of a ball bearing is given by: !vc D !rotor BN, where BN D {(Ri=.Ri C Ro)} Nb the BN number depends on the specifications and dimensions of the bearing. In Fig. 2, .x; y/ denote the displaced position .O0/ of the centre of the inner race. From simple geometrical analysis, the expression for elastic deformation .\u0131. i // is, \u0131. i / D .xcos i C ysin i 0/ (5) In case of \u0131. i / > 0, ball at angular position i is loaded giving rise to restoring force with nonlinear characteristics because of Hertzian contact [16]. If \u0131. i/ < 0, the ball is not in the load zone . LZ/, and restoring force from the ball is set to zero. The total restoring force is sum of restoring forces from each of the balls" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001063_robio.2011.6181298-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001063_robio.2011.6181298-Figure2-1.png", + "caption": "Fig. 2. Methods for moving the center of gravity", + "texts": [ + " The ZMP expresses the robot\u2019s balance as a force distribution on each foot\u2019s sole. In the following section we describe the PI control with ZMP on moving the center of gravity. In section III we present a forward, linear, and parallel\u2013to\u2013the\u2013floor sole movement for a humanoid robot. In section IV we shortly introduce our research platform and the ZMP on the soles of the feet. In section V we describe experimental procedures and results from section II and III. Finally, in section VI we summarize the presented methods and results. Fig. 2 shows the methods of controlling the displacement of the center of gravity used by Uda [10] and in this research. Most existing humanoid robots perform a lateral displacement of their center of gravity by controlling the rotation of the X axis of hips and ankles at the same time, as shown in Fig. 2(a). Nevertheless, this method of controlling two joints at the same time has a high lateral inertia that causes big variations in the ZMP, making control very difficult. The method where each joint moves discretely and successively has a smaller inertia and less variations in ZMP. Besides, when a joint on one side is moving to its desired position, the inertial force and the ZMP will variations decrease on the other side and since the angle difference between the hip joint and the ankle is big, the sole tends to get separated from the floor surface. In a similar way, if a variation is introduced in the displacement of each joint, the sole will become separated from the floor. In this research, after controlling the ankle joint with the ZMP in each cycle, only the angle that caused a displacement in the ankle will be used to control the hip joint, as shown in Fig. 2(b). It is possible to move the center of gravity using ZMP feedback with this method, as it reduces the variation in the ZMP. In addition, the center of gravity PI controlled to provide the ZMP feedback. Let us consider p(t) as the current ZMP, p0 as the target ZMP, E(t) as the variation between them, and \u03c8\u0301roll(t) as the manipulation variable of the rotation over the X axis needed to perform the control, as shown in Fig. 2(b). The center of gravity moves using PI control of the ankle joint, as described in the block diagram in Fig. 3. The control equation can be expressed as E(t) = po \u2212 p(t), \u03c8\u0301roll(t) = KPE(t) +KI \u222b t 0 E(t)dt. (1) In this research, we consider the proportional gain (KP ) as 0.0125, the integral gain (KI ) as 0.00025, and the sampling time as 0.001 [s]. III. INVERSE KINEMATICS OF LEG MECHANISM Within the present research, we implemented a forward, linear, and parallel\u2013to\u2013the\u2013floor sole movement for a humanoid robot", + " The \u03a3FR coordinate system is defined using the hip joint as the origin, and considers the forward direction to be the X axis, the left direction to be the Y axis, and the vertical line directed at the floor to be the Z axis, as shown in Fig. 4. We consider the position of the humanoid robot\u2019s hip joint as p1, the position of the knee as p2, the position of the ankle as p3, and the position of the foot as p4. The length of the humanoid robot\u2019s thigh is l1, the length of the calf is l2, and the distance from the ankle joint to the sole is l3. As shown in Fig. 2, the rotation of the hip joint and the ankle over the X axis are given by \u03c81 and \u03c83, respectively, and observing this from the XZ plane it is possible to obtain the line segments as the product of cos\u03c81, l1, and l2, respectively. We donate the rotation angle over the Y axis as \u03b81 for the hip joint, \u03b82 for the knee, and \u03b83 for the ankle joint. Considering a sliding walk, the rotation over the x and y axes of p4 will always be 0 [rad]. Using forward kinematics, p4 is given by p4 = \u23a1 \u23a3 x4 y4 z4 \u23a4 \u23a6 = \u23a1 \u23a3 l1c\u03c81 s\u03b81 + l2c\u03c81 s\u03b81\u03b82 (l1c\u03c81c\u03b81 + l2c\u03c81c\u03b81\u03b82)t\u03c81 l1c\u03c81c\u03b81 + l2c\u03c81c\u03b81\u03b82 + l3 \u23a4 \u23a6 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000010_13506501jet600-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000010_13506501jet600-Figure1-1.png", + "caption": "Fig. 1 Bushing and oil supply groove", + "texts": [ + " It is solved using the multi-level multi-grid technique [12]. Then, based on this model, results are obtained for three different bushing designs: first, a cylindrical bushing, which will be used as reference, second an \u2018hourglass\u2019-shaped bushing with either circular or parabolic profile, and third a bushing with conically chamfered edges. The bushing of the journal bearing is modelled as a hollow cylinder with inner radius RBi, external radius RBo, and length L. The position and the dimensions of the supply groove are given in Fig. 1. The global coordinate system (xG, yG, zG) is defined in such a way that the applied load is directed towards the negative side of the zG-axis. The shaft is considered misaligned in the (yG, zG) plane. Its position is defined by the tilting angle \u03b3 about the xG-axis. The attitude angle \u03d5J at the mid-plane, the eccentricity e, and the direction of the rotational speed \u03c9J of the journal are shown in Fig. 2. Assuming that the radius of the journal is very large compared to the film thickness, the curvature may be neglected and the journal surface can be developed onto a flat surface as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002138_gt2017-63815-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002138_gt2017-63815-Figure5-1.png", + "caption": "Figure 5. HIGH SPEED VISUALIZATION IMAGES DEPICTING TRANSITION FROM DIRECT-DROP FORMATION TO PRIMARY-LIGAMENTS/LAMINAR SHEET AND FINALLY TO RUFFLED SHEET/SECONDARY-LIGAMENTS.", + "texts": [ + " Locations \u20181\u2019 and \u20182\u2019 are the sharp edges of the bearing cage. The stationary surface of the bearing chamber is marked as surface \u20183\u2019. Oil supplied to the bearing is delivered via three sets of holes at axial locations under the cage and inner race. Some of the oil delivered to the location under the cage at the front of the rig exits through the gap between the cage and the inner race. A subsidiary test done with the front oil feed off shows that a negligible amount of oil from the other two feed locations exits at the front. Figure 5 shows a series of still images obtained with the high speed camera illustrating the oil shedding and break up behaviour as shaft speed is increased. The different oil break- up mechanisms are classified as (a) direct-drop formation, (b) primary-ligaments/laminar sheet and (c) ruffled sheet/secondary-ligaments. The transition from one mode to another is brought about by increasing shaft rotational speed. Each of these modes is explained below describing the instability mechanism responsible for its formation. a. Direct-drop formation At a low shaft speed of 1200 rpm (Figure 5a), the oil exiting between cage and inner race remains attached to the cage-surface between locations \u20181\u2019 and \u20182\u2019 (shown by the dotted line in Figure 4) due to surface tension effects. An oil torus is formed around location \u20182\u2019 and due to centrifugal action, drops are formed at the large diameter bulges of this torus which are dispersed 4 Copyright \u00a9 2017 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 08/23/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use circumferentially outwards as oil droplets. The term \u2018bulges\u2019 is used to represent the \u2018pearl-like\u2019 blob of fluid breaking-up from the edge of oil sheet. This can be seen in Figure 5a. This mode of oil disintegration is direct-drop formation, similar to drop formation mode in rotary cup oil shedding (Figure 1a). Although understanding droplet generation from the bearing cage at low shaft rotational speed may not be as important for aeroengine bearing chambers as high speed behaviour, it is useful from a CFD validation perspective. It is also important from an academic point of view to understand the fundamental fluid mechanics over the complete parametric ranges of bearing chamber operation. b. Primary-ligaments/laminar sheet As the shaft speed is increased to 1800 rpm (Figure 5b), centrifugal force overcomes the surface tension and viscous forces at location \u20181\u2019 to form an oil torus at this location (instead of flowing along the surface to location \u20182\u2019). Two breakup/shedding mechanisms are intermittently observed at this shaft speed. In the first, bulges from the torus become thin to form \u2018primary ligaments\u2019. By ligaments, we mean thread-like structures whose aspect ratio (length to thickness ratio) is >>1. This mode is similar to ligament break-up mode depicted in Figure 1b", + " In particular, each primary ligament granulates to an array of large main droplets from which secondary/satellite droplets are subsequently formed. The thinning of the primary ligaments to eventually form a conglomerate of droplets is driven by the aim to reduce system surface energy by reducing the surface area. A system always prefers to be in the minimum energy state in order to attain equilibrium. At the same shaft speed (of 1800 rpm) another oil break-up mechanism is observed where the torus forms into a \u2018laminar sheet\u2019 for some part of shaft rotational cycle (Figure 5c). The term \u2018laminar sheet\u2019 is selected because the sheet is quite smooth and relatively undisturbed. These sheets extend radially from the cage-edge till the equilibrium radius. At this radius, the centrifugal force equals the contracting action of surface tension. At this circular periphery a thick rim of oil is produced which further breaks down to threads/droplets. This type of disintegration mechanism has been termed \u2018rim disintegration\u2019 by Frazer [7]. A prominent characteristic of such sheets is that the sheet is not ruffled anywhere throughout its extent", + " Laminar sheet break-up is witnessed when the cage surface is fully wet whereas primary ligaments are observed when the cage surface is partially wet. In a recent study by Peng et al. [11] of liquid on a spinning disk, two break-up modes were also intermittently observed (namely direct drop and ligament break-up modes) and the authors concluded that the abovementioned wettability issue was responsible. As shaft speed is further increased to 2600 rpm the observed behavior is similar to 1800 rpm but the primary ligaments are observed to grow in number while simultaneously thinning to form fine threads (compare the ligaments in Figure 5b to those in Figure 5d). The laminar sheet is still intermittently observed. 5 Copyright \u00a9 2017 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 08/23/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use c. Ruffled sheet/secondary ligaments With a further increase in shaft speed, the proportion of time for which there is a sheet present increases until at 4500 rpm there is complete disappearance of primary ligaments and the sheet appears as illustrated by Figure 5e. The sheet break-up mechanism observed at this critical speed is fundamentally different from the laminar sheet formation of Figure 5c. In particular, at higher speed the shearing effect of the gas surrounding the radially propagating oil sheet superimposes exponentially growing waves extending from cage-edge and which are perpendicular to the liquid flowlines. These waves are the same as those observed by Frazer et al. [7] in the spinning cup study and can be either sinuous or dilatational or a combination of both. Frazer et al. [7] termed this type of sheet break-up mechanism \u2018wave disintegration\u2019. The oil sheet disintegrates through half wavelengths of oil being torn off", + " These secondary ligaments are torn-off from the edge of the sheet; they are formed due to wave disintegration followed by the subsequent formation of unstable threads. In contrast, primary ligaments bulge from the 6 Copyright \u00a9 2017 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 08/23/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use liquid torus formed at the edge of the rim which acts as controlling surface. As a result primary ligaments are smooth, almost equally spaced and very stable (compare Figure 5b and Figure 5e). Another difference between primary and secondary ligaments lies in the nature of the secondary/satellite droplets that are produced from the large main droplet that is torn off from the edges of each of these ligaments. The satellite droplets produced from parent droplet produced from secondary ligaments represent a bimodal/multimodal type of distribution where two or more distinct ranges of droplet sizes are observed. This behaviour has was observed by Wang et al [12] in their spinning disk study", + " As a result, for primary ligament break up, analytical expressions and empirical correlations can be developed to determine the droplet size using ligament number, diameter, and capillary wavelength as has been done by Kamiya et al. [10] and Wang et al. [12]. However, the less repeatable nature of the secondary droplets produced from secondary ligaments poses problems in deriving such empirical correlations. Within the shaft speeds tested in the present experimental work, the highest speed achieved was 7000 rpm. Secondary ligaments were evident at this speed also (Figure 5f). However, the sheet diminished in extent at this speed as compared to 4500 rpm (Figure 5e). Higher shaft rotational speeds have to be achieved to enable further study of the dynamics of these secondary ligaments and this is planned for future work. Comparison to prior published research Although, transition from direct-drop to ligaments and different types of sheet break-up has been reported in the past (for example in Hinze & Milborn [5], Frazer [7] and Frost [8]) the present experimental work is novel due to the fact that these observations are reported for the first time in the context of an aeroengine bearing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003075_tasc.2019.2891534-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003075_tasc.2019.2891534-Figure1-1.png", + "caption": "Fig. 1 Structure of the motor", + "texts": [ + " Finally, the temperature rise test of the motor was completed on the experimental platform, and the test results were compared with the simulation results to verify the accuracy of the analysis method. P 1051-8223 (c) 2018 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. The research object of this paper is a PMSM for industrial robots with an output power of 400W. Some parameters of the motor are shown in Table \u2160. The structure of the motor is shown in Fig. 1. The stator adopts concentrated winding with slots number of 12. The rotor is a permanent magnetic ring with poles number of 10. The simulation parameters is shown in Table \u2161. TABLE \u2161 SIMULATION PARAMETERS Design parameter Value DC voltage (V) 297 Switch frequency (kHz) 10 Rated voltage (V) 220 Rated current (A) 2.4 Rated speed (rpm) 3000 Number of poles 10 Rotor inertia (kgm2) 2.16*10-5 Ld, Lq (mH) 7.5 The diagram of vector control system is shown in Fig. 2. The system uses a double closed loop structure of current loop and speed loop" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001893_insi.2017.59.3.143-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001893_insi.2017.59.3.143-Figure3-1.png", + "caption": "Figure 3. Shape of the magnetic field concentrator", + "texts": [ + " The magnetic focus detection method relies on magnetic concentrators to collect the diverging magnetic field in a homogenised space and lead it onto the testing path. Thereby, the testing element can easily measure the variation in the magnetic field. The magnetic concentrator is made of a material with high magnetic conductivity. The structural design of a magnetic concentrator is of great importance if the magnetic concentrator is to focus the magnetic field in a proper manner. The magnetic concentrator structure shown in Figure 3 offers better magnetic focusing performance in wire ropes with small diameters[14]. The magnetic concentrator features a semi-circular structure divided Figure 1. Principle and structure of magnetic field balance detection 144 Insight \u2022 Vol 59 \u2022 No 3 \u2022 March 2017 Insight \u2022 Vol 59 \u2022 No 3 \u2022 March 2017 145 into upper and lower parts and it encases the wire rope along its periphery. Two protruding platforms, which are placed opposite each other, are placed on one side of the magnetic concentrator. A magnetic semi-circular ring is placed between the platforms and the detection elements are placed between the semi-circular magnetic concentrators and the semi-circular magnetic ring" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000209_tac.2009.2020643-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000209_tac.2009.2020643-Figure6-1.png", + "caption": "Fig. 6. .", + "texts": [ + " It is easy to know that the criterion function changes its sign on any control curve . Hence, the system (7) is globally controllable. Case 5: The eigenvalues of are two conjugate pure imaginary numbers. We assume , , , . For any control curve of (7), if does not pass through the origin, then there is a point in the outer side of (namely, the side which does not include the origin, see Definition 4.1 in Section IV). Since the trajectory of the system passing through is a circle and centered with the origin, must pass through twice at least as shown in Fig. 6. This implies that the criterion function must change its sign on the control curve . If passes through the origin, two ends of must pass through any nonzero trajectory of , since the two ends of extend to infinity. This also implies that the criterion function must change its sign on the control curve . Hence, the system (7) is globally controllable. Now, aside from an exceptive case, all cases that is uniformly controllable have already been discussed. This completes the proof of Theorem 2.1. IV" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002109_1350650117723484-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002109_1350650117723484-Figure5-1.png", + "caption": "Figure 5. Compressive stress distribution.", + "texts": [ + " Therefore, new values of displacements at kth iteration will be presented as k\u00fe1x k\u00fe1y k\u00fe1z k\u00fe1x k\u00fe1y 8>>>>>>< >>>>>>: 9>>>>>>= >>>>>>; \u00bc kx ky kz kx ky 8>>>>>>< >>>>>>: 9>>>>>>= >>>>>>; \u00fe k\u00fe1x kx k\u00fe1y ky k\u00fe1z kz k\u00fe1x kx k\u00fe1y ky 8>>>>>>< >>>>>>: 9>>>>>>= >>>>>>; \u00f020\u00de The result converges when the increment of displacements is less than the criteria value. After that, contact angles and contact forces are readily acquired. Contact stress and stiffness of the bearing Bulk failure of rolling members is generally not a significant factor in a rolling bearing design; however, the destruction of the rolling surfaces is a significant factor.6 Hence, compressive stress occurring on the contact area of the ball and raceways needs to be considered. Figure 5 illustrates an ellipse contact geometry and compressive stress distribution on the contact region. According to Harris,6 the compressive stress at any point in the ellipse contact region can be determined by \u00bc 3Q 2 ab 1 x a 2 y b 2 1=2 \u00f021\u00de where a and b are the semi-major and semi-minor axes of the ellipse, respectively. Clearly, the maximum compressive stress is achieved at the center \u00f0x \u00bc 0, y \u00bc 0\u00de of the elliptical contact area max \u00bc 3Q 2 ab \u00f022\u00de By knowing the semi-major and semi-minor axes of the ellipse and contact forces, the surface compressive stress can be easily pointed out" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002897_irsec.2017.8477423-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002897_irsec.2017.8477423-Figure5-1.png", + "caption": "Figure 5. Dimensions of studied 8/6 Switched Reluctance Machine", + "texts": [ + " In the above expressions \u03c6, i, R, L, \u03c9, \u03b8, V, T, W(\u03b8, i) are flux linkage, phase current, phase r\u00e9sistance, phase inductance, rotor speed, rotor position angle, phase voltage, phase torque, and co-energy, respectively. Because of the magnetic saturation effect, the flux in the stator phases varies according to the rotor position \u03b8 and the current of each phase. III. STRUCTURE OF THE STUDIED MACHINE The structure of studied machine is a four phase DSSRM 8/6. The parameter dimensions are given by Fig. 5. 8 designates the number of poles in the stator (Ns) and 6 is the number of poles in the rotor (Nr), the choice of Ns and Nr is important since they have significant implications on the torque. The number of turns in windings is 260 and the material of the SRM stator and rotor is steel M36 with nonlinear B-H characteristic, the B-H curve is shown in Fig. 6. IV. NONLINEAR FINITE ELEMENT MODELING OF SRM For magnetostatic model two types of formulations are considered: the scalar model, and the vector model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000661_raad.2010.5524601-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000661_raad.2010.5524601-Figure5-1.png", + "caption": "Figure 5. a) Kinematic schematic of the robot b) Kinematic schematic of the cell", + "texts": [ + " Taking into account the various constraints led us to explore various solutions, showing that the cutting of left and right half carcasses is not possible without repositioning and reorienting them (Fig. 2). The second workspace is dedicated to ham boning. Most cutting operations call for the bones to be accessible, which requires the integration of an external axis. The definition of robotic tasks led us to use a vertical position of the ham, maintained between the tibia-fibula and femur bones. Geometric modeling is necessary for simulation and optimization. A TCS (Traveling Coordinate System) [10] type model was created (Fig. 5a)). This model shows the six successive transformations in the order of the axes of rotation imposed by the TCS model. Offsets, corresponding to the definition of the tool from the center of connection 6, should be added to the final element. They are represented by three parameters atool, btool and tool, which define the position and orientation of the Tool Center Point (TCP). The geometrical model is represented by the following transformations: tooltooltooltool R bTaTR RaTR bTRbT RbTR ZYXOZYXO ZYXOZYXOZYXO ZYXOZYXOZYXO ZYXOZYXOZYXO ZYXOZYXOZYXO out y out out y out xx zyy zxy xyz \u23af\u23af\u23af\u23af \u2192\u23af \u23af\u23af\u23af \u2192\u23af\u23af\u23af\u23af \u2192\u23af\u23af\u23af\u23af \u2192\u23af \u23af\u23af\u23af \u2192\u23af\u23af\u23af\u23af \u2192\u23af\u23af\u23af\u23af \u2192\u23af \u23af\u23af \u2192\u23af\u23af\u23af\u23af \u2192\u23af\u23af\u23af\u23af \u2192\u23af \u23af\u23af\u23af \u2192\u23af\u23af\u23af \u2192\u23af\u23af\u23af\u23af \u2192\u23af )( '6'6'6'6 )( \"6\"6\"6\"6 )( 6666 )( 5555 )( '4'4'4'4 )( 4444 )( '3'3'3'3 )( 3333 )( '2'2'2'2 )( 2222 )( '1'1'1'1 )( 1111 )( 0000 '6 '6\"6\"66656'5 545'44'44434'3 3'33323'22'22 212'11'111001 \u03b8 \u03d5 \u03d5\u03d5 \u03d5 \u03d5\u03d5 (1) This model allows us to write the homogeneous operators used in the calculation of the effector coordinates based on joint variables. The complete cell, including the robot and turntable (external axis), is modeled in the same way as with the TCS method. The external axis adds a joint variable 0e (Fig. 5b)). The new Direct Kinematic Transformation (DKT) solution is direct by adding this axis in series. In order not to change the notations of the other joints, the joint value is denoted q7 although it is placed at the beginning of the driveline. The relative position of the turntable is described by an architecture parameter which must be also be integrated into the model. The DKT of the complete cell is obtained by multiplying together the different homogeneous operators. The Inverse Kinematic Transformation (IKT) is determined from the geometric and vector relationships of its architecture" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001335_pvp2010-25811-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001335_pvp2010-25811-Figure8-1.png", + "caption": "Fig. 8 Schematic diagram for the relationship of underhead bearing friction torque bT , thread friction torque tT and pitch torque pT", + "texts": [ + " The actual bearing friction torque, which is approximately equal to the twisting torque in the bolt, will quickly be reduced to a level equal to the pitch torque pT in just one cycle. When self-loosening is about to occur, the thread friction torque will have the same direction as the bearing friction torque, but opposite to the pitch torque pT . The relationship is tpb TTT \u2032\u2212=\u2032 (1) where bp F p T \u03c02 = is pitch torque component, bT \u2032 is the actual underhead bearing friction torque component, tT \u2032 is actual thread friction torque as shown in Fig.8, p is thread pitch, and bF is the bolt tension. 3.2 Underhead Friction and Bearing Surface The underhead friction force is always opposite to the direction of the relative velocity vector on the contact surface. The schematic diagram of the friction force distribution and its resultant shear force bsF and the underhead bearing friction moment bT is shown in Fig.9. Reference [16] gives the following relationships about sliding bearing frictional torque bT and the transverse bearing friction shear force bsF as functions of newly introduced underhead sliding speed-torotational speed ratio bbb v \u03c9\u03b7 /1= \u222b\u222b ++ +=== \u03c0 \u03b8\u03b7\u03b7 \u03b8\u03b8\u03b7 \u00b5 \u03b7 2 0 22 2 0 1 sin2 )sin( )( rr dr drr q T fR bb b r rbb b bbTb e i (2) \u222b \u222b ++ +=== e i r r bb b bb bs bbFb rr dr rdr q F fR \u03c0 \u03b8\u03b7\u03b7 \u03b8\u03b8\u03b7 \u00b5 \u03b7 2 0 220 3 sin2 )sin( )( (3) where 1bv is a relative translation velocity (along the x direction) between the bolt head and the joint bearing surface, 0bq is the average pressure on the contact bearing surface, er and ir are the external and internal radii of the bearing contact area, respectively, r is the radius and \u03b8 is the angle in polar coordinate system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002307_012055-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002307_012055-Figure3-1.png", + "caption": "Figure 3. Block diagram of the experimental setup.", + "texts": [ + " The accelerometer BC 110 transmits the readings to the signal amplifier, which, amplifying the signal, transmits further along the circuit to the analog-to-digital converter, which transmits the signal further to the personal computer. The thermometer is located closest to the shell, but does not touch it, transmits the readings on the software \"Zet-lab\", where in the compartment with the data of the oscillations the dependence of some parameters on others is reflected, in the real time. Software \"Zetlab\" allows you to display and record fluctuations in the real time. The heating element creates a temperature gradient. A block diagram of the experimental setup for carrying out the experiment is shown in Figure. 3. 3 The experiment is aimed at studying the natural oscillations of a plate in \"rest\" and revealing the dependencies of forced and natural oscillations on the effect of an attached mass or mass system. Forced oscillations, going into their own oscillations, were set by impact with a test hammer AU03. Also there was a contactless sensor for measuring the vibration of the plate, not shown in the diagram. This sensor is a checker, serves to check and reject erroneous data of the accelerometer BC110" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001232_tmag.2012.2198203-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001232_tmag.2012.2198203-Figure6-1.png", + "caption": "Fig. 6. Prototype of two-DOF resonant actuator.", + "texts": [ + " The position of mover is calculated by the time step. Fig. 4 shows the flowchart for this coupled analysis. The vector control is taken into consideration in this analysis. Fig. 5 shows the FEM model without air regions. The analyzed region is of the whole region because of the symmetry. The number of tetrahedron elements, edges, and unknown variables are about 306 400, 367 200, and 347 800, respectively. Table I shows the analysis conditions. The number of steps is 9900, time division is 100 s, and total CPU time is about 240 hours. Fig. 6 shows the prototype. As described at Section II-A, the -axis is supported with a flat spring and the -axis is supported by the coil spring and the mover can be moved independently on each axis. This prototype is assumed to move with a range of mm in the -direction and mm in the -direction, respectively. In the experiment, the actuator is controlled by a microcontroller. Fig. 7 shows the analyzed and measured results of the amplitude when the actuator was operated to resonate in the axis. As can be seen, both results show a good agreement" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002801_s40997-018-0239-9-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002801_s40997-018-0239-9-Figure1-1.png", + "caption": "Fig. 1 Drive coordinate system", + "texts": [ + " Combined with aforesaid studies of the theory, this paper proposes the calculation method of tooth width of EHCF gear and the influencing factors and variation of undercutting radius, pointing out that radius and tooth width are analyzed by changing the different parameters of EHCF gear pair. Based on the motion relationship of EHCF gear, the relative motion of pitch curves of gear pair at point P is pure rolling during meshing process, so the coordinate system of gear pair is established as in Fig. 1 (Kang et al. 2012). As shown in Fig. 1, O1 X1Y1Z1, O2 X2Y2Z2 and O3 X3Y3Z3 are coordinate systems which are rigidly connected with the frame of the cutting machine. Coordinates O0 1 X0 1Y 0 1Z 0 1, O0 2 X0 2Y 0 2Z 0 2 and O0 3 X0 3Y 0 3Z 0 3 are rigidly connected with the HNC gear, the fictitious HCF gear and the EHCF gear, respectively. The distance between plane Y2O2Z2 and plane Y1O1Z1 is R. e is the eccentricity of EHCF gear. According to the motion relationship between the EHCF gear and the HNC gear, the HNC gear rotates around axis O1Z1 with an angular velocity of w1 in clockwise direction, and the HNC gear moves around axis O1Z1 with a velocity of vs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001046_wcica.2012.6358389-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001046_wcica.2012.6358389-Figure3-1.png", + "caption": "Fig. 3: The cross section force of the flying boat", + "texts": [ + " The forces and moments in Eq(1) will be estimated in the following sections. III. THE FORCES AND MOMENTS ACTING ON THE FLYING BOAT When the flying boat takes off from water, it is mainly carried by water dynamic forces and aerodynamic forces. As its speed increases, its aerodynamic forces increase gradually, and the water forces decrease to zero when it leaves water. It is needed to define a two-dimensional cross section for any strip method, and the cross sections of the flying boat are defined as showed in Fig.3. According to the research works of Karman (1929) and Wagener (1932), the forces acting on an arbitrary cross section can be divided into three components: hydrodynamic force due to the change of fluid momentum, viscous lift force associated with cross flow drag and buoyancy force related to displacement volume of the section. Added mass The sectional added mass sam at any transverse section depends on its geometric shape and its half beam length sb , and can be estimated by Eq(2): 22sa a sm k b (2) where is the added mass coefficient" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001695_jae-2011-1356-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001695_jae-2011-1356-Figure1-1.png", + "caption": "Fig. 1. Proposed meal assistance robot.", + "texts": [ + " In order to solve these problems, we develop a new meal assistance robot having the following features: first, our robot is handled using acts of user\u2019s eyes. The mechanism is called eye-interface. Second, we design the robot so that the arms move in orthogonal coordinate axes. Third, we use ultrasonic motors (USMs) so that the robot has electromagnetic compatibility. In Section 2, we introduce mechanical structure and a control method of our robot. In Section 3, we propose the eye-interface. In Section 4, we demonstrate experiment results. Section 5 concludes this paper. Figure 1 shows a meal assistance robot proposed in this paper. In this section, we show the mechanical structure and a control method of the robot. \u2217Corresponding author: Kanya Tanaka, Graduate school of Science and Engineering, Yamaguchi University, 2-16-1, Tokiwadai, Ube 755-8611, Japan. E-mail: ktanaka@yamaguchi-u.ac.jp. 1383-5416/11/$27.50 2011 \u2013 IOS Press and the authors. All rights reserved The robot has four axes to move a plate, a pusher, a shutter, and a spoon (see Fig. 2). The plate is a part for putting foods" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000294_icsmc.2009.5346897-Figure19-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000294_icsmc.2009.5346897-Figure19-1.png", + "caption": "Fig. 19. A bicycle model and its coordinate systems and parameters on the lateral plane.", + "texts": [ + " 11) The equation of motion being constrainted is projected on the space which makes the column vector of D a base vector.The equation of motion of only the flexibility after being constrainted is obtained by carrying out coordinates conversion of the ingredient vector. ( DT MDq\u0308 + DT MD\u0307q\u0307 = DT h ) A model figure, the Projection Method of a parameter and the constraint conditions which become important in the modeling of the bicycle in 3-dimensional space are shown below. B. A model figure and the account method of a parameter. The model of the bicycle is shown in Fig.18 and Fig.19. Parameters of the bicycle is shown in Tab.III. \u2022 The relation between the arc by the front wheel and the track by the rear wheel is rF \u03c8F sin(\u03c6F \u2212 \u03c6R) = lb\u03c6R. (2) \u2022 The relation between inclination of the front wheel and the rear wheel is rR\u03c9xR cos \u03b8R =rF \u03c9xF cos \u03b8F cos(\u03c6F \u2212 \u03c6R). (3) The constraint conditions between the front wheel and the handle are held as follows: \u2022 The positional constraint that the handle is connected to the front wheel is xH = Rz(\u03c6H)Rx(\u03b8H)Ry(\u03c8H) \u23a1 \u23a3 \u2212rh 0 lh \u23a4 \u23a6 + xF , (4) where xF = [ xF yF zF ]T is the vector of COG of the front wheel and, xH = [ xH yH zH ]T is the vector of COG of the handle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003836_tie.2019.2950863-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003836_tie.2019.2950863-Figure2-1.png", + "caption": "Fig. 2: Different forces, tensions and motions on the IMs.", + "texts": [ + ".) (3) where fL and f are the mechanical looseness and voltage supply frequencies, respectively. Also, p and \u03b3 are the integer numbers. In addition, s is the rotor slip. In order to model and analyse the IM , it is necessary to describe the effective components on the rotor motion in different operating conditions. In this section, assuming that the air-gap function is independent of the rotor depth, the governing equations on the dynamics of the rotor motion with TDF are derived. According to Fig. 2(a), the position vector of the rotor mass center rG in the TDF along ro, \u03b8r and \u03d5s directions can be declared by rG =(ro cos\u03d5s + e cos \u03b8r)ax+(ro sin\u03d5s + e sin \u03b8r)ay (4) 0278-0046 (c) 2019 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. where \u03d5s and \u03b8r are the minimum air-gap and first rotor bar angles with respected to the stator reference. Also, e, ax and ay are the distance between the rotor and mass centers and the unit vectors along with the positive directions of x and y-axes, respectively", + " The radial and rotational motions of the rotor which are subject to the EM force and torques influenced by the air-gap variation between the rotor and stator. Using EulerLagrange equation, the governing equations for the radial and rotational motions of the rotor can be derived as [17] d dt \u2202Et \u2202q\u0307k \u2212 \u2202 (Et \u2212 Ev) \u2202qk = Qqk \u2212 \u2202Ef \u2202q\u0307k (10) where, qk is the generalized coordinates, ro, \u03b8r, \u03d5s. Also, Qqk is the EM force or torques from variations of the magnetic energy Em in the all directions as Qqk = \u2202Em \u2202qk , qk = ro, \u03b8r&\u03d5s (11) Regarding to Fig. 2(b), it can be concluded that Qro , Q\u03d5s and Q\u03b8r are equal to the radial force fr and the torques of the \u03c4\u03d5s and \u03c4\u03b8r , respectively. By substituting the kinetic, potential and loss energies in (7)-(9) into Euler-Lagrange equation in (10), the TDF dynamic of the rotor motions can be expressed as mr\u0308o + cr\u0307o + krro = fr \u2212mg sin\u03d5s +mro\u03d5\u0307 2 s+ +me(\u03b8\u03072r cos\u03d5r \u2212 \u03b8\u0308rsin\u03d5r) (12) (me2+J\u03b8r )\u03b8\u0308r + c\u03b8r \u03b8\u0307r=\u03c4\u03b8r \u2212me(g cos\u03d5r \u2212 r\u0308o sin\u03d5r)+ +mero(\u03d5\u0307 2 s sin\u03d5r \u2212 \u03d5\u0308s cos\u03d5r)\u2212 2mer\u0307o\u03d5\u0307s cos\u03d5r (13) mr2o\u03d5\u0308s+c\u03d5s \u03d5\u0307s = \u03c4\u03d5s \u2212 2mror\u0307o\u03d5\u0307s \u2212mgro cos\u03d5s\u2212 \u2212mero(\u03b8\u0308r cos\u03d5r \u2212 \u03b8\u03072r sin\u03d5r) (14) By simultaneously solving mechanical equations of (12)-(14) and EM force and torques, it is possible to identify the instantaneous position of the rotor in the machine air-gap. In this section, various types of the EM force and torques on the rotor are introduced to the identification of the EM force and torques imposed to the IMs. Fig. 2(b) shows the IM state in a differential variation on the rotor motion. In this figure, Os, OR and \u03d5s are centers of the stator and the rotor and the minimum air-gap angle, respectively. As shown in Fig. 2(b), the EM radial force fr moves the rotor on the minimum air-gap direction with \u03b4ro displacement. Differential variation of the minimum air-gap angle \u03b4\u03d5s is produced by EM torque \u03c4\u03d5s in the direction of \u03d5s (around stator center). Also, the EM torque on the rotor axis (\u03c4\u03b8r ) differentially turns the rotor at the fixed minimum air-gap position with \u03b4\u03b8r. In general, the stored magnetic field energy Em result in this movement is Em = 1 2 iTL (\u03d5s, \u03b8r, ro) i (15) where \u201cT\u201d denotes the transpose of the matrices and L(\u03d5s, \u03b8r, ro) is the inductance matrix which is defined as L (\u03d5s, \u03b8r, ro)= ( Lss (\u03d5s, ro)) Lsr (\u03d5s, \u03b8r, ro) Lrs (\u03d5s, \u03b8r, ro) Lrr (\u03b8r, ro) ) (16) In this matrix, Lss (\u03d5s, ro) is the stator self-inductance matrix with order 3 \u00d7 3 ", + " In addition, i is the current vector which is defined as i = ( iTs iTr )T (17) where is and ir are the vector of the stator winding and the rotor bars currents which are considered as is = ( ia ib ic )T (18) ir = ( ib1 \u00b7 \u00b7 \u00b7 ibm \u00b7 \u00b7 \u00b7 ibn )T (19) 0278-0046 (c) 2019 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. in which ia, ib and ic are the currents of the stator windings and ibm is the mth rotor bar current. The small deviation of the magnetic field energy energy in the directions of all differential variations (Fig. 2(b)) is \u03b4Em = fr\u03b4ro + \u03c4\u03b8r\u03b4\u03b8r + \u03c4\u03d5s \u03b4\u03d5s (20) Thus, the EM radial force of fr and magnetic torques of \u03c4\u03b8r and \u03c4\u03d5s can be obtained by equality of Eqs. (15) and (20) as Qqk = 1 2 iT \u2202L (\u03d5s, \u03b8r,ro) \u2202qk i, qk = ro, \u03b8r&\u03d5s (21) The instantaneous currents of stator windings and rotor bars can be obtained by deriving from the electric model of the IM . According to Fig. 3, the voltage equations of the stator windings may be expressed in vector form as us = Rsis + d dt \u03bbs (22) where us and Rs are the stator input voltage vector and the resistance matrices which are defined as us = ( ua ub uc )T (23) Rs = rn + rs rn rn rn rn + rs rn rn rn rn + rs (24) where ua, ub and uc are source voltages of the stator windings" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001381_kem.490.237-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001381_kem.490.237-Figure3-1.png", + "caption": "Fig. 3. Axial section of a gear", + "texts": [], + "surrounding_texts": [ + "Due to simultaneous cutting of the both sides of the tooth space (pinion and gear), the bearing contact bias occurs after assembling of a transmission made in that way. Such an effect is undesirable since it causes loud work of the transmission and non-uniform transfer of the motion which leads to the faster fatigue wear. In order to remove the bias the helical motion is applied during pinion finishing. It is realized by the axial offset of the fixed headstock of the workpiece connected with the generating gear rotation. The generating system of the pinion is the technological hypoid gear obtained thanks to the hypoid offset of the workpiece\u2019s axis by the E value with respect to the cradle\u2019s axis. The cradle with inclined toolhead (tilt) creates the bevel generating gear. For the pinion cutting the head\u2019s axis inclination is applied in order to compensate the difference between the pressure angle and tool profile so in the initial position the c S system is rotated with respect to the d S system by the angle of j around d X axis. The head\u2019s axis inclination angle (in degrees and minutes) and the value of the hypoid technological offset E is determined on the basis of the gear geometry analysis and calculation cards. Positive values mean down-shift by the pinion with left inclination line direction of a tooth or up-shift by the pinion with right inclination line direction of a tooth. The negative value means upshift by the pinion with left inclination line direction of a tooth or down-shift by the pinion with right inclination line direction of a tooth. Mathematical model of tooth flank surface Mathematical notation of the tooth flank is apparent from the equation of the surface of action of the tool, kinematics, and accepted treatment technological system. The following discussion presents the side of the tooth surface obtained with the use of a technological system with a bevel generating gear - this is the case more generally in comparison with a ring generating gear. While processing the envelope, the equation of the tooth flank, which is the bounding surface of the utility, is determined from the system of equations [1, 2, 6]. This system includes the equation of the family of tool surfaces and the equation of meshing, resulting from the method for determining the kinematic envelope: ( ) ( ) , , , , 0 t t t t t t s s \u03b8 \u03c8 \u03b8 \u03c8 \u22c5 = 1 t1 1 1 r n v (1) where: ( ), ,t t ts \u03b8 \u03c81r - determining the vector function of the family tool surfaces system bounded with treating pinion ( 1S ), 1 n - the unit normal vector defined in 1S , ( ), ,t t ts \u03b8 \u03c8t1 1v - the relative velocity vector defined in 1S . Based on the defined technological model, the family of tool surfaces is determined as follows: ( ) ( ) ( ), , ,t t t t t ts s\u03b8 \u03c8 \u03c8 \u03b8= \u22c51 1t tr M r (2) where: ( ),t ts \u03b8tr - vector equation of the tool surface referred to the system associated with the tool, tS , ,t ts \u03b8 - curvilinear coordinates surface form, ( )t\u03c81tM - the transformation matrix being the product of a transformation matrix representing the rotations and translations of homogeneous coordinate systems included in the technological gear model t\u03c8 - parameter of motion (in this case - angle of the cradle rotation). Vector equation of the tool surface as a function of curvilinear coordinates ,t ts \u03b8 shows the relationship (3), which involves the processing of the active side of the tooth. ( ) ( ) ( ) cos sin , sin sin cos t wk t wk t t t wk t wk t wk r s s r s s \u03b8 \u03b1 \u03b8 \u03b8 \u03b1 \u03b1 + = + \u2212 tr (3) where: wk\u03b1 - the angle of the external blades, wkr - the radius of the cutterhead. Based on the model of technological gear it is designated a family of the tool surfaces according to equation (2), for which it determines conversion matrix equation (4). ( ) ( )( ) ( )1t t t\u03c8 \u03c8 \u03c8 \u03c8= \u22c5 \u22c5 \u22c5 \u22c5 \u22c5 \u22c5 \u22c51t 1w wr rh hm mk kc cd dgM M M M M M M M M (4) where: ( )t\u03c81tM - conversion matrix, ij M - the elementary transformation matrices representing the rotations and translations of homogeneous coordinate systems, ij - subscript indicating the direction of the transformation from system jS to system iS . Technological gear model is also used to determine (in the chosen system, for example 1(S or )mS a unit normal vector and relative velocity. The normal vector to the surface of the tool sets at any of the predefined layouts. In this way we can get the components of meshing equation. Solving equations (1) by eliminating one of the variables such as: ts from the meshing equation and then substituting into the equation of the family of tool surfaces ( )1 , ,t t ts \u03b8 \u03c8r we can obtain the tooth flank surface equation in two-parametric form (5). ( ) ( )( ), , , ,t t t t t t ts\u03b8 \u03c8 \u03b8 \u03c8 \u03b8 \u03c8=1 1r r (5) where: ( ),t t\u03b8 \u03c81r - the equation of the pinion tooth surfaces in the two-parametric form, ( ),t t ts \u03b8 \u03c8 - the variable ts in the function of other parameters. Model of technological gear is designed to create the flank surface of the gear and pinion teeth, which will be used for the analysis of meshing for constructional spiral bevel gear. In order to obtain a tooth surface as the family of tool surfaces it should be designated in the system rigidly bounded with cutting pinion S1 , which represents the envelope to family of surfaces of the cut gear tooth ( )\u03a3 1 . Family of tool surfaces is shown in the following way: ( ( )) ( ( ( ) ( )) )\u03c8 \u03c8 \u03c8\u2190 \u2190 \u2190 \u2190 \u2190 \u2190 \u2190 \u2190 \u2190= \u22c5 \u22c5 \u22c5 \u2212 + \u22c5 \u22c5 \u22c5 \u2212 \u2212(t) (t) 1 1 w 1 t w r r h h m m k t c d d t t k c r hr L L L T L L L r T T (6) In this equation the vector record (t) 1 r , (t) m r , (t) gr , concerns a family tool surfaces ( )\u03a3 t (as evidenced by the superscript ( t )), set respectively in the cutting gear system S1 , the basic system (stationary) and the tool system. Other signs are bounded with the transformations of coordinate systems used in the model of technological gear. The applied rotation and translation matrixes of the coordinate system are as follows: cos i 0 sin i 0 1 0 sin i 0 cos i , \u2190 = \u2212 d g L (7) cos j sin j 0 sin j cos j 0 0 0 1 , \u2190 = \u2212 c d L (8) U cos q U sin q 0 , \u2190 \u2212 \u22c5 = \u2212 \u22c5 k c T (9) cos sin 0 ( ) sin cos 0 0 0 1 , \u03c8 \u03c8 \u03c8 \u03c8 \u03c8\u2190 \u2212 = t t m k t t t L (10) 1 B1 1 0 A X p ( ) ,\u03c8 \u2190 = \u2212 + h m t" + ] + }, + { + "image_filename": "designv11_33_0002472_rpj-06-2017-0120-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002472_rpj-06-2017-0120-Figure4-1.png", + "caption": "Figure 4: Envelope of one dental structure.", + "texts": [ + " Instead, a certain range (conditioned by the workspace yield influence) of achievable average manufacturing speeds can be established. In order to describe the workspace yield influence, two influential factors are used, and a new geometrical entity is defined called \u201cpart\u2019s envelope\u201d. The part\u2019s envelope is a block with edges (a, b, c) parallel to individual coordinate axes. The edges\u2019 lengths are equal to the maximal part\u2019s D ow nl oa de d by I N SE A D A t 0 0: 13 2 6 A pr il 20 18 ( PT ) dimensions in individual axes, according to the part\u2019s orientation in the machine\u2019s workspace as shown in Figure 4. The first factor is defined as the ratio between the part\u2019s volume and the volume of the part\u2019s envelope. It is named the \u201cVolume Ratio\u201d and is used to describe the influences of the part\u2019s geometry on the workspace yield (1) (Brajlih et al., 2010): Volume Ratio= Part\u2019s volume Part\u2019s envelope volume (1) The Volume Ratio is a parameter of the object\u2019s geometrical complexity, which also defines the apparent density of the object. To calculate the average manufacturing speed of the machine, we need to define the utilization of the whole machine\u2019s workspace", + ", Joachim Schulz, J., Graf, P., Ahuja, B., Martina, F. (2016), \u201cDesign for Additive Manufacturing: Trends, opportunities, considerations, and constraints\u201d, CIRP Annals - Manufacturing Technology, Vol. 65 pp. 737\u2013760. D ow nl oa de d by I N SE A D A t 0 0: 13 2 6 A pr il 20 18 ( PT ) Figure 3: Comparison between a one dental structure in the working space of the machine and a fully occupied working space with twenty-one dental structures. D ow nl oa de d by I N SE A D A t 0 0: 13 2 6 A pr il 20 18 ( PT ) Figure 4: Envelope of the one dental structure. D ow nl oa de d by I N SE A D A t 0 0: 13 2 6 A pr il 20 18 ( PT ) Figure 6: The average manufacturing speed of the MLab Cusing\u2122 machine D ow nl oa de d by I N SE A D A t 0 0: 13 2 6 A pr il 20 18 ( PT ) element Time (h) x (mm) y (mm) z (mm) STL volume (mm 3 ) Envelope Volume (mm 3 ) 1 pcs 0.97 26.28 14.32 8.5 400 3199 21 pcs 5.38 82.85 84.43 8.5 8330 59458 D ow nl oa de d by I N SE A D A t 0 0: 13 2 6 A pr il 20 18 ( PT ) Volume Ratio Tray Ratio AMS (cm 3 /h) 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003741_oceanse.2019.8867249-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003741_oceanse.2019.8867249-Figure2-1.png", + "caption": "Fig. 2. Designed of the battery holder and sliding mass", + "texts": [], + "surrounding_texts": [ + "The overall system was designed through an iterative testing and validating process considering the requirements established according to the environmental characteristics of the Peruvian coastline and the UWG state-of-art. Those requirements are presented as follows. \u2022 The main structure must house the electronics in a watertight compartment. \u2022 The shape and configuration of the hull must ensure the buoyancy and stability of the whole system. \u2022 The UWG must resist depths of 150m maximum accord- ing to the Peruvian coastal shelf. \u2022 the UWG must have a minimum time of autonomy of seven (07) days." + ] + }, + { + "image_filename": "designv11_33_0000988_s00542-010-1189-3-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000988_s00542-010-1189-3-Figure1-1.png", + "caption": "Fig. 1 Mechanical structure of the disk-spindle system in a computer hard disk drive Fig. 2 Coordinate system of the coupled journal and thrust bearings", + "texts": [ + " It shows that the groove design optimized using the proposed method has a small whirl radius in the steady state. It also shows that it has very little displacement due to the shock excitation, and that it quickly recovers to the equilibrium state. Robust design of the disk-spindle system has been one of the most important issues for computer hard disk drives (HDDs), because the vibration of the disk-spindle system prohibits the magnetic head from reading or writing the data on a rotating magnetic disk. Figure. 1 shows the mechanical structure of the disk-spindle system in a HDD. The vibration of the diskspindle system is strongly affected by the fluid dynamic bearings (FDBs), which are composed of the coupled journal and thrust bearings that generate the radial and axial pressure to support the rotating disk-spindle. They provide the damping coefficients to suppress the vibration of the diskspindle system, but they also excite the disk-spindle system with a unique form of vibration called \u2018\u2018half-speed whirl\u2019\u2019 due to the instability of the FDBs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000928_12.978271-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000928_12.978271-Figure1-1.png", + "caption": "Figure 1: Schematic diagram of the sensor device. Figure 2: Light spot through the bearing", + "texts": [ + " In this study, an optical fiber sensor approach is presented to measure the rotation speed of the cage in oil lubricated bearing. Several factors that have potential effects on the measurement of the senor were determined, and the approach is verified by experiments. To measure the cage rotation speed and rolling elements, the transmitter of the optical fiber sensor is placed on one side of the bearing, and the receiver of the sensor is mounted on the other side. The light beam directly points to the gap between the cage and the ring. The schematic diagram of the optical sensor measurement device is shown as figure 1. When the rolling elements running cross the light beam, the light spot and shade of the rolling elements will be formed alternately on the sensor receiver, and these impulse light signals will be captured. Through counting the frequency of the impulse signals, the rolling element passing frequency and cage speed can be obtained. For conventional rolling element bearing with cage, if the number of rolling elements is n, then the frequency of rolling element running pass through the optical sensor light beam is n times of the cage rotation speed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002015_tmech.2017.2713397-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002015_tmech.2017.2713397-Figure1-1.png", + "caption": "Fig. 1 Schematic of a PM motor", + "texts": [ + " The method is validated on a Direct-Drive Dual-Disc (D4) spindle motor for duplex lathe-cutting of hard-to-machine materials, where harsh bearing reactions due to cutting forces and rotor/workpiece inertia cause vibrations and downgrade of the cutting quality [5][8]. The D4 motor that has 248 independently controllable current inputs provides an essential testbed for experimentally investigating the effects of different motor designs, control methods and implementation algorithms on external-load compensation. II. MOTOR WITH SIX-DOF FORCE/TORQUE COMPENSATION Fig. 1(a) schematically shows a typical permanent magnet (PM) motor consisting of a stator with NE electromagnets (EMs), a mechanical bearing and a rotor with NP (NE) PMs to allow spinning (at speed \u03c9) around its Z-axis. Equation (1) describes the dynamics of the rotor (inertia J) and the bearing (radius rb and viscous damping b), where the subscripts l, b and e of the force F and moment T (or torque about Z-axis) denote the contributions from the load, bearing reaction and electromagnetic actuation respectively: + e b l ZJ b e T (1a) where e e Ze T , 0 sign( )b b Z be T (1b,c) 2 20, ,l b e ne n X Y T T T (2) and 0l b e F F F ", + " > REPLACE THIS LINE WITH YOUR PAPER IDENTIFICATION NUMBER (DOUBLE-CLICK HERE TO EDIT) < 3 parametric effects on load compensation, and 3) real-time solutions for implementing the motor inverse model. A. Forward Force/Torque Model in PM motors The motor forward model, which provides a basis for controlled system design and analysis, determines the six-DOF force and torque components as a function of current inputs for a given design geometrical parameters (NE, NP, hE, hP, rE, rP) where NE/2 and NP/2 may be an odd or even integer. Schematically, the stator EMs and rotor PMs are evenly spaced around the circumferences of two parallel planes with radii rE and rP, respectively (Fig. 1a), the locations of which are described by their respective position vectors, sj and ri with respect to a reference plane as shown in Fig. 1(b). For a magnetically linear motor design, the actuation (Fe, Te) can be computed from the sum of individual interactions (fij, \u03c4ij) between the jth (j=1, 2\u2026NE) EM and ith PM (i=1, 2\u2026NP). As derived in Appendix (A.5), the Lorentz force fij between the ith PM and current-carrying jth EM depends only on the separation angle 1cos /i j jij i r s sr between them. The Lorentz force fij acting on the ith PM can be expressed in terms of the kernel functions , , R T Z as (6a): ( ) ( ) ( )ij i j R ij R T ij T Z ij Zu e e e f (6a) ij j ij \u03c4 s f (6b) The locations of NE EMs (each with a local frame , ,Rj Tj Zje e e ) are given in stator coordinates, by (7a~d): 0 ; ; ; 0 10 0 j j j j j j E j E Rj Tj Zj E r C C S r S e S e C e h s (7a~d) where 2 ( 1) /j Ej N ; C and S represent cosine and sine; and the subscripts (R, T, Z) of the unit vector e indicate the components along the radial, tangential and Z directions of the local EM coordinate-frame respectively", + "5) is coordinate-independent, and proportional to and u with a coefficient depending only on the separation angle \u03c3 between the position vectors (s and r). For an EM and a PM with shapes symmetrical about \u03c3=0, T and Z can be characterized by an even and odd function respectively: ( ) ( ) ( ) ( ) ( ) ( ) 0 0 T R Z o T R Z o (A.6) The o value depends on the specific PM and EM designs. M S+ PM S n EM (with winding volume Ve) +n R h h R+ sr eR eT Fig. A1 EM and PM B. Actuating Force/Torque Orthogonality Consider the system matrix A with elements given in (9b, c) for the forward model (16a) characterizing the motor (Fig. 1) where Np and NE are multiple of four. Using (13) and (20) with the aid of Fig. 3(c~d), / 2 T ( / 2) ( / 2) 1 / 4 ( / 4) ( / 4) 1 2 0 0 E E E E E E N X n X n jX jn j N X j N n j N jX jn j N X j N n j k k k k k k k k n Y n Z K K K K (B.1) Thus T T T 0X Z Y Z X Y K K K K K K (B.2a,b,c) Similarly, T T T 0X Z Y Z X Y P P P P P P (B.3a,b,c) 2 1T 2 1 0 0 E j j E j j N jT jR jT j n Z E N jT jR jT j C k k k S n X r S k k k C n Y K P (B.4a) (B.4b) T 2 1T T 2 1 0 0 E j E j N E X Z E jZ j Z n N E Y Z E jZ j h r k S n X h r k C n Y K K K P K K (B" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000935_1.4737888-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000935_1.4737888-Figure1-1.png", + "caption": "FIG. 1. The structure of the shock tester: 1\u2014bolt; 2\u2014stopping block; 3\u2014threaded hole; 4\u2014impact table; 5\u2014thimble; 6\u2014pulse shaper; 7\u2014spring; 8\u2014upperstand; 9\u2014drop rod; 10\u2014cover; 11\u2014shell; 12\u2014spring; 13\u2014slide groove; 14\u2014location steps; 15\u2014handler; 16\u2014cover; 17\u2014pin; 18\u2014cross-ring; 19\u2014stopping pin; 20\u2014washer; 21\u2014bolt; 22\u2014sliding ring; 23\u2014spring; and 24\u2014anvil.", + "texts": [ + " In this paper, the working principles and structure design of a shock tester are described and key parameters are analyzed in detail. We present results of tests with a prototype device in which a material with a high specific modulus was used. A velocity gain can be achieved through collisions between vertically-stacked masses. This phenomenon can be used to develop very-high acceleration and low-cost shock testing machines. Our study proposes a practical design based on concepts developed by Rodgers et al.1516 The design of the shock tester is shown in Fig. 1. The expected acceleration range of the shock tester is in excess of 100 000 g and the expected widths of the acceleration pulses are not less than 50 \u03bcs. The maximum expected velocity of the test object is in excess of 25 m/s. During operation, the full apparatus is vertically connected to a concrete foundation by bolts through the anvil (24). The article under test is attached to the impact table (4) by a bolt through a threaded hole (3). The impact table is suspended by a spring (7) and thimble (5), and slides along the hole in the upper-stand (8)", + " Although the experiments and results provided a rudimentary validation of our design, because of the prototype apparatus and the randomness of the manual excitation, these experiments cannot determine all the device properties including the measuring range of acceleration, accuracy, reproducibility (excited by spring with a constant compression deformation), and acceleration pulse durations. Therefore, our future focus will be to complete the construction of the remaining components and to build the complete shock tester as shown in Fig. 1. We will then carry out a more in-depth investigation of the properties of the shock tester. In particular, the theoretical analysis and experimental study on acceleration pulse shaping will be presented in our future papers. This paper proposed a simply constructed but efficacious shock tester for the simulation of high-g level shock environments. The operating principles and structural design of the shock tester are described. In particular, some key parameters are analyzed in detail based on a simplified mechanical model and the constraints rising from the size and shape of the testing articles, and the manual operating requirement" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001477_s0036024412070345-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001477_s0036024412070345-Figure4-1.png", + "caption": "Fig. 4. CVs of 5.0 \u00d7 10\u20133 mol L\u20131 K4[Fe(CN)6] at NG/NC/GCE (1), NC/GCE (2), and bare GCE (3). Supporting electrolyte: 0.1 mol L\u20131 KCl; scan rate: 50.0 mV s\u20131.", + "texts": [ + " The major diffraction peaks can be indexed as the gold face centered cubic (fee) phase based on the data of the JCPDS file [20]. The NG/NC showed a typical peak of (002) phase of NC with a character istic peak at 2\u03b8 = 26.4\u00b0 [21], and the peaks appeared at 38.3\u00b0, 44.7\u00b0, 64.7\u00b0, and 78.6\u00b0 can be assigned to (111), (200), (220), and (311) crystalline plane dif fraction peaks of gold, respectively. While diffraction peaks of CNP and graphite are found at 24.7\u00b0 and 26.6\u00b0, respectively. The cyclic voltammetry (CV) of modified GCE in the K3Fe(CN)6\u2013K4Fe(CN)6 system is shown in Fig. 4. Compared with the bare GCE, NC/GCE, and NG/NC/GCE shows higher peak current. The increase in voltammetric response of ferrocyanide is simply due to the intrinsic properties of the NG/NC, which works to produce a larger peak current than that of the bare GCE. The real active surface area will be 1460 RUSSIAN JOURNAL OF PHYSICAL CHEMISTRY A Vol. 86 No. 9 2012 YANG SONG et al. estimated. In a reversible process, the following Ran dles\u2013Sevcik formula at 298 K [22] has been used: ip = 2.69 \u00d7 105n3/2Ac0 v 1/2, where, ip refers to the anodic peak current, n is the electron transfer number, A is the microscopic surface area of the electrode, D0 is the diffusion coefficient of K3Fe(CN)6, C0 is the bulk concentration of K3Fe(CN)6 and v is the scan rate" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003725_embc.2019.8857155-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003725_embc.2019.8857155-Figure1-1.png", + "caption": "Figure 1. Sagittal section of 3D FE model illustrating geometry, loading configuration and boundary conditions, plus regions of interest for actuation.", + "texts": [ + " The model was generated from Magnetic Resonance Imaging (MRI) scans of a person with unilateral transtibial amputation (Institutional Ethics Approval references: ERGO 29927 and 41476). The scans were segmented (Scan IP 2017.06, Synopsys Inc., USA) into bones, soft tissue, tendon and meniscus structures and the 3D meshed model imported into ABAQUS 6.14 FEA software for analysis. The model was parameterized to enable variation of the loading, and liner stiffness at the three regions of interest of the residuum, as shown in Fig. 1. This provided training data and biomechanical rationale for tissue injury risk prediction as well as a basis for structuring the different actuators within the overall controller model, as regards coupled vs. uncoupled subsystems. The actuation components implemented at the residuumsocket interface were initially modelled as translational spring and damper systems. The damping and spring stiffness coefficient values were considered constant parameters of the actuation component subsystem and the manipulated component was x, a measure of the expansion/contraction of the spring and damper elements", + " These correspond to linear-elastic and nearly incompressible models with Poisson\u2019s ratio, \u03bd = 0.49, and Young\u2019s Modulus, E = 200 kPa, 380 kPa and 500 kPa, respectively [12]. Quasi-static 2D sagittal plane loading was applied to the model, with an axial force (FZ), an anterior-posterior force (FX), and an extensionflexion moment (MY) for approximation of typical prosthesis use during the stance phase of gait [18], [19]. The loads were applied to the base of the socket whilst the model was constrained at proximal femur and patella tendon (Fig. 1). Further, the liner stiffness was kept constant and several loading cases were applied to the model to assess the changes in the interface pressure and the compressive strain at each of the three limb locations, throughout different stages of the stance phase. The interface pressure-tissue strain relationships with variation of load at each of the residuum locations were also investigated. Analysis of the FE model factorial study based on the 95th percentile estimations for compressive strain showed little interaction between regions, with a variation of <2% strain with change in liner stiffness" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003937_j.measurement.2019.107431-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003937_j.measurement.2019.107431-Figure4-1.png", + "caption": "Fig. 4. Schematic diagram of arc pressure testing experiment system.", + "texts": [ + " g \u00bc Q Qnominal \u00f04\u00de Table 1 Processing parameters used in experiments. Arc current I/A Arc length Larc/mm Arc moving speed vf/mm s 1 Gas flow rate qs/L min 1 Circulating water flow rate W/g s 1 140 3 5 20 60 4 5 6 7 g \u00bc WC R1 0 Tout Tin\u00f0 \u00dedt UaIt \u00f05\u00de The calculation method indicates that the larger the shadow area is, the higher the total heat transferred from arc to calorimeter is. An arc pressure testing experiment system is developed to test arc pressure in real time. The structure of arc pressure testing experiment system is shown in Fig. 4. The anode plate is a copper block with a prefabricated test hole. The arc torch must pass uniformly over the test hole during the test. Accordingly, the arc pressure is transferred through the test hole and the test tube to the pressure sensor. The pressure sensor converts the arc pressure signals into voltage signal and transfers it to computer system. In order to avoid burning loss of copper block, circulating cooling water is used for continuous cooling of copper block. The arc continues to heat the calorimeter for 35 s in temperature test" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002365_j.matpr.2018.08.118-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002365_j.matpr.2018.08.118-Figure6-1.png", + "caption": "Fig. 6. Comparison of the virtual and experimental estimation of the lifetime on the very simplified model of an engine mount.", + "texts": [ + " Due to the large number of data points, a statistical analysis of the scattering was possible. The fit of the green line in Fig. 5b is a statistical evaluation with the reliability and lifetime function of the software JMP 11.0, (Statistical Discovery SAS, NC, US). The fit follows a Weibull distribution. With this data, one can start lifetime estimation by implementing these datasets into durability software. The lifetime modeling, was done in the software nCode from HBM on a simplified model of an engine mount (see Fig. 6). For the estimation of the lifetime in nCode several steps are essential. A FE analysis on the engine mount, using the OgdenN3 model on three different elongations, was performed and implemented in the FEInput glyph in nCode. The next step was the corresponding load mapping, so that the simulations were equivalent to the experimental validation test setup. The Ru-Ratio was constant at 0.1. The implementation of the LSWC is realized with the MaterialMananger provided by nCode in form of E-N curves, based on the Coffin-Manson-Basquin strain life relationship. The software calculates the damage by looking up the strain amplitude of the relevant cycle on that curve. Holzweber et al. / Materials Today: Proceedings 5 (2018) 26572\u201326577 26577 The results and the deviation between the prediction and the experimental are shown in Fig. 6. The slopes of both curves in the double logarithmic plot are quite alike and the curves are following a similar trend. Nevertheless, there is a deviation between the simulated and the experimentally determined results. The deviations seem to get larger in the high cycle region. It should be mentioned that the uniaxial limit as well as the endurance limit was not characterized, neither with the specimens or the component. The big variety of the mechanical material testing necessary to derive a valid material model gives a proper understanding of the complex material behavior of thermoplastic polyurethanes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001202_icra.2012.6225091-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001202_icra.2012.6225091-Figure5-1.png", + "caption": "Fig. 5. Angles between contact forces and \u2212fext.", + "texts": [ + " A pair of contacts at points p and q on the boundary of a polygon P , with contact normals n\u0302p and n\u0302q can balance an external wrench wext = (fext, \u03c4ext) (where fext 6= 0 and n\u0302p, n\u0302q and fext are not collinear), if and only if both the following conditions are met: \u2022 The vectors n\u0302p, n\u0302q and fext positively span the plane, and \u2022 the wrenches wext, (n\u0302p,p \u00d7 n\u0302p) and (n\u0302q, q \u00d7 n\u0302q) lie in a plane through the origin of wrench space. Recall that a grasp has quality Q if the maximum of the magnitudes of the contact forces needed to resist wext, equals 1/Q. Kruger and van der Stappen [13] showed that to balance wext with grasp quality greater or equal to Q, a grasp must also satisfy the following: 0 < \u03b8p \u2264 tan\u22121 \u221a 4 Q2\u2016fext\u20162 \u2212 1 (1) \u03b8p \u2264 \u03b8q \u2264 tan\u22121 tan \u03b8p Q\u2016fext\u2016 sec \u03b8p \u2212 1 , (2) where \u03b8p \u2208 (0, \u03c0) is the angle between n\u0302p and \u2212fext and \u03b8q \u2208 (0, \u03c0) is the angle between n\u0302q and \u2212fext (see Fig. 5). If n\u0302p is collinear with fext, then n\u0302q must also be collinear with fext and at least one of the contact normals must have the opposite direction to fext. In this case the quality constraint results in constraints on the contact positions, rather than the normal directions. a) Review of edge pair algorithm: Based on the above criteria, Kruger and van der Stappen developed an algorithm to compute all pairs of edges of P that admit frictionless two-fingered grasps that can balance a pure force fext with grasp quality at least Q" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000209_tac.2009.2020643-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000209_tac.2009.2020643-Figure2-1.png", + "caption": "Fig. 2. .", + "texts": [ + " This implies the criterion function changes its sign on the control curve (see [8]). Because the control curve is arbitrary, the system (7) is globally controllable by Proposition 3.2. Case 2: The eigenvalues of are two different real numbers with same sign. Without loss of generality, we assume that , . 1The two ends of a curve extending to infinity means that: , when and . Similar to Case 1, we can assume that there are two number such that , . By the same method to Case 1, we can draw the trajectories of Equation and a control curve in the phase plane as shown in Fig. 2. Obviously, we have the criterion function changes its sign on any control curve . Hence, the system (7) is globally controllable by Proposition 3.2. Case 3: One eigenvalue of is zero, the other is a nonzero real number. Without loss of generality, we assume that , . Similar to Case 2, we assume that there are two number such that , . By the same method, it is easy to know that the the criterion function changes its sign on any control curve as shown in Fig. 3. Hence, the system (7) is globally controllable by Proposition 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000772_1.4005510-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000772_1.4005510-Figure4-1.png", + "caption": "Fig. 4 Inverted pendulum with dc servo motor for sensing and actuation", + "texts": [ + " Figures 3(a) and 3(b) show that the PZT element considerably accelerates the vibration damping successfully stabilizing the first two modes within a span of 1 s. The lower rate of convergence implies lower required control effort for stabilization. This rapid vibration attenuation is imperative for safe operation of large, flexible structures. 6.2 Inverted Pendulum Stabilization. The second example of the application of the persistence based observer and controller considers an open-loop unstable system. Here, we look at stabilizing the inverted pendulum (Fig. 4) in the unstable, h\u00bc 0 configuration using a dc servo motor. This example is referred to the work in Ref. [20]. The servo motor can operate in both sensor and actuator modes, which allows for switching between the two modes, thus reducing control hardware required. In contrast to the work in Ref. [20], we consider smooth switching, which results in smooth state and control evolutions. The states of the system are x \u00bc \u00bdh; _h; i T , where i is the motor current, the input u is the control voltage, and the output y is the back-EMF from the servo" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000400_978-90-481-8764-5_3-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000400_978-90-481-8764-5_3-Figure1-1.png", + "caption": "Fig. 1 The second prototype of the muFly helicopter", + "texts": [ + " Using the linearized model is adequate since muFly will only operate close to the hover point. High velocities are not foreseen for the missions. The paper is organized as follows. In Section 2 the muFly micro helicopter and its hardware setup is shown, followed by the nonlinear model in Section 3. The identification process with the PEM is presented in Section 4 and the identification and verification results are shown in Section 5. muFly is a 17 cm in span, 15 cm in height coaxial helicopter with a mass of 95 g (Fig. 1). The two main rotors are driven by two lightweight brushless DC (BLDC) 28 Reprinted from the journal motors and are counter rotating to compensate the resulting torque due to aerodynamical drag. This allows to control the yaw by differential speed variation of the two rotors, where as the altitude can be controlled varying the rotor speed simultaneously. The motor speed is reduced by a gear to achieve a higher torque on the rotor side. A benefit of the coaxial setup is that one rotor can be used to help stabilizing the helicopter using a stabilizer bar" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001465_powereng.2011.6036541-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001465_powereng.2011.6036541-Figure1-1.png", + "caption": "Fig. 1. Example of split-phase alternators with N stator windings supplying N cascaded diode bridges: (a) N=2; (b) N=4.", + "texts": [ + " SYSTEM LAYOUTS FOR MULTIPHASE GENERATORS The most common solution to obtain a DC current (which can be possibly processed through downstream DC/AC converters) out of an AC alternator is to connect one or more rectifiers (either controlled or not) to the output terminals of the alternator, [2]-[4]. In particular, if the alternator has split-phase winding configurations (i.e. its stator phases are split into N three-phase sets, [2]-[4]), it is natural to connect each threephase set to a full bridge according to the configuration shown in Fig. 1. For example, in Fig. 1 the bridges are connected in series, but they may be connected in parallel as well [4]. The benefits associated to using multiple stator windings and rectifiers are highlighted in [3] and are mainly related to the enhancement of the output DC current, the torque ripple and the air-gap field harmonic content. Furthermore, the presence of multiple rectifiers provides the system with fault tolerance features. A criticality of the system, however, is constituted by the strong distortion caused by the rectification process in stator phase currents [5]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002839_j.mechmachtheory.2018.06.021-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002839_j.mechmachtheory.2018.06.021-Figure5-1.png", + "caption": "Fig. 5. Mathematical sketches for finding the envelope of a family of straight lines: (a) the rolling cone and the other essential factors; (b) the envelope.", + "texts": [ + " Spin motion is caused by the different speed distributions of the two contact areas and can be described as an angular velocity along the normal \u2217 Corresponding author at: School of Mechanical Engineering, Xihua University, No. 999 Jinzhou Road, Chengdu, Sichuan Province, 610039, PR China. E-mail address: tsingtau.lee@gmail.com (Q. Li). https://doi.org/10.1016/j.mechmachtheory.2018.06.021 0094-114X/\u00a9 2018 Elsevier Ltd. All rights reserved. Nomenclatures \u03b8 half cone-angle of roller \u03b3 titling angle of roller D distance ( Fig. 3 ) r radius of circle ( Fig. 3 ) C cone apex P arbitrary point ( Fig. 3 and Fig. 5 ) v movement speed of the cone apex \u03c9 angular speed of axis t time parameter \u03bb individual parameter that represents the length of CP A and B contact points on input disc and output disc A\u2019 and B\u2019 foot points ( Fig. 6 ) A\u201d and B\u201d foot points ( Fig. 6 ) \u03c9 i and \u03c9 o rotational speeds of input disc and output disc \u03c9 r rotational speed of roller i speed ratio R length ( Fig. 7 ) \u03c9 spin spin speed \u03c9 dg rotational speeds of driving component \u03c9 dn rotational speeds of driven component \u03b8dg angle between the contact area and the rotational axis of the driving component \u03b8dn angle between the contact area and the rotational axis of the driven component \u03c3 spin spin ratio \u03c9 spinT and \u03c3 spinT spin speed and spin ratio of half-toroidal CVT \u03c9 spinL and \u03c3 spinL spin speed and spin ratio of logarithmic CVT \u03c9 spinI and \u03c3 spinI spin speed and spin ratio of inner cone CVT \u03c9 spinO and \u03c3 spinO spin speed and spin ratio of outer cone CVT e the distance of the toroidal cavity from the disc rotation axis k aspect ratio Cr in creep coefficient at input contact point Cr out creep coefficient at output contact point \u03bcin tangential force coefficient at input contact point \u03bcout tangential force coefficient at output contact point \u03c7 in spin momentum coefficient at input contact point \u03c7out spin momentum coefficient at output contact point t in input traction coefficient t out output traction coefficient \u03b7 total efficiency direction at the point of contact [17,18] ", + " In this section, we present the calculation of an envelope of a family of straight lines and propose a new type of zero-spin CVT referred as inner cone CVT. On the basis of the discussion in Section 2.1 , the proposed method cannot automatically guarantee the zero-spin condition, because the aim of envelope theorem is to determine the curve that is tangent to each member of the curve family. In other words, the directions of common tangents cannot be controlled by the method. Thus, a rolling cone (tapered roller) is introduced, as shown in Fig. 5 (a). The generatrices of the cone are two straight lines that can be extended and subsequently intersected at a single point referred as cone apex. The roller rotates on its axis to transmit power. The rotation axis and the two generatrices intersect at the cone apex. For a component that is in contact with the rolling cone, the common tangent line (contact area) should be the generatrix of roller. In other words, the cone apex is the intersection point of the contact area and the roller rotation axis. By controlling the location of the cone apex on the disc rotation axis, the spin motion can be eliminated (see Section 2.1 ). Fig. 5 (a) shows the plane coordinate system XOY , in which the X -axis coincides with the disc rotation axis. The generatrix of the roller is supposed to be a rolling cone with a half-cone angle \u03b8 . The cone apex of the rolling cone is confined to X - axis and moves at the speed of v . Angle \u03b3 is the angle between the roller rotation axis and the Y- axis. Points O and C are the original point and the cone apex, respectively. Point P is an arbitrarily located on the roller\u2019s generatrix and represents the contact point between the generatrix and the envelope curve", + " The vectors OC and CP can be expressed as OC = [ v t 0 ] ; CP = [ \u03bb sin ( \u03b8 \u2212 \u03c9t ) \u03bb cos ( \u03b8 \u2212 \u03c9t ) ] (8) where \u03bb is an individual parameter that represents the length of CP and is equal to | CP |. OP = OC + CP = [ x y ] = [ v t 0 ] + [ \u03bb sin ( \u03b8 \u2212 \u03c9t ) \u03bb cos ( \u03b8 \u2212 \u03c9t ) ] (9) Then, the equations of the family of the straight lines can be derived. { x = v t + \u03bb sin (\u03b8 \u2212 \u03c9t) y = \u03bb cos (\u03b8 \u2212 \u03c9t) t \u2208 ( \u03b8\u2212\u03c0/ 2 \u03c9 , \u03b8 \u03c9 ] , (10) where \u03b8 , v and \u03c9 are constants, while t is an independent variable. The family of straight lines with a positive slope in Fig. 5 (b) is calculated with Eq. (10) . However, the family of straight lines with a negative slope is not considered, and this aspect will be discussed at the end of this subsection. The next step is determining the envelope. \u2202(x,y ) \u2202(\u03bb,t) = \u2223\u2223\u2223\u2223 sin (\u03b8 \u2212 \u03c9t) cos (\u03b8 \u2212 \u03c9t) v \u2212 \u03bb\u03c9 cos (\u03b8 \u2212 \u03c9t) \u03bb\u03c9 sin (\u03b8 \u2212 \u03c9t) \u2223\u2223\u2223\u2223 = \u03bb\u03c9 sin 2 (\u03b8 \u2212 \u03c9t) \u2212 v cos (\u03b8 \u2212 \u03c9t) + \u03bb\u03c9 cos 2 (\u03b8 \u2212 \u03c9t) = \u03bb\u03c9 \u2212 v cos (\u03b8 \u2212 \u03c9t) = 0 (11) Eq. (11) is simplified as \u03bb= v \u03c9 cos (\u03b8 \u2212 \u03c9t) . (12) By substituting Eq. (12) into Eq. (10) , the envelope of the straight lines with a positive slope can be represented as: { x = v t + v \u03c9 sin (\u03b8 \u2212 \u03c9t) cos (\u03b8 \u2212 \u03c9t) y = v \u03c9 cos 2 (\u03b8 \u2212 \u03c9t) (13) The envelope is also illustrated in Fig. 5 (b). The family of straight lines with a negative slope can be represented as: { x = v t \u2212\u03bb sin (\u03b8 + \u03c9t) y = \u03bb cos (\u03b8 + \u03c9t) (14) Similarly, the envelope equation of the family of straight lines with a negative slope is { x = v t \u2212 v \u03c9 sin (\u03b8+ \u03c9t) cos (\u03b8+ \u03c9t) y = v \u03c9 cos 2 (\u03b8+ \u03c9t) . (15) The envelope represented by Eq. (15) is paired with the envelope represented by Eq. (13) , as shown in Fig. 5 (b). The two envelopes are axisymmetric. Thus, the envelope of the straight lines with a positive slope is chosen as the research object in this study. Nonetheless, the findings of the present work are also applicable to the other envelope. Convexity\u2013concavity identification is necessary process because the introduced rolling cone ( Fig. 5 (a)) should come in contact with a disc with a convex-curved generatrix in quadrants I and II or a concave-curved generatrix in quadrants III and IV. In other words, the second-order derivative of y with respect to x in Eq. (13) needs to be derived. Indeed, d 2 y d x 2 = d dx ( dy dx ) = d dt ( dy dx ) \u00d7 dt dx , (16) where dy dx = ( dy dt )( dx dt ) . (17) From Eq. (13) , we can obtain dx dt = v \u2212 v \u03c9 [ \u2212\u03c9 sin 2 ( \u03c9t + \u03b8 ) + \u03c9 cos 2 ( \u03c9t + \u03b8 ) ] = v + v [ sin 2 ( \u03c9t + \u03b8 ) \u2212 cos 2 ( \u03c9t + \u03b8 ) ] , = 2 v sin 2 ( \u03c9t + \u03b8 ) (18) dy dt = \u22122 v cos ( \u03c9t + \u03b8 ) sin ( \u03c9t + \u03b8 ) ", + " (16) , it follows: d 2 y d x 2 = \u03c9 sin 2 ( \u03c9t + \u03b8 ) 1 2 v sin 2 ( \u03c9t + \u03b8 ) = \u03c9 2 v 1 sin 4 ( \u03c9t + \u03b8 ) . (22) It can be inferred that when v / \u03c9 > 0 the envelope is in quadrants I and II (y > 0, Eq. (13) ) and its second-order derivative is greater than zero ( Eq. (22) ). In contrast, when v / \u03c9 < 0 then y < 0, the envelope is in quadrants III and IV and its second derivative is smaller than zero. Thus, the concavity or convexity of the derived envelope is unsuitable for the contact of the introduced rolling cone, as shown in Fig. 5 (a). To overcome this drawback, a particular structure is needed. As shown in Fig. 5 (b), the straight lines, which are the generatrices of roller, are always under the envelopes. Furthermore, the lines with positive and negative slopes should be used in pairs. Therefore, the contact surface of the roller should take the form of an inner cone. As shown in Fig. 6 , two inner cone rollers are introduced with rotation shafts bolted together at the intersection point of the generatrices of the inner cone surfaces. On the left of the roller, an input disc is placed with a contact surface generated by the derived envelope", + " In this section, the parameters v and \u03c9 will be discussed separately and replaced with particular expressions. The influ- ence of the v / \u03c9 ratio to the convexity\u2013concavity of the envelope can then be avoided. In Fig. 7 , an outer cone roller is placed into the plane coordinate system XOY . Here, B is the assumed contact point on the roller, while O\u2019B (perpendicular to the generatrix of the roller) intersects X -axis at point O\u2019 . Furthermore, R represents the length of the segment O\u2019B . We suppose that the movement of the roller shown in Fig. 7 is the same as that shown in Fig. 5 (a). In Eq. (13) , the values of v and \u03c9 are assumed to be v = R t cos (\u03b8 \u2212 \u03c9t) (31) and \u03c9 = 1 t . (32) By substituting Eqs. (31) and (32) into Eq. (13) , we can obtain { x = R cos (\u03b8\u2212\u03c9t) + R cos (\u03b8\u2212\u03c9t) sin (\u03b8 \u2212 \u03c9t) cos (\u03b8 \u2212 \u03c9t) y = R cos (\u03b8\u2212\u03c9t) cos 2 (\u03b8 \u2212 \u03c9t) . (33) Thus, { x = R sin (\u03b8 \u2212 \u03c9t) + R cos (\u03b8\u2212\u03c9t) y = R cos (\u03b8 \u2212 \u03c9t) . (34) Eq. (34) represents a family of circular curves, and it can be illustrated more clearly by calculating ( x \u2212 R cos (\u03b8 \u2212 \u03c9t) )2 + y 2 = R 2 . (35) We separately generate the family of circular curves represented by Eq", + " On this basis, the durability performance of the logarithmic CVT has not been improved relative to the half-toroidal CVT. In Fig. 6 , the contact positions relative to the inner cone roller and disc both vary with the speed ratio. Thus, all regions of the working surfaces of the inner cone roller and the disc may experience contact. A similar conclusion can be derived with mathematical equations. For instance, in Eq. (10) , \u03bb is the independent variable that represents the distance from the cone apex to the contact point, similar to the one depicted in Fig. 5 . Additionally, in Eq. (12) , \u03bb varies with parameter t , which is the independent variable of speed ratio in Eq. (30) . Hence, the contact point of the inner cone roller varies with speed ratio. This conclusion holds true for the contact points of the input and output discs. Similarly, the contact ranges of the outer cone CVT components are same as that of inner cone CVT components. The manufacturability of the half-toroidal, logarithmic and outer cone CVTs are similar to one another. They consist of components with finishing exterior surfaces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001664_tmag.2011.2132756-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001664_tmag.2011.2132756-Figure1-1.png", + "caption": "Fig. 1. Dual-stage actuator mechanism.", + "texts": [ + " In Section II, plant nominal models for controller design are built and real models for analysis are also designed by using measurement data. Section III contains a new ideas for better time and frequency responses, which is a main research result in this paper. In a Section IV, test results verify effectiveness of the proposed design method and conclusion follows. DSA is a combined actuator. In the DSA systems, MA is connected to VCM through a hinge and two piezoelectric actuators as shown in Fig. 1. The moving directions of the two piezoelectric actuators are opposite to each other. Thus, this mechanism doubles actuator\u2019s force. In some research [7], one piezoelectric material was utilized for a sensor and suspension vibrations were successfully removed. In the case, the gain of the actuator was exactly reduced by 6 dB (equivalent to 1/2). Even though the two piezoelectric materials are used for the actuation, the force is still weaker than that of VCM. Moreover, the stroke of the MA is relatively shorter than that of VCM" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001243_s11044-012-9325-8-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001243_s11044-012-9325-8-Figure6-1.png", + "caption": "Fig. 6 Link with its reference frames", + "texts": [ + "1 Basic equations The Articulated-Body Algorithm is currently the fastest serial processing algorithm for a series of interconnected rigid-bodies when compared with the other leading algorithms: the Divide-and-Conquer Algorithm [10, 11], and the Hybrid Direct-Iterative Algorithm (HDIA) [2], Table 3. The ABA, developed by R. Featherstone [9] solves the dynamics equation system, which is a system of equation with n equations at n variables, with an O(n) number of operations. The main advantage of ABA is the link-by-link analysis instead of the regular wholesystem analysis. Defining the two reference frames shown in Fig. 6, it is possible to determine the velocities and accelerations of the (i + 1)-frame using Eq. (1) to Eq. (5): vO i+1 = vO i + \u03c9i \u00d7 rL i (1) \u03c9i+1 = \u03c9i + \u03b8\u0307 i (2) aO i+1 = aO i + \u03b1i \u00d7 rL i + \u03c9i \u00d7 ( \u03c9i \u00d7 rL i ) (3) \u03b1i+1 = \u03b1i + \u03b8\u0308 i + \u03c9i \u00d7 \u03b8\u0307 i (4) aCM i = aO i + \u03b1i \u00d7 rCM i + \u03c9i \u00d7 ( \u03c9i \u00d7 rCM i ) (5) where \u03c9i+1 and \u03c9i are the angular velocities of the (i + 1)-frame (at the base of the link, Fig. 6) and i-frame (at the top of the link, Fig. 6), respectively. The i-frame is fixed to the link and the (i + 1)-frame moves with the joint. vO is the velocity of origin of its frame and rL i is the position of the (i + 1)-frame relatively to the i-frame. aO i and aCM i are the accelerations of the origin of the frame and of the Center of Mass (CM), respectively. \u03b1 is the angular acceleration of the frame, \u03b8\u0307 and \u03b8\u0308 are the angular velocity and acceleration of joint, and rCM i is the position of the CM relatively to the i-frame. With the frames defined in Fig. 6, we can also define the balances of Forces and Moments, Eq. (6) and Eq. (7): \u2211 F = Fi+1 + Fi = miaCM i (6) \u2211 M = Mi+1 + Mi + rL i \u00d7 Fi+1 = H\u0307O i (7) where F and M are the applied forces and moments, respectively, and H\u0307O i is the time derivative of the angular momentum at the origin of the i-frame, Eq. (8): H\u0307O i = IO i \u03b1i + rCM i \u00d7 miaCM i + \u03c9i \u00d7 IO i \u03c9i (8) where IO i is the inertia tensor of link relatively to the i-frame and mi is the link mass. 3.2 Spatial vectors and articulated-body inertia For the ABA, Featherstone introduced a new notation called Spatial Vectors and Spatial Matrices" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002579_s0001925900002341-Figure12-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002579_s0001925900002341-Figure12-1.png", + "caption": "FIGURE 12. Notation for the tests under axial load.", + "texts": [ + " and the results of the tests on shouldered shafts subjected to torsion and to bending have been given in the previous two parts'1-2>. The present paper includes the results of the tests on shafts subjected to axial load, together with details of a supplementary investigation into the limiting value of the s.c.f., for a particular shoulder radius, as the depth of the shoulder is increased to infinity. A summary of the general conclusions for all forms of loading is given at the end of the paper. NOTATION (see Fig. 12) D large diameter of shaft d small diameter of shaft r shoulder root radius h shoulder depth (thus D = d + 2h) r , e? U; t > e@ (11) A B; t X t K X t 1 > e? U; t > e@ (12) The outer boundary would look like to \u201cFig (2)\u201d in the K; K plane. It would be referred to as \u201cHourglass\u201d. e? and e@ are errors which are initialized by a guess. For every two time step, an area is obtained from the intersection of two Hourglasses. An estimation of their overlap is demonstrated for whole simulation time in \u201cFig (3)\u201d. If the overlap doesn\u2019t exist, the boundary should be widen by increasing e? and e@. So, finally one could get the upper and the lower boundaries for optimal gains from the overlap boundaries or initial guess from the overlap center." + ] + }, + { + "image_filename": "designv11_33_0001967_jmech.2017.23-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001967_jmech.2017.23-Figure1-1.png", + "caption": "Fig. 1 The general dynamic model of the drivetrain system.", + "texts": [ + " Finally, the dynamic behavior and the disc brake parameters effecton the transmission error of the geared system were studied. Besides, the system resonance problemwas investigated using the sweep frequency analysis. 2. DRIVETRAIN MODELING 2.1 Model Description The drivetrain system is an essential element in a vehicle. The main function of the driveline is to convey power from the vehicle\u2019s engine to the drive wheels, through a transmission system. The drivetrain system consists of an engine, clutch, gearbox and disc brake. The general dynamic model of the drivetrain system is shown in Fig. 1. This model is represented by twenty two degrees of freedom. The system is made up of four main blocks. The first presents the engine; the second defines the vehicle clutch; the third describes the two stages of helical geared system and the last one represents the disc brake system. Since the shafts and bearings transmit the mechanical power from the engine to the other components, they will be taken into account in the modeling of the system. Each subsystem model is described in more details in the next section" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001448_978-94-007-1664-3_3-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001448_978-94-007-1664-3_3-Figure1-1.png", + "caption": "Fig. 1 Regulated aircraft damages used in its damage tolerance analysis.", + "texts": [ + " First of all in accordance with above-mentioned Russian idea \u201cto ensure acceptable damage tolerance characteristics satisfying the requirements considered design of the structure should be performed basing on the definite criteria of tolerable damage sizes and their propagation period\u201d [1]. Specific damage tolerance parameters were defined from the requirement that the calculated number of flight hours per one accident of the airplane with design goal of 50000 hours would be of 108 hours [1]. Recommended (regulated) damages are presented in Fig. 1 [1], [44]. In the stiffened structures those are two-bay skin cracks with broken central stiffener (stringer in the wing and fuselage, frame in the fuselage). Wing structure should carry limit load under simultaneous failure of spar chord and the panel skin crack of one inter-stringer distance size, as well as the spar web crack equal to the half of the web height. In the fuselage cut-out zones it is reasonable to consider the cut-out edge failure and adjacent skin crack of one inter-stringer (inter-frame) distance size to be a tolerable damage" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002327_ecce.2017.8096694-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002327_ecce.2017.8096694-Figure1-1.png", + "caption": "Fig. 1.- Schematic representation of the machine and sensors location.", + "texts": [ + " The paper is organized as follows: principles of the rotor position estimation using analog hall-effect sensors and 978-1-5090-2998-3/17/$31.00 \u00a92017 IEEE 3966 experimental results to demonstrate the viability of the concepts are presented in section II. A discussion of practical implementation issues is presented in section III, while conclusions are provided in section IV. II. Speed/position estimation using analog hall-effect sensors This section analyzes the principles of speed/position estimation using analog hall-effect sensors. Fig. 1 shows a schematic representation of the PMSM that will be used for the analysis and development of the method. The details of the ratings and dimensions of the test machine are shown in Table I. Fig. 2 shows the location of the three hall-effect sensors and the xyz coordinate system defined for each sensor. y-axis aligns with the radial direction, x-axis having a tangential direction. Magnetic flux density in the z-axis (axial direction) does not contain useful for position/speed estimation and will not be discussed further. As it can be observed from Fig. 1, Hall-sensors a, b and c are aligned with the field induced by the current flowing throughout phases a, b and c respectively, i.e. each Hall-sensor (a, b or c) is located at the geometric center of the corresponding phase coil. The phase shift among sensors is therefore 120 electrical degrees, which corresponds to 40 mechanical degrees in a 6-pole machine. Standard hall-effect sensors used in PMSM drives measure the magnetic flux density along one direction, i.e. they are 1D sensors. Typically y direction is used [3]-[6]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003003_iciibms.2018.8549921-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003003_iciibms.2018.8549921-Figure1-1.png", + "caption": "Fig. 1. Features from three axis", + "texts": [ + " For example, not switching to control home devices manually but just doing an action anywhere I am regardless of the environment, for example, brightness, camera, and others. It is a convenient user interface compared to the existing approaches. For this, we define some specific actions or gestures to control home devices, where a sensor named gyroscope in a smartphone is used to capture the gesture data representing user's motions. Using a smartphone is a relatively cheap and easy way to apply to a real environment. Using a gyroscope sensor, we can get some features, angular speeds by three axes, as shown in Figure 1. According to the speed, the value will change: High speed produces a big value, and slow speed makes a small value. In chapter 2, we discuss gesture recognition for home automation and experimental results in the next chapter. We will conclude the research in chapter 4 II. GESTURE RECOGNITION FOR HOME AUTOMATION In this research, we defined six types of gestures or motions to control home devices as below: As shown in Figure 2 horizontal grip and moving up and down is for TV ON and OFF. Vertical grip and moving up and down is to turn and off the air conditioner" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000954_cp.2012.0266-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000954_cp.2012.0266-Figure1-1.png", + "caption": "Fig. 1 Six-switch three-phase (SSTP) inverter topology. (a) Diagram. (b) Voltage vectors in ( ) plane.", + "texts": [ + "ue to many advantageous characteristics, interior permanent magnet (IPM) brushless AC (BLAC) machine is widely used in many applications [1, 2]. Fig. 1 presents a conventional IPM BLAC machine drive fed by a six-switch three-phase (SSTP) inverter where vector control (VC) technique together with maximum torque per ampere (MTPA) operation is often employed [3]. Basically, under VC methodology, electromagnetic torque of IPM BLAC machine is indirectly regulated via controlling stator phase currents to achieve high performance operation [4, 5]. Unlike VC strategy, both electromagnetic torque and stator flux linkage of IPM BLAC machine are simultaneously controlled under DTC technique [2, 6]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003881_012010-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003881_012010-Figure5-1.png", + "caption": "Figure 5. The selected case for 3D metal printing: Top left figure = the progressive die for the C-Bow Lower, top right figure = the station of this progressive die that was selected for 3D metal printing, and bottom figures = the shapes and dimensions of the selected punch & puller.", + "texts": [ + " This is illustrated in Figure 4. In this investigation, the tool concept comprises maraging steel (DIN 1.2709) 3D-printed both solidly and after topology optimization, and hardened to 55 HRC (see Table 2). The selected sheet material is 2-mm thick hot-dip galvanized DP600 (see Table 4). Due to the size limitations (in 3D metal printing by LPBF) mentioned above, the puller and the punch, which constitute a working station of the progressive die for the car body part C-Bow Lower, were selected. See Figure 5. The part, C-Bow Lower, is made in 1-mm thick hot-dip galvanized D600. This progressive die has been used (produced parts) in a couple of years. The puller and the punch shown in Figure 5 were selected for 3D-printing to evaluate the lead time, costs and performance compared to those of the existing conventionally made versions. Figure 6 displays the requirements set and the materials and manufacturing processes used for the conventional and 3D-printed versions of the puller and punch for the C-bow Lower progressive die in Figure 5. As shown in Figure 6, the requirements are the same, regardless of how the puller and punch are manufactured. The manufacturing process for the conventionally made puller and punch is the process that was used to make these portions of the die, as the progressive die for the C-Bow Lower was made. The 3D printed versions of the puller and punch were made in this study to compare 3D printing with conventional toolmaking. International Deep Drawing Research Group 38th Annual Conference IOP Conf. Series: Materials Science and Engineering 651 (2019) 012010 IOP Publishing doi:10", + " Both the solid and topology optimized tools managed 100,000 strokes in 2-mm thick DP600 with a burr height lower than 0.2 mm and were thereby approved. After 100,000 strokes, the maximum wear measured as a change in the profile radius was 0.100 mm on the solid tool and 0.196 mm on the topology optimized tool. Compared to the 3D-printed solid tool, the topology optimized and 3D-printed tool exhibits a lead time reduction by 29.6%. See Figure 10. Figure 11 displays the 3D-printed puller and punch in the progressive die shown in Figure 5. The puller and the punch were 3D-printed simultaneously (the same print) in maraging steel DIN 1.2709 (see Tables 1 & 2). It was selected to print a so-called honeycomb inner structure with a facade/outer shell thickness of 1.5 mm. Machining tests were conducted on the puller by milling at three different cusp heights \u2013 6 \u00b5m, 3 \u00b5m and 0.6 \u00b5m. No problems were encountered and the milling yielded the expected results. 2D and 3D surface roughness measurements were conducted directly after 3D printing and after 3D printing followed by milling at the above-mentioned three cusp heights. The surface roughness was Ra = 4.92 \u00b5m (Sa = 5.23 \u00b5m) directly after 3D printing and Ra = 0.71 \u00b5m (Sa = 0.85 \u00b5m) after 3D printing and milling at a cusp height of 0.6 \u00b5m. The average hardness was 56 HRC. Table 5 depicts a comparison of the lead time and total costs for the puller and the punch in Figure 5 made conventionally and by the 3D-printing inclusive manufacturing process (Figure 6). As displayed in Table 5, the lead time is reduced from 8 days for the conventionally made puller and punch to 3.7 days for the 3D-printing inclusive manufacturing of the same tools. The total costs (comprising material, machine, salary, and logistics costs) increase, as displayed in Table 5, from 26,000 to 31,000 SEK, in case the 3D-printing inclusive process is selected. The total cost of the 3D-printing inclusive process is based on the assumption that the depreciation period for the 3D printing machine is 5 years" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001440_acc.2011.5991404-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001440_acc.2011.5991404-Figure1-1.png", + "caption": "Fig. 1. Sketch of the MSA coordinate system in 2D.", + "texts": [], + "surrounding_texts": [ + "Information on the concentration is realized via a sensor attached to a mobile agent which provides (noise-corrupted) values of the concentration c(t,X, Y ) at a spatial point (Xs, Ys) of the domain \u2126. Therefore an adequate sensor model takes the form y(t) = c(t,Xs, Ys) = \u222b LX 0 \u222b LY 0 \u03b4(X \u2212Xs)\u03b4(Y \u2212 Ys)c(t,X, Y ) dX dY. (4) To incorporate the effects of a possibly moving sensing device, the sensor centroid is explicitly taken to be timedependent and thus y(t) = \u222b LX 0 \u222b LY 0 \u03b4(X\u2212Xs(t))\u03b4(Y\u2212Ys(t))c(t,X, Y ) dX dY." + ] + }, + { + "image_filename": "designv11_33_0003852_j.mechmachtheory.2019.103628-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003852_j.mechmachtheory.2019.103628-Figure3-1.png", + "caption": "Fig. 3. The geometry of tooth surface and grinding wheel.", + "texts": [ + " Partial derivatives of vector function V g l are taken with respect to variables \u03b8 and b , then coefficients of the first fundamental form of tooth surface can be obtained: E = \u2202V g l \u2202b \u00b7 \u2202V g l \u2202b = 0 F = \u2202V g l \u2202b \u00b7 \u2202V g l \u2202\u03b8 G = \u2202V g l \u2202\u03b8 \u00b7 \u2202V g l \u2202\u03b8 (15) Then partial derivatives of second-order are taken to get coefficients of the second fundamental form of tooth surface: L = n g c \u00b7 \u2202 2 V g l \u2202 b 2 M = n g c \u00b7 \u2202 2 V g l \u2202 b\u2202 \u03b8 N = n g c \u00b7 \u2202 2 V g l \u2202 \u03b82 (16) According to the differential geometry theory, the following two equations can be constructed using these coefficients: \u2223\u2223\u2223\u2223 \u03ban E \u2212 L \u03ban F \u2212 M \u03ban F \u2212 M \u03ban G \u2212 N \u2223\u2223\u2223\u2223 = 0 (17) \u2223\u2223\u2223\u2223 Edb + Fd \u03b8 Fdb + Gd \u03b8 Ldb + Md \u03b8 Mdb + Nd \u03b8 \u2223\u2223\u2223\u2223 = 0 (18) where \u03ban ( n = 1 , 2 ) is the principal curvature of tooth surface, db and d \u03b8 are the derivatives of tooth surface equation along two parametric curves, which can represent the tangent direction of two parametric curves, as shown in Fig. 3. The two principal curvature can be obtained by solving Eq. (17) . Solving the Eq. (18) , two ratio values of db / d \u03b8 can be obtained, which can represent two principal directions g 1 and g 2 . Since the b curve of the tooth surface is a set of straight lines, the tangent direction of b curve is one of the principal directions of the tooth surface. Assuming this direction is g 1 , then the principal curvature \u03ba1 = 0 . The two principal curvatures of the grinding wheel are set to be \u03bap and \u03baw respectively, in which \u03bap is the curvature in the t p (tangent direction of generator of the grinding wheel), and \u03baw is the curvature in the t w (circumferential direction of the grinding wheel)", + " The following two equations should be satisfied by the parameters of grinding wheel: \u03bc \u2264 | \u03b7h | (31) r w \u2265 cos \u03bc/ \u03ba2 hr (32) where \u03b7t is the slope angle of principal curvature curve corresponding to the b curve at the tooth toe of concave side, \u03b7h is the slope angle of principal curvature curve corresponding to the b curve at the tooth heel of convex side. After obtaining the geometric parameters of the working side of grinding wheel, the position of the grinding wheel can be determined according to the contact state between the grinding wheel and the tooth surface, and then the CNC tool location of tooth grinding can be calculated. As shown in Fig. 3 , while the generator of grinding wheel is tangent to the b curve of the tooth surface, the tip of the grinding wheel should also be tangent to the root cone of the gear to ensure the whole tooth surface being ground. Therefore, the tooth root curve can be regarded as the guide curve during tool path planning. According to the Section 3 , the tooth surface equation of the FFHHG can be expressed as a vector function containing two variables \u03b8 and b . The root curve of the tooth is the \u03b8 curve corresponding to the tip of the cutter blade" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001893_insi.2017.59.3.143-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001893_insi.2017.59.3.143-Figure8-1.png", + "caption": "Figure 8. Structural model of the defect detector: (a) holistic structural model; (b) structural model of the magnetic concentrator; and (c) structural model of the annular permanent magnet", + "texts": [ + " The defect detection sensor proposed in this paper was designed based on the research conducted on the model simulation and the 146 Insight \u2022 Vol 59 \u2022 No 3 \u2022 March 2017 Insight \u2022 Vol 59 \u2022 No 3 \u2022 March 2017 147 flaws in the wire ropes were recreated, including a loss of metallic cross-sectional area of 5% and 3, 5, 7, 9 and 12 broken wires. The running speed of the wire rope relative to the sensor was 1 m/s. The detection signal waveform detected by the conventional magnetic flux leakage defect detection sensor is presented in Figure 10(a). Figure 10(b) presents the detection signal waveform collected by the magnetic concentrator model mentioned above. Figure 8(a) shows a holistic structural model of the sensor and Figure 8(b) shows the structural model of the magnetic concentrator. Figure 8(c) shows the structural model of the annular permanent magnet. An experimental prototype was developed and an online experimental detection platform was built based on the structural models mentioned above. As shown in Figure 9, an N35 permanent magnet was adopted in the magnetiser. It was 15 mm wide and 10 mm thick. Compared to a conventional N48 permanent magnet (with strong remanence), it is 10 mm narrower and over 20 mm thinner. In the experimental platform, different proposed sensor. From the Figures, it can be seen that the proposed sensor detected six flaws while the conventional sensor missed one flaw" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001144_vppc.2012.6422577-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001144_vppc.2012.6422577-Figure3-1.png", + "caption": "Fig. 3. Space vectors in the stationary and rotating reference frames", + "texts": [ + " In the space vector control system, the three-phase current is transformed to the d-q axis current by constructing a rotating reference frame. The negative-sequence current also rotates at synchronous speed, but its direction of rotation is opposite of motor's direction of rotation. Therefore, a reverse rotating reference frame is applied to calculate negative sequence current as well as calculation of d-q axis current. This calculation is easily implemented by negating the position signal from motor position sensor.Fig.3 shows relationship of different reference frames. In the new reference frame, the rotating negative sequence current vector is transformed into a stationary vector. The complex quantities of the stationary vector in is defined as In = Ind + jInq (14) The transformed is described as [. 1 [cos () :: \ufffd sinO 2 COs( () + - n\") 3 sine () + \ufffd n) 3 The negative sequence real part ind and complex part inq both contains direct components and alternating components, which is the mapping of positive sequence current" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003587_1350650119868910-Figure18-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003587_1350650119868910-Figure18-1.png", + "caption": "Figure 18. Noise test setup: (a) multipurpose transmission tribotester, and (b) bearing installation.", + "texts": [ + " This is due to the phenomenon that the increased rotational speed leads to an increment in the maximum film temperature, as illustrated in Figure 17. In turn, the increased temperature makes the lubricant viscosity drop, in which case the LCC of the lubricant decreases and further friction noise of the bearing increases. Therefore, a reasonable low rotational speed can be selected to reduce the APL of the compound textured bearings. With the multipurpose transmission tribotester, the noise experiment of the bearing at the varied rotational speed is made using the noise tester. Figure 18(a) gives the schematic of the multipurpose transmission tribotester, in which the control cabinet controls the power consumption of this equipment and the torque sensor transduces the torque provided by the driving motor. The base restrains the deformation and vibration of the machine to reduce the external excitation impact on the experimental result. Figure 18(b) shows the bearing installation in which the acceleration sensor records the vibration acceleration of the rotational shaft and the supply pipe filled lubricant into the bearing during the experiment. The material of the used bearing is GCr15, whose inner radius and outer radius are separately 25.0 and 29.0mm, and whose elastic module and Poisson\u2019s ratio are 208GPa and 0.30. The material of the shaft is 40Cr steel and H2050 bearing seat type is used for supporting the bearing. Besides, 11 textures are distributed on the inner surface bush, whose length and interval are 30mm and 10 , respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002377_978-3-319-68826-8_7-Figure7.68-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002377_978-3-319-68826-8_7-Figure7.68-1.png", + "caption": "Fig. 7.68 a Execution scheme of an elementary behaviour of GS type, b execution scheme of an elementary behaviour of OA type, c execution scheme of a complex CB behaviour", + "texts": [ + " These algorithms are called reactive algorithms, or sensor signal-based algorithms. Examples of such algorithms include systems generating behavioural (reactive) control, inspired by the animal world [9, 10, 38, 39, 47, 71, 72]. These are simple behaviours which can be observed, e.g. in insects. Among them, there is goal-seeking (GS) task (e.g. seeking for a food source), and \u201ckeep to the centre of the open space\u201d tasks which can be viewed as an obstacle avoiding (OA) task. The implementation schemes for individual elementary behaviours are presented in Fig. 7.68a, b. No behavioural control system, whether of OA or GS type, enables the implementation of a complex task such as \u201ccombination of behaviours OA and GS\u201d (CB), shown schematically in Fig. 7.68c. The implementation of such a task is possible by way of combination of GS and OA behavioural controls, along with a relevant switching function. Another examples of local methods of generation of robots\u2019 motion trajectories are e.g. potential fields method [1, 60, 62] or elastic band approach [8]. The sensory systemofWMRPioneer 2-DX,whose design is discussed inSect. 2.1, was equipped with an additional Banner LT3PU laser rangefinder, installed on the upper panel of the WMR frame and marked as sL in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003422_j.triboint.2019.06.022-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003422_j.triboint.2019.06.022-Figure3-1.png", + "caption": "Fig. 3. Orientation of valve train components.", + "texts": [ + " A constant lubricant supply pressure of 2 bars was maintained in the engine head with the help of a feedback Proportional Integral Derivative (PID) and hydraulic pump controller ABB-ASC 150 whereas the sensing of the lubricant pressure was accomplished by using a Piezoresistive transducer. The oil film thickness at cam/roller contact has been measured using electrical capacitance technique for which the roller finger follower assembly should be insulated electrically from the rest of engine head components. The orientation of under research valve train is shown in Fig. 3. If the hydraulic lash adjuster (HLA) and the valve cap are electrically insulated from the rest of the engine head, the roller finger follower assembly will be insulated, completely. The insulated roller of the assembly will act as one plate of the capacitor system whereas the lobe of electrically grounded camshaft would be the second plate to measure the oil film thickness. The electrical insulation of required engine valve train components to monitor film state is shown in Fig. 4. The outer diameter of barrel of HLA was reduced from 12mm to 11" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002583_1.4039943-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002583_1.4039943-Figure3-1.png", + "caption": "Figure 3. The effect of the 3D rotations on the OF values", + "texts": [ + " This is clearly illustrated in Figure 2; objects that are close to the camera seem to move faster than objects which are relatively farther. For example, the pixel\u2019s movement of the landing pattern in 2 (a) is faster than its movement in 2 (b). Therefore, in order to obtain an accurate OF measurement, an accurate measurement of the altitude has to be precisely measured. Similarly, the 3D rotational components of the camera must be accurately measured while computing the OF components. (a) (b) Figure 2. The correlation between the depth of scene and the OF values Figure 3 illustrates the effect of the 3D rotational motion of the helicopter on the OF readings. Figure 3 (a) shows the helicopter when no 3D rotation is applied, while in 3 (b) the helicopter is rolled by a small angle. Clearly, in 3 (a) the pattern fills the view completely while it is partially seen in 3 (b) due to the roll motion. Therefore, in the latter case, and assuming that no translational velocities are applied in both scenarios, the measured OF values have to be accurately compensated for the 3D rotations. In this study, the proposed modeling approach uses the PX4FLOW smart kit which has 3D gyros and ultrasonic sensor onboard to compensate for the 3D rotations and the altitude variations at each time instant" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000917_20110828-6-it-1002.02534-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000917_20110828-6-it-1002.02534-Figure6-1.png", + "caption": "Fig. 6. K(2) gain surface", + "texts": [], + "surrounding_texts": [ + "J. M. Biannic and P. Apkarian. Missile autopilot design via a modified LPV synthesis technique. Aerospace Science and Technology, 3(3):153\u2013160, 1999. ISSN 1270-9638. Stephen Boyd, Laurent El Ghaoui, Eric Feron, and Venkataramanan Balakrishnan. Linear matrix inequalities in system and control theory. SIAM, 1994. M. Chilali and P. Gahinet. H-infinity design with pole placement constraints: an LMI approach. IEEE Transactions on Automatic Control, 41(3):358\u2013367, 1996. A. Fujimori, F. Terui, and P. N. Nikiforuk. Flight control designs using v-gap metric and local multi-objective gain-scheduling. In AIAA Guidance, Navigation and Control Conference, 2003. AIAA 2003-5414. P. Gahinet, A. Nemirovskii, A.J. Laub, and M. Chilali. The lmi control toolbox. In Proceedings of the 33rd IEEE Conference on Decision and Control, pages 2038\u2013 2041, 1994. S. L. Gatley, D. G. Bates, M. J. Hayes, and I. Postlethwaite. Robustness analysis of an integrated flight and propulsion control system using \u00b5 and the \u03bd-gap metric. Control Engineering Practice, 10(3):261\u2013275, 2002. ISSN 0967-0661. Jung-Yub Kim, Sung Kyung Hong, and Sungsu Park. LMI-based robust flight control of an aircraft subject to CG variation. International Journal of Systems Science, 41(5):585\u2013592, 2010. C. H. Lee, T. H. Kim, and M. J. Tahk. Design of missile autopilot for agile turn using nonlinear control. In Proceeding of the KSAS Conference, pages 667\u2013670, 2009. D. J. Leith and W. E. Leithead. Survey of gain-scheduling analysis and design. International Journal of Control, 73(11):1001\u20131025, 2000. Kenneth H. McNichols and M. Sami Fadali. Selecting operating points for discrete-time gain scheduling. Computers & Electrical Engineering, 29(2):289\u2013301, 2003. ISSN 0045-7906. P. K. Menon and E. J. Ohlmeyer. Integrated design of agile missile guidance and autopilot systems. Control Engineering Practice, 9(10):1095 \u2013 1106, 2001. ISSN 0967-0661. Andy Packard. Gain scheduling via linear fractional transformations. Systems & Control Letters, 22(2):79\u2013 92, 1994. ISSN 0167-6911. Wen qiang Li and Zhi qiang Zheng. Robust gainscheduling controller to LPV system using gap metric. In International Conference on Information and Automation, pages 514\u2013518, 2008. Wilson J. Rugh and Jeff S. Shamma. Research on gain scheduling. Automatica, 36(10):1401\u20131425, 2000. ISSN 0005-1098. David Saussie, Lahcen Saydy, and Ouassima Akhrif. Gain scheduling control design for a pitch-axis missile autopilot. In AIAA Guidance, Navigation and Control Conference, 2008. AIAA 2008-7000. C. Scherer, P. Gahinet, and M. Chilali. Multiobjective output-feedback control via LMI optimization. IEEE Transactions on Automatic Control, 42(7):896\u2013 911, 1997. J. S. Shamma and James R. Cloutier. Gain-scheduled missile autopilot design using linear parameter varying transformations. Journal of Guidance, Control, and Dynamics, 16(2):256\u2013263, 1993. J.S. Shamma and M. Athans. Gain scheduling: potential hazards and possible remedies. IEEE Control Systems Magazine, 12(3):101\u2013107, Jun. 1992. Spilios Theodoulis and Gilles Duc. Missile autopilot design: Gain-scheduling and the gap metric. Journal of Guidance, Control, and Dynamics, 32(2):986\u2013996, 2009. Ajay Thukral and Mario Innocenti. A sliding model missile pitch autopilot synthesis for high angle of attack maneuvering. IEEE Transactions on Control Systems Technology, 6(3):359\u2013371, 1998. D.P. White, J.G. Wozniak, and D.A. Lawrence. Missile autopilot design using a gain scheduling technique. In Proceedings of the 26th Southeastern Symposium on System Theory, pages 606\u2013610, Athens, 1994. Kevin A. Wise and David J. Broy. Agile missile dynamics and control. Journal of Guidance, Control and Dynamics, 21(3):441\u2013449, 1998. G. Zames and K. El-Sakkary. Unstable systems and feedback: The gap metric. In Proceedings of the 18th Allerton Conference, pages 380\u2013385, 1980. Paul Zarchan. Tactical and Strategic Missile Guidance. Progress in Astronautics and Aeronautics Series, 2007. K. Zhou and J.C. Doyle. Essentials of Robust Control. Prentice Hall, 1997. P. H. Zipfel. Modeling and Simulation of Aerospace Vehicle Dynamics. AIAA Education Series, 2000." + ] + }, + { + "image_filename": "designv11_33_0002886_j.matpr.2018.06.311-Figure5.3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002886_j.matpr.2018.06.311-Figure5.3-1.png", + "caption": "Figure 5.3 : At C.R. 16.5 Fatigue life and Structure analysis", + "texts": [ + " / Materials Today: Proceedings 5 (2018) 19497\u201319506 19501 4.5: Variation Of Temperature With Crank Angle Structural Analysis Steady state thermal analysis consider the before structural analysis, apply boundary conditions (heat transfer coefficient, heat flux) on the piston and we observe the maximum and minimum temperatures, total heat flux of cylinder head can use a steady-state thermal analysis to determine temperatures, thermal gradients, heat flow rates, and heat fluxes in an object that are caused by thermal loads that do not vary over time.Figure5.3&-4 shows a steady-state thermal analysis calculates the effects of steady thermal loads on a system or component .At Compression Ratios 16.5:Apply the Boundary conditions on the cylinder head , we observe the results [14 -18] .The fixed supports are shown in detail in the figure and the pressure is applied at the inlet and outlet valves of the cylinder head as shown in above Fig 4.1 and below Fig 5.1 (a-b) 5.1 Based on thermal analysis: By applying the boundary conditions for 16.5 the heat transfer analysis is carried out as shown in below Fig 5", + "5 heat transfer analysis is carried out. 19502 K. Satayanarayana et al. / Materials Today: Proceedings 5 (2018) 19497\u201319506 5.2 Based on thermal analysis:- Fig (a) : Cylinder head Total heat flux Fig (b) . Cylinder head Temperature distribution: FIG: 5.2: At compression ratio 17.5 heat transfer analysis we observe from the plots maximum temperature in the cylinder head is 211.86 0C and minimum temperature in the cylinder head is135.66 0C.Maximum heat flux in the cylinder head is 1.25 w/m2 as show n in below Fig5.3 to fig 5.7 K. Satayanarayana et al. / Materials Today: Proceedings 5 (2018) 19497\u201319506 19503 Fig (a) Structural Analysis: Vonmisses Stresses Fig (b) Total deformation Fig 5.4 At compression ratio 17.5 Structural analysis For compression ratio 18.5 based on thermal analysis: 19504 K. Satayanarayana et al. / Materials Today: Proceedings 5 (2018) 19497\u201319506 Fig (a) Cylinder head vonmisses stresses Fig (b) Cylinder head total deformation Fig5.6. FIG: At compression ratio 18.5 Structural analysis For structural analysis ,the cylinder head pressure of 51" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003584_icra.2019.8793779-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003584_icra.2019.8793779-Figure2-1.png", + "caption": "Fig. 2: Hands (Barrett, PR2, Shadow) and added noise for the validation test. Top row, moving the object, middle row, joint angle variation, bottom row, grid samples.", + "texts": [ + " From this, a grid of points is generated in the tangent space at the desired resolution (generateGrid). Each point is then projected onto the robot geometry in the direction of the tangent plane\u2019s surface normal (morphPoints). This defines a set of points (from the robot geometry). This projection only has to happen once. For the results presented here, we used a 20 \u00d7 20 grid sampling rate, one grid for each finger link, placed so that it spans the link, and one grid for the palm (Barrett and Shadow, see Figure 2). We calculated these values shortly before the fingers made contact. Algorithm GENERATE GRID() Input: Mesh mhand X,Y Resolution N User specified points pl, xl, yl, l \u2208 [1, N ] Output: Grid Points linkli,j , l \u2208 [1, N ], i \u2208 [1, X], j \u2208 [1, Y ] 1: for i in [1, N ] do 2: gplane \u2190 generateGrid(mhand,pi, xi, yi, X , Y ) 3: linkli,j \u2190 morphPoints(linkli,j , mhand, gplane) 4: return {linki,j}1,N The signed distance feature requires the pre-computation of a Truncated Signed Distance Field (TSDF) d : R3 \u2192 R of the object", + " Validation test 1: Comparison to existing metric We describe the hands, objects, and grasps used, the noise model used to generate variations, and the deep learning model. Note that the goal is to learn the expected value of the \u03b5-metric \u2014 mean and standard deviation. 1) Hands, objects, and grasps: To demonstrate the handagnostic nature of our feature representations we conduct three experimental simulations using hands with varying kinematic structures: The fully anthropomorphic 20-DOF Shadow Hand, the opposable 4-DOF Barrett Hand, and the PR2 Gripper, (see Figure 2). We placed grids on each link: 2 for the PR2, 7 for the Barrett, and 15 for the Shadow Hand. The objects (shown in Figure 3) used to perform our experiments were selected from a set of fundamental shapes, with three sizes for each one. Our grasps were specified as pre-shapes (the grasps shown in Figure 2). To generate the final pose, the grasp is executed in GraspIt!, which essentially closes the fingers at equal speeds until they make contact with the object. This does not reflect \u201creal life\u201d, because the object would normally shift when grasped, but it suffices for these tests. We calculate the \u03b5-metric for the final grasp, and use the initial hand configuration (pre-noise) to compute our representations. 2) Noise generation: For all three hands we introduced noise by changing the pose of the hand relative to the object" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001856_1077546316689745-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001856_1077546316689745-Figure1-1.png", + "caption": "Figure 1. Physical model of the gear pair in the rattle condition: (a) \u2018\u2018correct\u2019\u2019 contact, (b) \u2018\u2018no\u2019\u2019-contact, and (c) \u2018\u2018incorrect\u2019\u2019 contact.", + "texts": [ + " Before describing the signal processing procedure, a brief reminder about rattle modeling, and particularly about the shape of a sample representing the time history of the gear relative motion, has to be performed. The relative motion of an unloaded gear pair under rattle conditions can be modeled, along the line of contacts, by referring to a single-degree-of-freedom system, considering the driven wheel gear forced to vibrate by a periodic motion imposed at the driving gear, as reported by Wang et al. (2001). Figure 1 shows the scheme and physical model of a gear pair, where the subscript 1 indicates the driving gear and subscript 2 refers to the driven wheel. The figure shows three characteristic situations occurring during rattle; in particular, Figure 1(a) shows the \u2018\u2018correct contact\u2019\u2019 when the pinion pushes on the driven wheel; in Figure 1(b) no contact takes place because the driving gear decelerates in respect to the driven wheel, and so the only interaction is due to a damping effect of the oil lubricant interposed between the teeth (Brancati et al., 2005); finally, in Figure 1(c) the driven wheel bumps on the driving gear due to its inertia, determining an \u2018\u2018incorrect contact\u2019\u2019 phase, anomalous for the transmission of motion. Under rattle conditions, being the contacts alternated along the two flanks of teeth, a time history of the gear relative motion along the contact line takes the form of a series of trapezoidal waves. Figure 2 shows, as an example of what has been said above, a typical time history of four rattle cycles obtained respectively from an experimental test and a theoretical simulation code" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002218_17515831.2017.1378852-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002218_17515831.2017.1378852-Figure1-1.png", + "caption": "Figure 1.\u00a0(a) schematic of two rotating bodies coming into contact, positive separation. (b) close up view of the contacting region, negative separation.", + "texts": [ + " Contact mechanics are also considered at the interface between the two bodies under frictionless conditions, no flow is permitted D ow nl oa de d by [ G ot he nb ur g U ni ve rs ity L ib ra ry ] at 0 6: 15 0 1 N ov em be r 20 17 across the interface and as such fluid is confined to each of the contacting bodies and dry conditions apply. In this paper, the contact between two rotating bodies is investigated, the case studies conducted analyse the following: (i) soft elastic contact; and (ii) poroelastic contact. The remainder of this section outlines the equations, parameters and assumptions which define each of these models. In Figure 1(a) two elastic bodies with given inner R1,i, R2,i and outer radii R1,o, R2,o are rotating with angular velocities \u03a91, \u03a92 about their origins. Figure 1(b) illustrates that when the separation \u03b4 between the two bodies is reduced beyond zero they come into contact and are deformed such that there is no penetration. As a result, the contacting region is defined in x by where y = 0 and in which the contact pressure is positive, outside this region the contact pressure is zero. Since the angular velocities of the bodies are constant the elastic deformation in contact can be calculated using steady-state assumptions in which the bodies are compressing and sliding against each other" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003893_0954406219888240-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003893_0954406219888240-Figure7-1.png", + "caption": "Figure 7. Assembly drawing of universal joint.", + "texts": [ + " Accordingly, the values of two end points of Side A in the o-UVW coordinate system are A1 \u00bc d 2 , c 2 , 0 T , A2 \u00bc d 2 , c 2 , L4 T \u00f02\u00de In the o-uvw coordinate system, the values of two end points of Side B are B01 \u00bc c 2 , d 2 , 0 T , B02 \u00bc c 2 , d 2 ,L4 T \u00f03\u00de According to the mapping relation of two coordinate systems, the values of two end points of Side B in the o-UVW coordinate system can be solved as B1 \u00bc R1B 0 1 \u00bc \u00bdB1U,B1V,B1W T, B2 \u00bc R1B 0 2 \u00bc \u00bdB2U,B2V,B2W T \u00f04\u00de where B1U \u00bc c 2 cos \u00fe d 2 sin sin , B1V \u00bc d 2 cos , B1W \u00bc c 2 sin \u00fe d 2 cos sin , B2U \u00bc c 2 cos \u00fe d 2 sin sin \u00fe L4 sin cos , B2V \u00bc d 2 cos L4 sin , B2W \u00bc c 2 sin \u00fe d 2 cos sin \u00fe L4 cos cos . Non-interference domain of the universal joint In order to determine the relationship between the workspace and the geometric parameters, a new assembly drawing is given as shown in Figure 7, where the assemble dimensions are involved, and four assumptions are given: 1. The block of cross trunnion is ideal and no interference occurs between the cross trunnion and two hinges; 2. The gap c is always greater than the diagonal diameter d: c> d; 3. The length L4 is greater than or equal to the half of diagonal diameter: L45 d/2, where we assume that the length L4 and the half of the diagonal diameter d are equal; 4. The ranges of two Euler angles ( , ) are limited within: [ p/2, p/2]. In the study by Zhang et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002595_j.triboint.2018.05.003-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002595_j.triboint.2018.05.003-Figure1-1.png", + "caption": "Fig. 1. Ball-on-plane contact geometry used to represent ball-on-disk experiments by Fryza et al. [1].", + "texts": [ + " The numerical model employed assumes isothermal, time-dependent conditions. The effect of the lubricant rheology is also investigated by means of the Eyring [25] and the modified Carreau-Yasuda [26] nonNewtonian constitutive equations. Film thickness predictions are then directly compared with those measured by Fryza et al., both in quantitative and qualitative manner. The ball-on-disk configuration used in the experiments by Fryza et al. is represented here by a ball-on-plane contact geometry. As illustrated in Fig. 1, the stationary rolling entrainment velocity is along the x-axis, whereas the transverse sliding speed on the ball acts along the yaxis direction. The lubricant flow between the ball and the plane is described by the generalized Reynolds equation introduced by Najji et al. [27] for isothermal point contacts with non-Newtonian fluid rheology: \u239c \u239f \u239c \u239f\u2207 \u2207 = \u2202 \u2202 \u23a1 \u23a3 \u23a2 \u239b \u239d \u2212 \u2032 \u239e \u23a0 \u23a4 \u23a6 \u23a5 + \u2202 \u2202 \u23a1 \u23a3 \u23a2 \u239b \u239d \u2212 \u2032 \u239e \u23a0 \u23a4 \u23a6 \u23a5 + \u2202 \u2202 p x u \u03b7 \u03b7 u h \u03c1h y v \u03b7 \u03b7 v h \u03c1h t \u03c1h(\u03b5 ) ( )e e s e e s 2 2 (1) where: = \u239b \u239d \u239c \u2032 \u2212 \u2032 \u239e \u23a0 \u239f\u2032\u03c1 \u03b7 \u03b7 \u03b7 \u03b5 1 e e e 2 and Notation a Hertz contact radius, m aCY Carreau-Yasuda parameter \u2032E Reduced elastic modulus, Pa fy Oscillation frequency for v2, Hz fy Oscillation stroke length for v2, m G Carreau-Yasuda parameter, Pa h Lubricant film thickness, m h0 Rigid body normal approach, m ha Average film thickness, m hc Central film thickness, m hm Minimum film thickness, m Har Average film thickness ratio, = =h t h t( )/ ( 0)a a Hcr Central film thickness ratio, = =h t h t( )/ ( 0)c c Hmr Minimum film thickness ratio, = =h t h t( )/ ( 0)m m nCY Carreau-Yasuda parameter p Lubricant pressure, Pa ph Hertz contact pressure, Pa Rx Reduced x-axis radius of curvature, m Ry Reduced y-axis radius of curvature, m t Time, s u Mean surface velocity along x-axis, = +u u 2 1 2m \u22c5 s-1 u1 Lower surface velocity along x-axis, m \u22c5 s-1 u2 Upper surface velocity along x-axis, m \u22c5 s-1 ur Reference speed, m \u22c5 s-1 us Sliding velocity along x-axis, = \u2212u u2 1, m \u22c5 s-1 v Mean surface velocity along y-axis, = +v v 2 1 2m \u22c5 s-1 v1 Lower surface velocity along y-axis, m \u22c5 s-1 v2 Upper surface velocity along y-axis, m \u22c5 s-1 vs Sliding velocity along y-axis, = \u2212v v2 1, m \u22c5 s-1 W Applied normal load, N x Coordinate, m y Coordinate, m z Coordinate, m \u0391 Pressure-viscosity coefficient, Pa -1 \u03b3\u0307e Effective lubricant shear rate, s-1 \u03b5 Dimensionless group \u03b7 Lubricant effective viscosity, Pa \u22c5s \u03bc Lubricant Newtonian viscosity, Pa \u22c5s \u03bc2 Carreau-Yasuda second Newtonian viscosity, Pa \u22c5 s \u03bcr Reference viscosity, Pa \u22c5 s \u2207 Nabla operator \u03a1 Lubricant density, kg \u22c5m-3 \u03c1r Reference density, kg \u22c5 m-3 \u03c40 Eyring lubricant non-Newtonian stress, Pa \u03c4e Effective lubricant shear stress, Pa Zr Roelands equation parameter \u222b \u222b \u222b= \u2032 = \u2032 =\u2032\u03b7 dz \u03b7 \u03b7 z \u03b7 dz \u03b7 z \u03b7 dz1 , 1 , 1 e h e h e h 0 0 0 2 The above integrals take into account the variation of viscosity across the film due to the shear-thinning non-Newtonian behaviour of the lubricant, which is represented either by the Eyring sinh constitutive equation as proposed by Johnson and Tevaarwerk [25]: \u239c \u239f= \u239b \u239d \u239e \u23a0 \u03b3 \u03c4 \u03bc \u03c4 \u03c4 \u02d9 sinhe e0 0 (2) or by the modified Carreau-Yasuda equation described by Bair [26]: = + \u2212 \u23a1 \u23a3\u23a2 + \u239b \u239d \u239e \u23a0 \u23a4 \u23a6\u23a5 \u2212 \u03b7 \u03bc \u03bc \u03bc \u03c4 G ( ) 1 e aCY nCY aCY 2 2 1 1/ (3) with the effective shear rate and shear stress respectively defined as: = + = +\u03b3 \u03b3 \u03b3 \u03c4 \u03c4 \u03c4\u02d9 \u02d9 \u02d9 ,e zx zy e zx zy 2 2 2 2 (4) The individual shear rate components are given by: \u239c \u239f \u239c \u239f = \u239b \u239d \u2212 \u239e \u23a0 + = \u239b \u239d \u2212 \u239e \u23a0 + \u2202 \u2202 \u2202 \u2202 \u03b3 z u \u03b3 z v \u02d9 \u02d9 zx \u03b7 p x \u03b7 \u03b7 \u03b7 \u03b7 s zy \u03b7 p y \u03b7 \u03b7 \u03b7 \u03b7 s 1 1 e e e e e e ' ' (5) and the corresponding shear stresses by the relationships: = = \u03c4 \u03b7 \u03b3 \u03c4 \u03b7 \u03b3 \u02d9 \u02d9 zx zx zy zy (6) From these equations, Newtonian lubricant behaviour simply follows from the equality =\u03b7 \u03bc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000537_978-3-642-02525-9_6-Figure6.11-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000537_978-3-642-02525-9_6-Figure6.11-1.png", + "caption": "Fig. 6.11 Schematic illustration of the dielectrophoretic assembly of nanowires onto a template with additional 3-D structures formed in a resist. After assembly, the correctly assembled nanowires are fixed in a plating process and the resist is removed (after [6.44], c\u00a9 Macmillan 2008)", + "texts": [ + " The system was designed in such a way that 3 \u03bcm polystyrene particles were guided towards areas of weakest electrical field, which was directly underneath the lines of an interdigitated electrode array on the opposite substrate. There, the particles were permanently bonded by a chemical reaction. Dielectrophoretic assembly lends itself very well to the assembly of nonisotropically shaped objects such as nanorods or nanowires. The electric field can additionally align the wires in a desired orientation. Mayer and coworkers have applied this method to align semiconductor and metal nanowires on substrates [6.44]. The electrodes were covered with photoresist, which had openings at the desired binding sites (Fig. 6.11). Nanowires were directed to and adsorbed onto those sites. The topographic structure of the assembly sites helped to maintain the wires in the correct positions upon drying. The assembled wires were fixed in a plating process and lift-off of the resist layer removed those nanowires that adsorbed onto undesired positions. This combination of methods can significantly reduce the error count and increase the yield of the assembly process (Fig. 6.12). Part A 6 .2 A very versatile variant of dielectrophoretic assembly was demonstrated by Chiou and coworkers [6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000837_jfm.2012.45-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000837_jfm.2012.45-Figure1-1.png", + "caption": "FIGURE 1. (Colour online available at journals.cambridge.org/flm) (a) Schematic diagram for a moving object in the inertial and the body-fixed reference frames, OX1X2 and ox1x2. The point of interest, Q, is at z in ox1x2 but at Z = Zo + ei\u03b8z in OX1X2. Here, Zo is the position of the moving reference frame in OX1X2 and \u03b8 is the angle between OX1 and ox1. (b) Illustration of the integration contour: on the object surface C or around the internal singularities, C \u2217j .", + "texts": [ + " Finally, \u00a7 5 summarizes this work and comments on possible directions for future research. The conventional two-dimensional Lagally theorem that considers sources and dipoles is revisited here in terms of complex variables to include free vortices, following the approach and the sign convention of Milne-Thompson (1968). Let a solid object move and rotate at U and \u2126 relative to a fixed inertial reference frame OX1X2 in an unbounded inviscid and incompressible fluid at rest at infinity, as depicted in figure 1(a). A Cartesian coordinate system ox1x2 is fixed to a point o on the body and its position relative to OX1X2 is Zo. The positions of a point Q in both reference frames, Z and z, are related by Z = Zo + ei\u03b8z, where \u03b8 is the angle between OX1 and ox1. In the following derivations, variables in the moving reference frame and any function in terms of the variables therein will be denoted with over-tilde. The flow field, v\u0303 = v\u03031 + iv\u03032, is irrotational in OX1X2 and can be described by v\u0303 = \u2212dw\u0303/dz, where the complex velocity potential, w\u0303 = \u03c6\u0303 + i\u03c8\u0303 , is composed of the potential and the stream functions, \u03c6\u0303 and \u03c8\u0303 ", + " The second term can generate the added mass force, A\u03b1j dU\u03b1/dt, for unsteady object motion but remains non-zero for a steady object due to U\u03b1 dA\u03b1j/dt. This latter force component arises as a result of changes in object relative configuration. The last term of (2.8) accounts for the temporal variations of the strengths and positions of the internal image singularities. By transforming the line integral along C onto circles, C \u22171 , C \u22172 , . . . and C \u2217j , centred at each internal singularity (see figure 1b), the integrals in (2.8) can be evaluated by the residue theorem. The final force formula in OX1X2 is found, in terms of the property of each internal singularity, as F =\u22122\u03c0\u03c1 \u2211 j all ( \u00b5j dfj dZ \u2223\u2223\u2223\u2223 Zd j + mj fj|Zs j \u2212 i\u03baj fj|Zvj ) \u2212 d dt 3\u2211 \u03b1=1 [U\u03b1(A\u03b11 + iA\u03b12)] \u2212 2\u03c0\u03c1 d dt \u2211 j img (\u00b5j + mjZ s j \u2212 i\u03bajZ v j ). (2.9) The function fj = \u2212d(w \u2212 wj)/dZ in the first summation is the flow velocity induced by all the other singularities at an internal singularity location Zj. Thus, each fj is different and will be referred to as the partial velocity function hereinafter", + " (2007), where the equations of motion for both the cylinder and the free vortex are given explicitly after contour integration of the conventional Blasius theorem. With the current generalized Lagally theorem, identical formula describing the relative motion can be obtained from (3.4) using Nv = 1, \u03ba1 = \u03ba and L0 = 0 as dZc dt = i\u0393 \u03c0a2(\u03c1s/\u03c1 + 1) [ Zv1 \u2212 a2 l2 1 (Zv1 \u2212 Zc) ] , (3.5) where \u0393 = 2\u03c0\u03ba is the circulation of the free vortex. Hence, the evolution trajectories for the point vortex and the cylinder presented in the figure 1 of their work can be exactly reproduced. 3.3. With a vortex pair: the moving F\u00f6ppl problem Next, we consider the moving Fo\u0308ppl problem, where a pair of vortices of opposite strengths, \u03ba1 = \u2212\u03ba2 = \u03ba , move behind a non-rotating cylinder of radius a and density \u03c1s along OX1 at velocity U, as depicted in figure 2. This vortex pair is in a symmetric configuration with respect to OX1 with a half-span of H and a horizontal distance D from Zc. The complementary set-up where a uniform flow passes a fixed cylinder with a downstream vortex pair is the classic Fo\u0308ppl problem, well examined in the literature (Fo\u0308ppl 1913; Milne-Thompson 1968; Marshall 2001)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000871_20120919-3-it-2046.00024-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000871_20120919-3-it-2046.00024-Figure2-1.png", + "caption": "Fig. 2. AUV coordinate systems", + "texts": [ + " Section 5 presents conclusions of the current work. The vehicle is a torpedo-like vehicle, 1.742 m long, with 0.234 m maximum cross-section diameter and 52.5 kg mass. It is equipped with a thruster for cruising and fully moving control surfaces (rudder and stern), to steer the vehicle in marine environment. The complete AUV nonlinear model is composed of six differential nonlinear ordinary equations that represent the dynamics of the underwater vehicle and six equations for coordinate transformations between inertial frame and body frame. Figure 2 shows the inertial frame for position and orientation coordinates (x, y, z, \u03c6, \u03b8 and \u03c8 ) represented by the vector \u03b7, and a moving reference system fixed at the gravity centre of the vehicle, used to measure linear and angular velocities of the AUV (u, v, w, p, q and r) represented by the vector \u03bd. Capital letters (X , Y , Z, K, M and N) are used for representing the external forces and moments (SNAME, 1950). Therefore, the dynamic model of the vehicle is described in six DOF according relative to the body and inertial frames (Fossen, 2002): M\u03bd\u0307 + C(\u03bd)\u03bd + D(\u03bd)\u03bd + g(\u03b7) = \u03c4 (1) \u03b7\u0307 = J(\u03b7)\u03bd (2) The dynamic equation (1) is derived with the principles of Newtonian and Lagrangian Mechanics (Fossen, 2002), where M is the system inertia matrix that includes the mass added and rigid body mass terms, C(\u03bd) is the Coriolis-centripetal matrix that includes added mass terms, D(\u03bd) is the damping matrix, g(\u03b7) is the vector of gravitational forces and moments, and \u03c4 is the vector of control inputs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003771_012024-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003771_012024-Figure3-1.png", + "caption": "Figure 3. 10 MW drivetrain MBS model", + "texts": [ + " The MBS model are composed of various rigid and flexible bodies, and the connections between the bodies are modelled with joins or force elements. Dynamic behaviors of mechanical or electromechanical system can be described or predicted by conducting the MBS simulation that has been widely used in the fields of vehicles, robots, and wind turbines, etc. In this study, the numerical model of the 10 MW medium speed drivetrain is established using a MBS simulation software, SIMPACK[26], as shown in Figure 3. The numerical simulation tool has been widely used and verified for modelling and dynamic analysis of wind turbine drivetrains, as demonstrated in various studies (e.g. Refs. [27, 28, 29, 30]). Topological diagram of the drivetrain model, illustrating the connection relationships between different components and degree of freedoms (DOFs) of each component, is presented in Figure 4. Many studies (e.g. Refs. [31, 32, 33]) have demonstrated the significant effects of flexible body modelling on dynamic response of drivetrain" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000085_imece2010-38641-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000085_imece2010-38641-Figure4-1.png", + "caption": "Fig. 4: Geometrical parameters of the end mill cutter and material removal [2]", + "texts": [ + "7% 5 30 a mm \u03b2 = = F 2.3% 3.2% T 2.4% 2.8% In this section, an improved model of cutting force coefficients in peripheral milling, presented by Yucesan & Altintas [2], is considered. This model is used for 3D simulation of cutting forces, with exponential dependency of the cutting force coefficients to the cut chip thickness. (a) (b) (c) Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/30/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 5 Copyright \u00a9 2010 by ASME Figure 4 shows the cross section of the end mill cutting tool and the geometry of the material removal. A ruled surface which is formed by two generating curves represents the helical flute of the end mill [2, 28]. In this figure, \u03b1 and \u03b8 denote position angle of a point on the cutting edge and tool rotation angle, respectively. Tool rake angle and radial immersion angle are represented by r\u03b1 and en\u03b1 , respectively. Considering the geometry of the problem through defining two generating curves to construct the ruled surface of the helical flutes, and calculating the friction and normal forces, cutting force components are found as follows (more details about the geometric model and formulation of the cutting force can be found in Refs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000209_tac.2009.2020643-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000209_tac.2009.2020643-Figure1-1.png", + "caption": "Fig. 1. .", + "texts": [ + " Case 1: The eigenvalues of are two real numbers with different sign. Without loss of generality, we assume that , . Because is uniformly controllable, we have . Then there is a number such that , , since is bounded. Without loss of generality, we assume that there is a number such that , . For any control curve of (7), we assume that passes through point . Then separates the plane into two disjoint sides denoted as Side L and Side R, respectively (see Section IV). lies in the sector formed by and as shown in Fig. 1, where and are the trajectories of equations and passing through , respectively. Obviously, there is a trajectory of Equation which passes through the control curve from Side R into Side L in Quadrant I. Similarly, there is a trajectory of Equation which passes through the control curve from Side L into Side R in Quadrant III. This implies the criterion function changes its sign on the control curve (see [8]). Because the control curve is arbitrary, the system (7) is globally controllable by Proposition 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003074_iros.2018.8593610-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003074_iros.2018.8593610-Figure2-1.png", + "caption": "Fig. 2. An external external force applied to the CoM of a Nao robot", + "texts": [ + " The force is projected onto the CoM frame, and therefore it can be directly used, for instance, in a pattern generator. The observer design and implementation architecture are given in Section IV. This observer was also validated on a Nao humanoid robot through several experimental scenarios. The results of these experiments are analyzed in Section VI. To study the impact of an external force on the ZeroMoment Point (ZMP) formulation, we consider the two following cases: This case study is illustrated in Fig. 2. The ZMP can be found by considering the forces acting on the simplified LIPM model, which are the gravity, the inertial force and the external force. Let the ZMP on the horizontal ground be given by the following vector1: p = [ px py ]T (1) where px and py are respectively the projections of the ZMP on the x and y axis. To compute p, one can use the following formula: p =N n\u00d7 \u03c4 o (fo|n) (2) where the operators \u00d7 and (.|.) design respectively the cross and scalar products, and 1As a general rule, bold parameters are vectors or matrices \u2022 N is a constant matrix N = [ 1 0 0 0 1 0 ] (3) " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000599_fld.2303-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000599_fld.2303-Figure3-1.png", + "caption": "Figure 3. Conservation of flow rate at a groove\u2013ridge region.", + "texts": [ + " The balance of the boundary terms at the connecting nodes can be simplified by imposing the balance of flow rate across the interface of the element: Qe,e+1 = Qe n+Qe+1 1 = 6 u [ h|e+1 x=x1 \u2212 h|ex=xn ] = { 0 for the node without a step height 6 shu for the node with a step height sh (12) The two-dimensional governing equation for HGJB is solved by the spectral element method, following the procedure used for the one-dimensional instance above. Discontinuity in the groove\u2013 ridge region of a two-dimensional HGJB is modeled with two adjacent elements, as shown in Figure 3. The volume flow rate at point p, located at the interface between element 1 and element 2, can be expressed as: q1x =\u2212 h3 12 p x \u2223\u2223\u2223\u2223 e=1 + uh 2 , q1y =\u2212 h3 12 p y \u2223\u2223\u2223\u2223 e=1 (13) q2x =\u2212 (h+cg)3 12 p x \u2223\u2223\u2223\u2223\u2223 e=2 + u(h+cg) 2 , q2y =\u2212 (h+cg)3 12 p y \u2223\u2223\u2223\u2223\u2223 e=2 (14) Conservation of volume flow rate at point p yields: q1x +q1y =q2x +q2y . (15) Substituting Equations (21) and (22) into Equation (23) yields:[( (h+cg) 3 p x ) + ( (h+cg) 3 p y )] e=2 \u2212 [( h3 p x ) + ( h3 p y )] e=1 =6 cgu (16) Therefore, the difference between the boundary terms of two neighboring elements at the groove\u2013ridge region is given by: Qe,e+1= \u222b 6 cg u \u00b7d Le (17) where the length of two adjacent elements is Le" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002343_icrera.2017.8191184-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002343_icrera.2017.8191184-Figure1-1.png", + "caption": "Fig. 1. Basic wingdingstructure of SRM", + "texts": [], + "surrounding_texts": [ + "Conventionally, the phase wingdings of SRM consist of an even number of series connected windings [12], as shown in Fig. 978-1-5386-2095-3/17/$31.00 \u00a92017 IEEE 1. Thus, central-tapped windings are formed, which can be easily developed in SRM as shown in Fig. 2(a). Fig. 2(a) and Fig. 2(b) show the central-tapped node of the winding and the full-bridge converter. The investigated integrated topology is shown in Fig. 3. A 6/4 poles SRM has phase A, phase B and phase C. La1, La2, Lb1 and Lb2 are winding inductances of phase A and phase B. NA and Nb are winding central-tapped nodes of phase A and phase B. S0~ S11 are twelve MOSFETs, D0~ D11 are twelve parasitic diodes paralleled with MOSFETs. Compared with traditional full-bridge converter, a relay is added between switch S8 and switch S10 to divide the voltage stage of dc-link and battery. The topology has two operating modes: charging mode and driving mode, switched by the relay. Each mode works independently. In driving mode, NA and NB has no connection with one another. In charging mode, the nodes NA and NB are connected to AC power grid. III. SRM DRIVING MODE In driving mode, the relay is turned off. Battery provides power to motor independently. System is totally disconnected from AC grid. The topology is working as traditional full-bridge inverter. Fig. 4 presents three working modes: excitation mode, energy recycling mode and freewheeling conduction mode. When a motor phase is powered by dc-link voltage, the phase is working in the excitation mode. The phase voltage as (1) where Uab is the phase voltage, R is phase resistance, L is phase inductance, i is phase current, is the rotor angular position and is the angular speed [6]. The phase inductance varies as a function of rotor position, and the electromagnetic torque is expressed as (2) where Te is the phase electromagnetic torque. The mechanical motion equation of SRM is given by (3) where J is the combined moment of inertia of the motor and load, B is the combined friction coefficient of the motor and load, is the total electromagnetic torque, and is the load torque [6]. IV. CHARGING MODE In charging mode, relay is turned on. The nodes NA and NB are connected to 50Hz AC power grid. The inverter is separated into a full-bridge converter and cascade buck-boost circuit. The topology is rebuilt as PFC converter and cascade circuit. Diodes D0~ D7 functions as a Rectifying bridge, as shown in Fig. 5. The windings of phase A and phase B perform the role of a filter inductances. A.PFC Buck-Boost converter Due to different levels of battery module and utility voltage applied in different areas of world, adopting single rectifier converter is not enough to meet the variations. Thus, an application with wider range of voltage shall be created and adopted. It is noticed that the PFC is capable of adjusting different voltage ranges, which can be utilized to remedy the limitations of single rectifier converter. Fig. 6 shows the proposed buck-boost converter for single phase utility. The winding of phase C is used as an inductance of the buck-boost converter. The switch S8, diode D10, diode D11 and phase C winding make a buck converter, and the switch S8, switch S9, diode D10 and phase C winding make a boost converter. After rectifier, the DC-link voltage will be different from the voltage of battery, where the buck-boost converter is required. B. Inductor current of Buck-Boost operation mode The operation modes of proposed buck-boost converter are categorized into three modes, as shown in Fig.6 according to switch pattern, and explained as follows: Mode : When Q8 turn on, Q11 turn off. The inductor voltage Vlc is (4) The applied inductor voltage causes the inductor current to rise when Vdc is higher than Vbattety or fall when Vdc is lower than Vbattety, From the AC input voltage and the DC battery voltage, the rising slope and falling slope of the inductor exist. Mode : When Q8 and Q11 turn on. The inductor voltage Vlc is (5) The rectified input voltage is always positive value and varied, i.e., the magnitude of rising slope is not fixed. Mode : When Q8 and Q11 turn off. The inductor voltage Vlc is (6) The battery voltage does not varied in the short term. Thus, the inductor current is fixed negative slope. C.PFC control scheme The closed-loop control is shown in Fig. 7. The control methodology of PFC boost circuit is simple. The controller is of hysteresis type. It consists outer voltage or current controller and inner current controller to regulate the output and shape the current. The output error signal I or U is processed by PI controller and its output after hysteresis controller to determine the switching states of the MOSFETs. In charging mode, the rotor could be moved by torque because the flow of current in the motor phase windings. A mechanical lock can be used in our experiments. And the main control circuit can be fully satisfied on the basis of the original drive main circuit, without additional control circuit, and the actual control is simple and convenient. V. SIMULATION AND EXPERIMENTAL RESULTS Simulation of the charging system is performed using MATLAB-Simulink. In this paper, the voltage control is achieved by a closed-loop control with hysteresis switching control of the converter. Fig. 8(a) provides the simulated waveforms when the converter gets power from 110 V, 50 Hz grid and the output voltage is 300V. Fig. 8(a) presents the waveforms of the input voltage and current, the response of the output current and phase current. It can be observed that the current and voltage have high quality sinusoidal waveforms. The converter demonstrates power factor is 0.999. The voltage ripple in the dc link capacitor is 3V. The corresponding THD for circuit is 2.19%. Fig. 8(b) provides the simulated waveforms when the circuit converters 110 V, 50 Hz grid voltage to 500V on the output. This corresponds to the occasion when the battery is low in state of charge, the operation in buck mode. To verify the feasibility of the integrated charger topology and charging scheme, the implemented system consisted of hardware circuit, software program and a test 6/4 SRM motor under resistive load. Fig. 9 presents the measured input voltage and current waveforms using boost-converter charging mode. It is clearly that the input current is fairly sinusoidal. The charging operation with good power quality by the topology and charging scheme is observed from the result. VI. CONCLUSION This paper describes an on-board integrated battery charger based on central-tapped winding switched reluctance motor for EV application. A additional relay are added when compared with the conventional propulsion system The new topology upgrades the full-bridge inverter circuit and repeated use of motor windings and other drive circuit components to use as charging system with PFC boost-buck function which reduce filtering requirements and achieve high conversion efficiency over the full voltage range of the battery pack. Detailed analysis in operation of the topology is given. Simulation and experimental results validate that the topology is feasible." + ] + }, + { + "image_filename": "designv11_33_0003519_tmag.2019.2925580-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003519_tmag.2019.2925580-Figure5-1.png", + "caption": "Fig. 5. (a) Particle boundary is a chain of counter-clockwisely oriented surfaces. (b) Cube boundary is a chain of counter-clockwisely oriented planes that are such parts of particle boundaries that are lying on the boundary of the cube [0, 1]3 and the dark-gray planes in the picture that are counterclockwisely oriented planes with corresponding holes.", + "texts": [ + " Instead, the geometry is expressed as the formal linear combinations of geometric entities for which algebraic operations are well-defined. Such an approach serves two purposes. First, it allows constructing the gap regions in a straightforward manner using the information about the bounding box and the particles inside. Second, the approach produces a suitable structure for the geometry to be converted into a format required by an external meshing program. First, we construct the representations of the particles and their boundaries. In Fig. 5(a), we can see one particle. Each plane in the boundary of the particle is considered as a counterclockwisely ordered tuple of vertices. One example of such ordered tuple of vertices could be [p4, p3, p3,t , p4,t ], visualized in Fig. 3(b). A particle is a tuple of planes that enclose the volume of the particle. In order to define the boundary of a particle, we have to express many oriented geometric entities at once. To do so, we use chains, formal linear combinations of geometric entities with integer coefficients. We define a boundary operator \u2202 that returns for a particle, a chain of its boundary planes with unit coefficients. Negative coefficients denote the opposite orientations. Next, we construct a representation for the gap region. In Fig. 5(b), we see the boundary of the bounding box that was earlier denoted as [0, 1]3. This boundary is a chain of oriented planes in a similar manner than the boundaries of the particles. Let us denote the boundary of the bounding box as b. If the particles are denoted as p1, . . . , pN , the boundary of the gap regions g, denoted by \u2202g, is given by \u2202g = b \u2212 N\u2211 i=1 \u2202pi (1) where \u2202pi stands for the boundary of the particle pi . We define the gap g as a tuple of its boundary planes. The whole space D is described by a formal linear combination D = g + N\u2211 i=1 pi " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001385_isci.2012.6222678-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001385_isci.2012.6222678-Figure7-1.png", + "caption": "Fig 7: Mounting of Proximity Probes", + "texts": [ + " The rotor shaft of machine fault simulator is supported by fluid film journal bearing. The length of the shaft between two journal bearings is 28.5 inches and the shaft diameter is of 0.5 inches. The lubrication oil used here is ISO 13 mineral oil and the pressure of oil is 12 psi. The displacement sensor used here are proximity probes which is mounted on the bearing housings to measure the displacement of shaft with respect to bearing housing. Displacement sensors are mounted on bearing housing in the horizontal and vertical directions as shown in fig 7. Experiment was conducted with three small disks installed on the shaft of the system that possesses smaller diameter with weight 767 grams. First disk was installed on 5.25 inches, second disk on 10 inches, third disk on 26 inches and centering device was installed on 14.25 inches from left side of journal bearing respectively. For this load condition, disk was installed at the shaft. firstly simulator ran up to over first critical speed and freeze data at that point and then simulator ran up to two times the first critical speed, and stayed at the maximum speed for a while and then coasted down" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002490_tcpmt.2017.2781723-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002490_tcpmt.2017.2781723-Figure3-1.png", + "caption": "Fig. 3. Target dimensions of towers and bridges.", + "texts": [ + " The measured materials exhibited lower permittivity than expected from our mixing calculation estimates. We found the ferrite powders disadvantageous due to their irregular shapes and high viscosity values after loading more than 24%. The ferrite powder with spherical particles had a better flow, but the size of the particles made it challenging to print fine features. The ferrite powder with smaller particle size had sharp edges that cause clogging on the microdispensing pen tips. C. 3-D Printed Dielectric Structures The target dimensions for a simple tower and bridge structure are depicted in Fig. 3. These dimensions were easily realized with the dielectric material because the small spherical particles made it ideal for dispensing through a small pen tip. Process parameters were determined empirically for each material independently. The goal of the process parameters was to dispense the material with as close to zero momentum as possible. There is a relation between viscosity, velocity of the pen tip, and dispense rate that affects the build [5]. If the structures are 3-D printed with too high of dispense rate relative to the velocity of the pen tip, the dispensed material can push the structures over during the build" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001692_iceceng.2011.6057034-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001692_iceceng.2011.6057034-Figure1-1.png", + "caption": "Figure 1. Architecture of the humanoid manipulator.", + "texts": [ + " In this paper, the inverse kinematics of the manipulator is analyzed, and a joint space planning method under multiple constraints is proposed to increase the success rate of playing by optimizing the joint angular acceleration. II. KINEMATICS ANALYSIS All of the commands, such as the position and velocity of the racket at the hitting moment, are given in the operation space by the visual system. In order to generate the joint trajectories, they should be transformed into joint space values with inverse kinematics. 3047 978-1-4244-8165-1/11/$26.00 \u00a92011 IEEE Fig. 1 shows the mechanical architecture of the 7-DOF humanoid manipulator. It has seven rotational joints which are driven by DC serve motors and speeded down by harmonic gear reducers. A racket is mounted on the end of the manipulator. Posture of each joint can be described by a dynamic coordinate xiyizi (i=1\u00b7\u00b7\u00b77). Joint 2, joint 4 and joint 6 rotate around their own x-axis while joint 1, joint 7 rotate around the y-axis and joint 3, joint 5 rotate around the z-axis. The rotational axes of joint 1, joint 2 and joint 3 intersect at the shoulder while the rotational axes of joint 5, joint 6 and joint 7 intersect at the wrist" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001157_iciea.2010.5516812-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001157_iciea.2010.5516812-Figure1-1.png", + "caption": "Figure 1. The laboratory set-up TRMS", + "texts": [ + " SYSTEM DESCRIPTION The Two Rotor MIMO System (TRMS) is a laboratory setup 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 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 block diagram of the TRMS model is shown in Fig. 2. It is suitable for using in the SIMULINK environment. The model of the DC motor with propeller is composed of a linear dynamic system followed by a static nonlinearity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003896_j.triboint.2019.106096-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003896_j.triboint.2019.106096-Figure5-1.png", + "caption": "Fig. 5. Exemplary representation of the ring discretization and the evaluation routine.", + "texts": [ + " Tribology International 143 (2020) 106096 energy density (Eq. (8)) characterizes the energy input for every area element. \ufffd E A \ufffd max \u00bcmax \ufffd 0:5 \u22c5 Z t\u00fet2b t \u03bc\u00f0t\u00de \u22c5 p\u00f0x; y; \u03c4\u00de \u22c5 \u00f0u1\u00f0\u03c4\u00de u2\u00f0\u03c4\u00de\u00ded\u03c4 \ufffd (8) Therefore, the contact time t2b is calculated in dependence of: (a) the contact width in rolling direction (y-axis) and (b) the circumferential speed. In the implemented evaluation routine the frictional power density is applied, therefore, with a moving coordinate system over the discretized ring surface. This approach is exemplary shown in Fig. 5. The result is consequently a spatially-resolved frictional energy for a ring rotation, in this case for the follower. The calculations are accordingly carried out in the time window between the first motion of the follower during the acceleration process and its first rotation. The assessment of the contact conditions is then based on the maximum value of the resulting value matrix. This methodology is exemplary shown in Fig. 6. While Fig. 6-top shows measured values, Fig. 6-bottom show the values determined during postprocessing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003253_012016-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003253_012016-Figure7-1.png", + "caption": "Figure 7. Design of the pressure sensor test specimen (a), housings: as built (b), threaded (c), polished (d), result of the force-distance monitoring: applied force and corresponding voltage in dependence of the time for a membrane thickness of 2.9 mm [8] (e).", + "texts": [ + " Number and size of the welding spots were varied (Figure 6 c, Table 1). The distance between two adjacent welding spots is 1.5 mm. In the samples with smaller welding spot width (samples 1 and 3) the strain gauge is completely connected to the base body. After welding, the additive manufacturing process is continued and the cover layer prepared. The cross-section of sample 3 demonstrates the successful integration of the strain gauge (Figure 6d). The integration of the capacitive differential pressure sensor using an active fluid medium is shown in Figure 7. To verify the integration concept housings were built by LBM. The membrane thickness of the as built housings varies from 3.7 mm to 4.1 mm. After separation from the build platform the defined membrane thickness of 2.7 to 3.2 mm is reached (Figure 7b). Each specimen is then threaded with a 6.35 mm (1/4 inch) bore diameter (Figure 7c). The flange surface is polished (Figure 7d). Then WTK IOP Conf. Series: Materials Science and Engineering 480 (2019) 012016 IOP Publishing doi:10.1088/1757-899X/480/1/012016 the active medium is filled into the cavity of the housing. Finally, the pressure sensor is screwed in. The pressure acting on the surface of the membrane of the housing is transformed by the medium into hydraulic pressure and measured by the pressure sensor. For verification of the measuring principle the manufactured test specimen are characterized in a universal compounding machine (PROMESS) with force-distance monitoring. The force was varied from 10 kN to 40 kN for each sample and the voltage of the pressure sensor recorded. A nearly linear relationship between the applied force and the measured voltage was verified (Figure 7e). Automotive parts made from high-strength steels formed by press hardening offer enormous potential for lightweight design. By use of this technology, the sheet metal is heated above 900 \u00b0C, the austenitizing temperature, and subsequently quenched during the forming operation to below 200 \u00b0C, forming a martensitic microstructure. A tool for forming of gear pans from coated heat-treatable boron steel (22MnB5) was designed and an innovative cooling system developed using thermal und forming simulations [3]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003253_012016-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003253_012016-Figure1-1.png", + "caption": "Figure 1. Scheme of the laser beam melting (LBM) process.", + "texts": [ + " A concept for integration of an encapsulated actuator system was developed and is presented. Manufacturing concept and materials Laser beam melting is a very versatile additive manufacturing technology and in this paper is used as base process. The starting point is a 3D CAD model, which is virtually sliced into thin layers with a layer thickness of about 45 \u00b5m. Based on this data the physical part is built. A thin layer of metal powder, stored in a dose chamber, is uniformly distributed across the working area by a coater blade (Figure 1). The light emitted from an ytterbium fiber laser system at the wavelength of 1070 nm is directed by a galvanometer scanner across the powder layer according to the cross section of the part. The machine used, a M2 LaserCUSING\u00ae from Concept Laser, applies an island exposure strategy, i.e. the segments of each layer are irradiated in stochastical succession. Upon irradiation the powder melts to form a melt pool, covering not only the top powder layer, but also the already solidified layer below. After laser exposure of the first powder layer, the build platform is lowered, the next powder layer is applied and the exposure process starts again" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000945_ssd.2010.5585533-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000945_ssd.2010.5585533-Figure9-1.png", + "caption": "Figure 9. Mesh of the stator yoke and of the two magnetic collectors.", + "texts": [ + " The CPAES static and rotating components are illustrated in figure 6. One can notice that each claw plate includes six poles. Moreover it is to be noted that the two claw plates are magnetically decoupled. In order to reduce the computation time, the FEA study domain is limited to a one pair of poles of the CPAES. Figures 7 and 8 show the stator and the rotor study domains, respectively. The meshing of both stator and rotor study domains has been automatically generated in the Ansys-Workbench environment. Figure 9 shows a mesh of the stator yoke and of the two associated magnetic collectors. A mesh of the stator lamination is illustrated in figure 10. Figure 11 shows a mesh of the rotor claws and the as sociated magnetic rings. 3.2.1. Main Flux Paths The flux moves through the magnetic circuit of the CPAES within three dimensional (3D) paths. These are described as follows: \u2022 axially in the stator yoke (see figure 12), \u2022 radially down through the collector and crossing the air gap down to the magnetic ring (see figure 12), \u2022 radially then axially in the magnetic ring (see figure 12), \u2022 axially-radially in the claws (see figure 13), \u2022 radially up in air gap facing the stator laminations (see figure 14), \u2022 radially up in the stator teeth and circumferentially in the stator core back (see figure 15), \u2022 radially down in stator teeth facing the two adjacent claws (see figure 16), \u2022 radially-axially in the adjacent claws (see figure 16), \u2022 axially then radially in the magnetic ring (see figure 17), \u2022 radially up through the air gap and the mag netic collector on the other side of the machine (see figure 17), \u2022 axially in the other side of the stator yoke (see figure 18)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001520_pes.2011.6039209-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001520_pes.2011.6039209-Figure2-1.png", + "caption": "Fig. 2. Schematic diagram of the two-mass equivalent model", + "texts": [ + "00 \u00a92011 IEEE The differential equations of the three-mass model are: ( ) ( )\u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23a9 \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23a8 \u23a7 \u2212= \u2212= \u2212\u2212= \u2212\u2212= \u2212\u2212= 12 12 2 2 1112 1 1 22122 2 2 2 2 2 2 tts t tgs gt ttmt t t tttgt t t gggtg g g dt d dt d DTT dt dH DTT dt dH DTT dt d H \u03c9\u03c9\u03c9\u03b8 \u03c9\u03c9\u03c9 \u03b8 \u03c9\u03c9 \u03c9\u03c9 \u03c9 \u03c9 (1) ( ) ( )\u23a9 \u23a8 \u23a7 \u2212+= \u2212+= 1212121212 22222 tttttt ggtgtgtgt DKT DKT \u03c9\u03c9\u03b8 \u03c9\u03c9\u03b8 (2) where 1tH , 2tH and gH are the inertia constants of the wind turbine blades, the wheel hub and the generator rotor, respectively; 1tD , 2tD and gD are the damping coefficients of the wind turbine blades, the wheel hub and the generator rotor, respectively; 1t\u03c9 , 2t\u03c9 and g\u03c9 are the angle speeds of the wind turbine blades, the wheel hub and the generator rotor, respectively, and fs \u03c0\u03c9 2= are the synchronous angle speed; 12tD , 12tK , 12t\u03b8 and 12tT are the damping coefficient, the shaft stiffness coefficient, the shaft twist angle and the shaft torque between the blades and the wheel hub, respectively; gtD 2 , gtK 2 , gt 2\u03b8 and gtT 2 are the damping coefficient, the shaft stiffness coefficient, the shaft twist angle and the shaft torque between the wheel hub and the generator rotor, respectively; mT is the mechanical torque of the wind turbine; gT is the electromagnetic torque of induction generator. Figure 2 shows the two-mass model of the drive train in which the bending flexibility of blades has been neglected [5]- [12]. The differential equations of the two-mass model are: ( )\u23aa \u23aa \u23aa \u23a9 \u23aa \u23aa \u23aa \u23a8 \u23a7 \u2212= \u2212\u2212= \u2212\u2212= tgs tg ttmtg t t ggtgg g g dt d DTT dt dH DTT dt d H \u03c9\u03c9\u03c9 \u03b8 \u03c9\u03c9 \u03c9 \u03c9 2 2 (3) ( )tgtgtgtgtg DKT \u03c9\u03c9\u03b8 \u2212+= (4) where tH and gH are the inertia constants of the wind turbine the generator rotor, respectively; tD and gD are the damping coefficients of the wind turbine the generator rotor, respectively; t\u03c9 and g\u03c9 are the angle speeds of the wind turbine the generator rotor, respectively, and 21 ttt HHH += ; tgD , tgK , tg\u03b8 and tgT are the shaft damping coefficient, the stiffness coefficient, the twist angle and torque, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001346_j.proeng.2012.04.091-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001346_j.proeng.2012.04.091-Figure2-1.png", + "caption": "Fig. 2. The middle part of a riser with the clicker in the process (A-C) of a shot. The clicker is a 5 cm long and 0.5 cm wide piece of spring steel that is fixed to the riser. In the process of a shot first the archer draws the bow holding the arrow between the bow and the clicker, so that the clicker presses the arrow lateral against the riser. During aiming the archer holds the stretched bow while the clicker presses the arrow lateral against the bow (A). At the end of the aiming phase the archer does the final pull: he/she pulls the arrow back until the clicker slips over the arrowhead and causes a click (B). Then the archer shoots by releasing his/her hand from the string and the arrow is accelerated (C) [2]", + "texts": [ + " A \u201cleave-one-out\u201d cross validation procedure revealed consistent estimate of the model (all corrected R\u00b2 were in a range of 0.66 and 0.75 with p<0.03). It has been shown that a highly precise timing in arrow release in terms of small CVs of clicker time is important for high mean scoring. \u00a9 2012 Published by Elsevier Ltd. Keywords: Recurve archery; arrow release; measurement system; clicker time; motor program * Corresponding author. Tel.: +43-1-4277-48886; fax: +43-1-4277-48889. E-mail address: mario.heller@univie.ac.at. The process of shooting with a recurve bow (see Fig. 1) can be described as follows (see Fig. 2): The archer draws the bow, pulls the arrow to the clicker, fixes in this position and aims. Then he/she pulls the arrow through the clicker (the so called \u201cfinal pull\u201d) and releases the shot. From a biomechanical point of view, the archer has to cope with the breakdown of the static balance of forces between the external tension and his/her muscular forces at the time of shooting [1]. The final pull and the release of the shot have been of interest in some earlier studies. In most cases, electromechanical devices were directly attached to the bow to detect the moments of clicker closure, arrow release, and contact-loss of the arrow with the bowstring with high temporal resolution [3-8], which is very time-consuming and complex to affix. Another approach was to fix an acceleration sensor at the bow riser detecting mechanical vibrations [1-2, 9]. Most authors report that the light sound of the click (as described in Fig. 2) is the stimulus for the archer to extend his or her pull fingers, which induces the release of the bowstring [3-8], and, moreover, having a quick reaction to the clicker\u2019s fall (sound) is intended to be directly related with the performance of the archer [10]. However, this is questioned by Edelmann-Nusser and Gollhofer [9] who investigated highly skilled archers from the German National Team under several experimental conditions (normal vs. extended shots, reaction tests, and lengthened arrows)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001080_elan.201100249-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001080_elan.201100249-Figure1-1.png", + "caption": "Fig. 1. Schematic expanded drawing of modular stack cell, each compartment is 1 cm in height and 1 cm in inner diameter.", + "texts": [ + " Buffered cathode solution: 50 mM PIPES buffer, 100 mM NaCl; pH 7.0 [39]. Shewanella was used as the bacteria as it has been shown to be a reasonable test bacterium to evaluate MFCs [50\u2013 53]. Shewanella oneidensis MR-1 (MR-1) was grown in sterile Luria\u2013Bertani (LB) broth (15 mL) inoculated with stock cultures retrieved from 80 8C storage. LB cultures were incubated aerobically for 20 hours at 30 8C and at a rotation rate of 140 rpm. This broth was then used directly to inoculate the anode as described previously [48]. The two chamber stack cell MFC (Figure 1) was used as described previously with the exception that an influent and effluent port were incorporated into module 1 and module 2 to permit pumped delivery of the substrate to and through the chitosan-carbon nanotube scaffold [48]. Inoculation of the chitosan-carbon nanotube scaffold anode was accomplished by directly filling the anode compartment with fresh inoculum and allowing the biofilm to grow [48]. The cathode employed is this study is a platinum sputtered carbon felt disk (d=1 cm, h=6 mm)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003893_0954406219888240-Figure20-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003893_0954406219888240-Figure20-1.png", + "caption": "Figure 20. Parameters of the spherical joint.", + "texts": [ + " Due to the complexity of the accurate mathematical models of the joint clearance and the contact deformation, the stiffness magnitude is the main concern in this paper, and the given joint clearance and contact deformation models are approximate. Since the joint clearances and joint contact deformations are closely related to the joint structure, the approximate clearance model and the contact deformation model of the ideal and combined spherical joints will be respectively discussed as follows. Approximate clearance model of the ideal spherical joint As shown in Figure 20, if the gap in the spherical joint is denoted by c1, the joint clearance will equal to the value of this gap and along the direction of the contact force due to the property of pure rotation motions. Therefore, the maximum clearance model of the ideal spherical joint can be simply defined as Pcle S \u00bc \u00bdc1, c1, c1; 0, 0, 0 T \u00f019\u00de where the first three components denote the translations, the remaining three components denote the rotations, and this clearance model is also applicable in the following discussions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001986_iet-epa.2016.0565-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001986_iet-epa.2016.0565-Figure2-1.png", + "caption": "Fig. 2 Structure diagram of dual skew rotor", + "texts": [ + " Main rotor slot harmonics result from the interactions of fundamental MMF with rotor slot permeance harmonics, and interactions of fundamental permeance with rotor MMF harmonics [4]. The highorder rotor slot harmonics arise from the interactions of the high harmonics of rotor MMF with the high rotor slot permeance harmonics. The magnitude of high-order rotor slot harmonic magnetic flux density is smaller than the main rotor slot harmonic magnetic flux density. The high-order rotor slot harmonics are ignored due to its small magnitude. The structure diagram of dual skew rotor is shown in Fig. 2. The dual skew rotor was composed of two single skew rotors. The two parts are skewed in inverse direction, staggered in the circumferential direction and connected by intermediate ring [12]. In the dual skew rotor, the bars are skewed inversely. So, the distributions of rotor slot harmonics will be divided into two parts. Axial length of intermediate ring is short and ignored for simplifying analysis. The harmonic MMF of dual skew rotor winding can be written as the following equations: f \u03bc1 \u03b8, t = F\u03bc cos \u03bc\u03b8 \u2212 \u03c9\u03bct \u2212 \u03c6\u03bc \u2212 \u03bcbsk1 R Z l1 (1) while in Z>0 region f \u03bc2 \u03b8, t = F\u03bc cos \u03bc\u03b8 \u2212 \u03c9\u03bct \u2212 \u03c6\u03bc \u2212 \u03bcbsk2 R Z l2 \u2212 \u03bc\u03b1 (2) while in Z<0 region \u03c9\u03bc = \u03c91 1 \u00b1 k2Z2 p (1 \u2212 s) (3) The method which approximately MMF multiply permeance was adopted to calculate air-gap permeance slot harmonic flux distribution of dual skew rotor winding" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002756_s12555-017-0569-1-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002756_s12555-017-0569-1-Figure4-1.png", + "caption": "Fig. 4. One-link manipulator with a flexible joint.", + "texts": [ + " We can see that state trajectories converge to the consensus function. Fig. 3 depicts the trajectories of time-varying edge weights, one can see that all weights converge to finite values. From the simulation results, we can see that this multi-agent network achieves adaptive consensus. Example 2: practical case In this case, a one-link manipulator with revolute joints actuated by a DC motor, also as shown in [11] and [14], is presented to prove the applicability of the proposed method. The schematic diagram of this one-link manipulator is shown in Fig. 4. It is assumed that the multi-agent network is consisted of six identical one-link manipulators. The dynamics of each one-link manipulator can be modeled by (1) with A = 0 1 0 0 \u221248.6 \u22121.25 48.6 0 0 0 0 1 19.5 0 \u221219.5 0 , B = 0 21.6 0 0 , g(xi) = 0 0 0 \u2212\u03b2 sin(xi3) , where xi = [xi1,xi2,xi3,xi4] T with i = 1, 2, ..., 6, and \u03b2 = 0.3. The communication topology is the same as the one in Example 1 and the initial states are given as follows: x1(0) = [0.6,\u22121.0,\u22121.6,3.0]T , x2(0) = [0.5,2.5,\u22122.6,1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000341_jnn.2009.se10-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000341_jnn.2009.se10-Figure2-1.png", + "caption": "Fig. 2. Diagram of the heated CPE. (A) scenograph (B) side elevation (a and b) connect to the function generator; (c) connect to the potentiostat; (d) working area of the electrode; (e) PMMA clump; (f) carbon paste.", + "texts": [ + "23 In all heating experiments the frequency was adjusted to 100 kHz, because the high frequencies could prevent the voltage of alternating current from interfering with the reactive process on carbon paste electrode (CPE). For electrically heating of an electrode inside an electrolyte solution, alternating current has been used instead of direct current, because of avoiding polarization.28 Changing the output of the function generator can change the temperature of the working electrode rapidly and easily. The diagram of the heated CPE is shown as Figure 2, one hole (length 20 mm, diameter 2 mm) was drilled horizontally in a polymethyl methacrylate (PMMA) clump (20 mm\u00d710 mm\u00d75 mm), and another upright hole (diameter 2 mm) was drilled in the middle of the horizontal hole. The carbon paste was prepared by mixing raphite powder and mineral oil (60/40 %w/w). Both the holes were filled with carbon paste and three electrical connections (see a, b, c in Fig. 2) were glued into the holes with epoxy resin, of which a and b connect to the function generator for heating and leave c to connect with the potentiostat. A segment (2 mm\u00d710 mm\u00d72 mm) was resected from the middle of PMMA clump to form a groove, a disk of carbon paste (about 2 mm in diameter) was then unfolded, which was the working face of heated CPE (see d in Fig. 2). 2.4. Immobilization of Ru(bpy)2+3 in MWNT/Nafion-Modified CPE 1 mg of MWNT was dispersed in 10 mL of 0.5 wt% Nafion aqueous alcoholic solution, and the composite solution was radiated by ultrasonic for 30 min for welldistributed. Before modification, the working area of CPE was carefully polished with weighing paper and rinsed with pure water. The MWNT/Nafion composite film was produced by dropping 10 L of the composite solution onto the working area of CPE and evaporated the solvent at the room temperature in the air" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002377_978-3-319-68826-8_7-Figure7.70-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002377_978-3-319-68826-8_7-Figure7.70-1.png", + "caption": "Fig. 7.70 Scheme of WMR Pioneer 2-DX in the measurement environment", + "texts": [ + " Distance to obstacles is measured by the sensory system of WMR Pioneer 2- DX and consists in cyclical activation of measurements carried out by individual ultrasonic sensors su1, . . . , su8 in a pre-defined sequence. The maximal measured distance to an obstacle is dmax = 6 [m]. Distance to obstacles located in front of the WMR was measured by the laser rangefinder irrespectively of measurements conducted with the use of ultrasonic sensors. A schematic of WMR Pioneer 2-DX in the measurement environment is shown in Fig. 7.70, where the following notations were used: l, l1 \u2013 dimensions resulting from WMR\u2019s geometry, \u03b2 \u2013 frame rotation angle, A (xA, yA) \u2013 point located at the intersection of the WMR frame symmetry plane and the axis of driving wheels, G (xG, yG) \u2013 selected goal of theWMR\u2019smotion, x1, y1 \u2013 axes of the movable system of coordinates connected with point A in such a way so that axis x1 is identical to the symmetry axis of the WMR frame, pG \u2013 straight line passing through the points A and G, \u03c8G \u2013 angle between the line pG and axis x1, \u03d5G \u2013 angle between the line pG and axis x of the immovable system of coordinates, dLi, dRi \u2013 measurements of distance to obstacles, executed by the sensory system of the WMR, selected from all measurements and assigned to distance measurement groups for the left and right side of the WMR frame, i = 1, 2, dF \u2013 measurement of distance to an obstacle in front of the WMR frame, executed by the laser rangefinder, dF = dLr , \u03c9Li, \u03c9Ri \u2013 angles between axes of individual measurements and axis x1 of the movable system of coordinates" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000625_1.4006456-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000625_1.4006456-Figure2-1.png", + "caption": "Fig. 2 System at the impact moment; collision with inclined wall", + "texts": [ + " 4 Ground Impact Map To solve the baton multiple impacts problem, we use the impulse-momentum method along with the energetic coefficient of restitution [16] to determine the post-impact velocities. The continuous dynamics of the hybrid system in Eq. (7) provides the continuous state vector x = x{t) for x{t) 0D and t e [tk-\\,tk), k eN. When the state vector x{t) reaches the hyperplane D at time i\u0302 the impact event would take place. The preimpact and post-impact states are denoted by x'^, and x^, respectively. Figure 2 demonstrates the components of velocities and impulses during the impact. 4.1 Impulse-Momentum Equations. When the collision of mass mi with the inclined wall takes place, the tangential and normal impulses at the inclined wall are denoted as ff' and F\\\\ respectively. These impulses can be expressed in the spatial reference frame ({X, Y, Z}) with the following components: = F{' cos7 \u2014 ' sin 7 = Fy'' sin y + F\u00a1 ' cos y (8) (9) =Px -LRQX (5) The linear momentum equations of the system (written along the surface and normal to surface) from the preimpact state to the post-impact state are written as" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000162_iemdc.2009.5075177-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000162_iemdc.2009.5075177-Figure3-1.png", + "caption": "Fig. 3. Schematic of a 4-pole IPM wind generator with straight type NdBFe permanent magnets.", + "texts": [ + "44 a s m d p E N f k k\u03c6= (6) where Na is the number of turns per phase in the stator winding, fs is frequency of the generated voltage in the stator, m \u03c6 is the peak permanent magnet (PM) flux per pole, kd is the stator winding distribution factor, and kp is the stator winding pitch factor. The generated frequency \u03c9s is directly related to the shaft speed of the wind turbine by the following expression 2 2 s m s P f\u03c9 \u03c9 \u03c0= = (7) where P is the number of rotor PM poles of the IPM machine. The IPM wind generator is a salient pole ac machine in which the saliency is created by the orientation of magnets arrangement inside the rotor assembly in stead of built-in variable air gaps of a typical salient pole synchronous machine. Figure 3 shows the schematic of a 4-pole IPM wind generator. The NdBFe magnetic materials are arranged in straight magnet form below the squire cage type dampers. Figure 4 shows the magnet flux distribution of an IPM machine using a straight type arrangement of magnets inside the 4-pole rotor. Figure 5 shows the per phase voltage equivalent circuit diagram of an IPM generator. The steady state per phase voltage E0 at the air gap of a P-pole three phase balanced IPM generator can be expressed as ( )0 0 ( )" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000945_ssd.2010.5585533-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000945_ssd.2010.5585533-Figure3-1.png", + "caption": "Figure 3. Field winding inserted in both sides of the ma chine between the armature end-windings and the housing.", + "texts": [ + " As a result, the brush-ring system and the associated mainte nance problem have been discarded, which represents cru cial cost and availability benefits. The machine is equipped by a three phase armature winding. In the manner of conventional claw pole alterna tors, the stator is made up of a laminated cylindrical mag netic circuit as shown in figure 1. The field winding is simply wound in a ring shape. Figure 2 shows the photo of one half of the stator field winding. The two halves of the field winding are inserted in both sides of the machine between the armature end-windings and the housing as illustrated in figure 3. They are con nected in series in such a way to produce additive fluxes. Following the transfer of the DC-excitation winding from rotor to stator, appropriate changes of the magnetic circuit have been introduced. These concerned mainly the rotor where the two iron plates with overlapped claws facing the air gap tum to be magnetically decoupled. Figure 4 shows a photo of the rotor of the CPAES. Following the removal of the field winding from rotor to stator and for the sake of an efficient flow of the flux, two magnetic collectors have been included in the stator mag netic circuit" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001073_ecce.2012.6342517-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001073_ecce.2012.6342517-Figure1-1.png", + "caption": "Fig. 1. Stator circuit of induction machine with stator inter-turn fault and a specific distributed stator winding with possible fault locations", + "texts": [ + " The stator inter-turn fault has been modeled in various ways based on the simplicity and accuracy measures in literature. Simplified q \u2212 d models of induction machine with stator inter-turn fault are developed without considering fault location in [4], [5], and with considering fault location in [6]. The stator inter-turn fault is modeled in [4], [5] considering only the number of faulted turns as a fault parameter. The schematic of stator circuit of induction machine with interturn fault is shown in Fig. 1. However, fault location affects the flux distribution and the machine inductances. The model with consideration of fault location is proposed in [6]. The fault location parameter reflects in the machine inductance equations of the model. A specific distributed stator winding with some possible fault locations, used in [6] to include the fault location effects in the model, is shown in Fig. 1. The rotating frame transformation normally used for linearizing q\u2212 d machine models results in time variant models of stator inter-turn faulted machine. These models are time periodic in the steady-state. Hence, the eigenvalue analysis tool for analytical studies of transients is not directly applicable to q \u2212 d models of faulted machines. One approach to the analysis of time-periodic systems is the \u2018Floquet\u2019 theory. This approach is explained in [7], and is applied to a rectifier-generator system", + " These analytical studies and results of experiments and simulations show that the positive and negative sequence current transients contain little information of the stator inter-turn fault. While this is as expected, this paper provides an analytical framework to quantify this conclusion. This paper also provides analytical validation for coupling in transients of positive and negative sequence currents due to presence of fault in the machine. The aim of this paper is the development, validation and study of the dynamic phasor small-signal model. While this model could also be used to develop diagnosis methods, this issue has not been addressed in this paper. Fig. 1 shows the schematic of stator winding with stator inter-turn fault. To derive the simplified dynamical model, the healthy stator and rotor windings are assumed to be balanced and sinusoidally distributed. The saturation of the magnetic materials of the machine is neglected. Air-gap is assumed to be uniform. Space harmonics are neglected. The fault loop resistance is proportional to the number of faulted turns. The leakage inductance of the fault loop is proportional to the square of the number of faulted turns", + " The eigenvalue study has been done on different rating machines, covering a wide range of machine ratings. The parameters of these machines are given in Table I. These parameters are also given in [12]. Table II shows the eigenvalues of healthy machine for operating points at stall and no-load speed. All eigenvalues of the healthy machines are complex conjugate pairs. Similarly, Table III shows the listing of eigenvalues for induction machine with stator inter-turn fault at stall and no-load speed. In this study, the fault location is considered at coil 2\u22122\u2032 (Fault2 in Fig. 1) and fault severity is the short-circuit of entire coil. The complex conjugate pair \u03bb9,10 is due to the fault, and does not exist in the healthy machine. This complex conjugate pair is far away on real axis from other eigenvalues. The participation factor analysis of faulted machine eigenvalues shows that the eigenvalues \u03bb9,10 are mainly dominant on the fault current transients and eigenvalues \u03bb1,..,8 are mainly related to the stator and rotor transients. The eigenvalues with less change in their imaginary part for stall and no-load conditions of the machine are associated with stator transients", + " This indicates the decoupling of positive and negative sequence current transients in healthy machines. The participations of eigenvalues \u03bb9,10 are small in the states of stator and rotor. This indicates less effect of fault on stator and rotor states. Fig. 3 shows the experimental set-up. The induction motor used for the experiments has several taps at different locations on the stator winding. Using these taps, it is possible to create stator turn faults at different locations. These taps are shown in Fig. 3. Fault-1 shown in Fig. 1 can be created by connecting taps 1 and 2, which involves entire coil 1 \u2212 1\u2032 in the fault. Similarly, Fault-2 and -3 can be created by connecting taps 2 and 3, and 3 and 4 respectively. These faults have the same severity but they are at different locations on the phase-A winding. The parameters of the machine used in the experiments are given in Table IV. These are obtained by laboratory tests. The load on the induction motor can be changed by loading the dc generator coupled to the induction motor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001385_isci.2012.6222678-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001385_isci.2012.6222678-Figure2-1.png", + "caption": "Fig. 2: Shaft off centre in journal bearing [19]", + "texts": [ + " This is also reported that different types of flexible coupling show different frequency composition in the vibration response of same misalignment condition [27]. It shows that mystery still surrounds the exact response of vibration for the misalignment. Oil whip can occur due to the design of journal bearing, misalignment of the shaft or due to sudden shock or impact to the machine [14]. If the shaft which is used on machine is moved from center position due to load, misalignment, or eccentricity, then the clearance on one side of bearing will be greater than that on the other side and the clearance on the other side of bearing will be lesser. Fig 2 shows the shaft off center position in journal bearing. Here, Shaft alignment is the positioning of the rotational centers of two or more shafts such that they are co-linear when the machines are under normal operating conditions. There are two components of misalignment which is angular and offset. Offset misalignment, is sometimes referred as parallel misalignment. This is typically measured at the coupling center. Angular misalignment is sometimes referred as gap or face, is the difference in the slope of one shaft, and usually occurs in the moveable machine, as compared to the slope of the shaft of the other machine, usually in the stationary machine" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000654_ectc.2011.5898636-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000654_ectc.2011.5898636-Figure2-1.png", + "caption": "Figure 2. Sketch of the CPW line with meander load, where L, W, and S are the width of the signal line, the length of the signal line, and the spacing between the signal line and the grounds. The width of the meander lines is 3m, and the spacing between each meander line is 10m.", + "texts": [ + "00 \u00a92011 IEEE 1030 2011 Electronic Components and Technology Conference analyzer (PNA-N5230A), up to 26GHz (limited by the system). Changes of impedance of the CPWs due to adsorption of target molecules can be noted in the scattering parameters (S-parameters). Thus the presence of adsorption of molecules and surface interactions will be monitored. nanosheets about 1 nm thick in average; the concentrated samples showed overlapping of many layers of graphene, ensuring a complete surface coverage. A number of coplanar wave guides (CPWs) sensing platforms with various configurations, as shown in figure 2, were designed by using Agilent advanced design system (ADS) and were fabricated on a glass substrate (AF-45) to eliminate the substrate loss. Ti/Au was deposited to fabricate the CPW. The 200nm thick Ti was used to improve the adhesion between the Au layer and the glass substrate. Au was selected as the conductor because of its chemical inertness. Standard lift-off processing was employed to pattern the gold film into CPWs. The purpose of using glass substrate is to suppress the substrate loss at high frequencies", + " l<<1, the load impedance can be determined directly by the S-parameter 11 11 1 1 S S ZL (4) Figure 3 shows the measured load impedances as a function of frequencies. The three devices A1, B1, and C1 resonate at ~ 6.0 GHz, 10.5 GHz, and 20.8 GHz, respectively, manifested by reaching to the maximum of the resistance, and approaching to zero of the reactance. Since the lengths and the characteristic impedance of the devices are the same, the red-shift of these resonant frequencies thus cannot be caused by the CPWs. On the other hand, the load consists of many turns of meander lines, as shown in Figure 2. At high frequencies, these meander-shaped lines show significant value of inductance and capacitance, forming the LC resonances. More turns of the meander lines, the larger inductance and capacitance, lead to lower resonant frequency. II. Extraction of Substrate Effective Permittivity The results from adding chemicals on the CPW devices were compared and presented in Figures 4 and 5. Significant red shifts of LC resonant frequencies (Fig. 4) and decrease of resistance (Fig. 5) at these resonant frequencies were observed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000777_s0025654411040017-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000777_s0025654411040017-Figure6-1.png", + "caption": "Fig. 6. Fig. 7.", + "texts": [], + "surrounding_texts": [ + "By setting \u03b4 = 0 in Eqs. (3.6) and (3.7) and by introducing new variables z1 and z2 by formulas z1 = z + a1\u03d5 and z2 = z \u2212 a2\u03d5, we obtain the equations z\u2022\u20221 + a11(c1z1 + \u03b31z \u2022 1) + a12(c2z2 + \u03b32z \u2022 2) = 0, z\u2022\u20222 + a21(c1z1 + \u03b31z \u2022 1) + a22(c2z2 + \u03b32z \u2022 2) = 0, (4.1) MECHANICS OF SOLIDS Vol. 46 No. 4 2011 a11 = 1 + a2 1 \u2212 a1f1, a12 = 1 \u2212 a1a2 \u2212 a1f2, a21 = 1 \u2212 a1a2 + a2f1, a22 = 1 + a2 2 + a2f2. (4.2) The dimensionless variables ak and fk (k = 1, 2) in formulas (4.2) are the ratios of the corresponding variables introduced in the preceding sections to the radius of inertia \u03c1; i.e., we assume that the radius of inertia \u03c1 is equal to unity. The variables z1 and z2 have the following physical meaning. These are small deviations of the points O1 and O2 from their equilibrium positions described in Sec. 3. Equations (4.1) and (4.2) and relations (1.1) with the use of the Lyapunov theorem permit solving the above stability and instability problem for any fixed v. Let us write out the characteristic polynomial of the linear system (4.1) as p4 + p3(a11\u03b31 + a22\u03b32) + p2(a11c1 + a22c2 + \u03b31\u03b32\u0394) + p(\u03b31c2 + \u03b32c1)\u0394 + c1c2\u0394 = 0, (4.3) where \u0394 = a11a22 \u2212 a12a21. Applying the Routh\u2013Hurwitz theorem to the polynomial (4.3), we arrive at the following result. Proposition 1. For the asymptotic stability of the equilibrium in the case of rectilinear and uniform motion of the vehicle, it is necessary and sufficient that the rolling friction coefficients f1 and f2 (and hence the velocity v of the vehicle translational motion determined by (1.1)) satisfy the inequalities a11\u03b31 + a22\u03b32 > 0, a11c1 + a22c2 + \u03b31\u03b32\u0394 > 0, \u0394 > 0, (4.4) (a11\u03b31 + a22\u03b32)[a11\u03b32c 2 1 + a22\u03b31c 2 2 + \u03b31\u03b32\u0394(\u03b31c2 + \u03b32c1)] \u2212 (\u03b31c2 + \u03b32c1)2\u0394 > 0, \u0394 = a11a22 \u2212 a12a21 = a(a + f2 \u2212 f1), a = a1 + a2. (4.5) The dimensionless parameters aij (i, j = 1, 2) are determined by formulas (4.2). Consider several particular cases where Proposition 1 applies. Proposition 2. If the system is completely symmetric, i.e., if \u03b3k = \u03b3 = 0, ck = c, fk = f, k = 1, 2, a1 = a2 = a0, (4.6) then inequalities (4.4) hold for each f ; i.e., the asymptotic stability of the vehicle equilibrium considered above is preserved for any wheel rolling friction coefficients and hence for any velocities of its translational motion. The proof is by a straightforward verification of inequalities (4.4) under conditions (4.6). In this case, it follows from (4.5) that \u0394 = a2 > 0 and the first two inequalities in (4.4) are satisfied, because they are reduced to the following ones: 2\u03b3(1 + a2 0) > 0, 2c(1 + a2 0) + \u03b32a2 > 0. The last inequality in (4.4) has the form c2(1 \u2212 a2 0) 2 + 4\u03b32c2a2 0(1 + a2 0) > 0. Thus, the stability conditions (4.4) are satisfied for any f \u2265 0, and the proof of Proposition 2 is complete. Proposition 3. Assume that all conditions (4.6) except for the last are satisfied; i.e., a1 =a2 (the case of incomplete symmetry; i.e., the car body center of mass is displaced). Then, for a1 < a2, the stability conditions (4.4) are satisfied for every f \u2265 0, and for a1 > a2, the stability is violated for f > f0 = 1 a1 \u2212 a2 [ 2 + a2 1 + a2 2 + \u03b5(a1 + a2)2 \u2212 \u221a \u03b52(a1 + a2)4 + 4(a1 + a2)2 ] > 0, \u03b5 = \u03b32 c . (4.7) Remark. The quantity f0 in inequality (4.7) can be arbitrarily small for some values of the system parameters. For example, as \u03b5 \u2192 0 (small damping or large rigidity of the wheel) and for a1a2 = 1 and a1 > 0, we have f0 = (a1 \u2212 1)(1 + a2 1)/(a1 + a2 1), which can be arbitrarily small as a1 \u2192 1. MECHANICS OF SOLIDS Vol. 46 No. 4 2011 The proof is by a straightforward verification. Under the conditions given in Proposition 3, the instability inequalities (4.4) become \u03b3[2 + a2 1 + a2 2 + (a2 \u2212 a1)f ] > 0, c[2 + a2 1 + a2 2 + (a2 \u2212 a1)f ] + \u03b32(a1 + a2)2 > 0, \u0394 = (a1 + a2)2 > 0, x2 + 2x\u03b5(a1 + a2)2 \u2212 4(a1 + a2)2 > 0, x = 2 + a2 1 + a2 2 + (a2 \u2212 a1)f. (4.8) For a2 > a1, these inequalities are necessarily satisfied for any f > 0. (For the first three inequalities in (4.8), this is obvious, and for the last, this follows from the inequality x > 2 + a2 1 + a2 2 > 2(a1 + a2).) But if a2 < a1, then the first inequality in (4.8) is violated for f > f1 = (2 + a2 1 + a2 2)/(a1 \u2212 a2), the second is violated for f > f2 > f1, the third is always satisfied, and the last is violated for f \u2208 [f0, f00], where f0 is given by the formula in (4.7) and f00 is given by the same formula with the plus sign at the radical. It is easily seen that 0 < f0 < f1 < f00, which implies that all inequalities in (4.8) are violated for f > f0. The proof of Proposition 3 is complete. Proposition 4. Let f2 = 0, \u03b30 = 0, and f1 = f > 0; i.e., assume that the fore (driving) wheel does not experience any rolling resistance and there is no dissipation in it. In this case, if a1a2 < 1, then the instability arises for f > a1, and if a1a2 > 1, then the instability arises for f > (a1a2 \u2212 1)/a2. Proposition 5. Let f1 = 0, \u03b31 = 0, and f2 = f > 0; i.e., assume that the rear (driven) wheel does not experience any rolling resistance and there is no dissipation in it. In this case, if a1a2 < 1, then the instability arises for f > (1 \u2212 a1a2)/a1, and if a1a2 > 1, then the system is stable for any f > 0. The proof of Propositions 4 and 5 is by a straightforward verification of inequalities (4.4) in Proposition 1 under the assumptions stated above. Let \u03b31 > 0, \u03b32 = 0, and \u03b32 > 0. Then the following assertion holds. Proposition 6. Under the above conditions, the asymptotic stability domain D1 in the plane of the parameters (f1, f2) is given by the following inequalities. If 0 < a1a2 < 1, then 0 < f2 < 1 \u2212 a1a2 a1 , 0 < f1, c1a1f1 \u2212 c2a2f1 < c0, f1 \u2212 f2 < a1 + a2. (4.9) If 1 < a1a2 < \u221e, then 0 < f1 < a1a2 \u2212 1 a2 , 0 < f2, c1a1f1 \u2212 c2a2f2 < c0, (4.10) c0 = c1(1 + a2 1) + c2(1 + a2 2). (4.11) The proof of Proposition 6 is by a straightforward verification of inequalities (4.4) under the restrictions given in the statement. Note that if one of the inequalities in conditions (4.10), (4.11) is not satisfied, then the vibrations under study are unstable. Now let \u03b31 = \u03b32 = 0; i.e., assume that there is no dissipation in the wheels. In this case, Eq. (4.3) is biquadratic, the stability of system (4.1) can be only nonasymptotic, and the conditions for such stability are equivalent to the conditions that this biquadratic equation has only pure imaginary roots. These considerations lead to the following result. Proposition 7. For \u03b31 = \u03b32 = 0, the stability domain D0 in the plane of parameters (f1, f2) is given by the following inequalities: f1 > 0, f2 > 0, c0 > c1a1f1 \u2212 c2a2f2, a1 + a2 > f1 \u2212 f2, [c0 \u2212 (c1a1f1 \u2212 c2a2f2)]2 > 4c1c2(a1 + a2)(a1 + a2 \u2212 f1 + f2). (4.12) The proof of Proposition 7 follows from well-known necessary and sufficient conditions for the existence of pure imaginary roots of a biquadratic equation. Proposition 8. The inclusion D1 \u2282 D0 holds. MECHANICS OF SOLIDS Vol. 46 No. 4 2011 The proof of Proposition 8 is by a straightforward substitution of the extreme values in the linear inequalities (4.9) or (4.10) into inequalities (4.12) and verification that the latter are satisfied. Remark. Proposition 8 means that this nonsymmetric addition of dissipation narrows down the stability domain in the plane of the parameters (f1, f2). Graphically (schematically), the domains D1 and D0 in various possible cases are shown in Figs. 2\u201315, where digit 1 corresponds to the line f1 \u2212 f2 = a1 + a2, digit 2 corresponds to the line c1a1f1 \u2212 c2a2f2 = c0 = c1(1 + a2 1) + c2(1 + a2 2), digit 3 corresponds to the line f2 = |1 \u2212 a1a2|/a1, digit 4 corresponds to the line f1 = |1 \u2212 a1a2|/a2, digit 5 corresponds to the line f1 \u2212 f2 = \u2212|1 \u2212 a1a2|(a1 + a2)/(a1a2), and digit 6 corresponds to the parabola (c0 \u2212 c1a1f1 + c2a2f2)2 \u2212 4c1c2(a1 + a2)(a1 + a2 \u2212 f1 + f2) = 0. Figures 2, 4, 6, 8, 10, 12, and 14 correspond to the case of \u03b31 = \u03b32 = 0, while Figs. 3, 5, 7, 9, 11, 13, and 15 correspond to the case of \u03b31 = 0, \u03b32 = 0 or \u03b31 = 0, \u03b32 = 0 for all possible distinct values of the parameters a1, a2, c1, and c2. Now consider the case of weighted wheels, m1 = 0 and m2 = 0. Since we assume that the car body mass is m = 1, the quantities m1 and m2 are the respective dimensionless ratios of the wheel masses to the car body mass. Simple calculations show that the form of the vibration equations (4.1) remains the same and the coefficients aij (i, j = 1, 2) are determined by the functions a11 = \u03c12 \u2212 (a1 + am2)(\u2212a1 + f1) \u03c12 0 , a12 = \u03c12 \u2212 (a1 + am2)(a2 + f2) \u03c12 0 , a21 = \u03c12 + (a2 + am1)(\u2212a1 + f1) \u03c12 0 , a22 = \u03c12 + (a2 + am1)(a2 + f2) \u03c12 0 , where \u03c12 0 = m0\u03c1 2 and m0 = 1 + m1 + m2. Then one can readily compute that \u0394 = a11a22 \u2212 a12a21 = a(a + f2 \u2212 f1)/\u03c12 0. Proposition 1 remains valid under the above variations in aij , i, j = 1, 2, and \u0394. In a similar way, one can restate Propositions 2\u20138. Equations (4.1) and (4.2) describing the vehicle vibrations under the action of the rolling friction forces admit the following analogies (interpretations) to well-known systems in mechanics. Each of the coefficient matrices multiplying zk and z\u2022k (k = 1, 2) in (4.1) with (4.2) taken into account consists of MECHANICS OF SOLIDS Vol. 46 No. 4 2011 two parts. The first part (it is said to be unperturbed) does not contain the friction coefficients f1 and f2. The second (perturbed) part depends on these coefficients linearly. As follows from the above derivation of the equations of motion, the origin of the perturbed part significantly depends on the following two external causes. One of them is the interaction between the deformable wheels and the rough road, which causes the rolling resistance torques Mk1 and Mk2; the second cause is the torque M1 from the engine (external energy source) to the driving wheels ensuring their (and the driven wheels) permanent rolling if there are resistance torques mentioned above. If these causes are absent, then the system motion occurs for f1 = f2 = 0, and Eqs. (4.1) contain only the first parts of the above-mentioned matrices for which the solutions have the property of (asymptotic) stability (a conservative system MECHANICS OF SOLIDS Vol. 46 No. 4 2011 under the action of some forces with complete or incomplete dissipation). But if f2 1 + f2 2 = 0, then these causes result in the appearance of additional terms in Eqs. (4.1), which, according to [6], can be interpreted as positional (circular) forces (for the corresponding terms with z1 and z2); gyroscopic and nonconservative (depending on the velocities) forces (for the corresponding terms with z\u20221 and z\u20222); conservative positional forces (for the corresponding terms with z1 and z2); and dissipative (Rayleigh) MECHANICS OF SOLIDS Vol. 46 No. 4 2011 MECHANICS OF SOLIDS Vol. 46 No. 4 2011 forces (for the corresponding terms with z\u20221 and z\u20222). In the general case, such a decomposition of forces was performed in [7]. Thus, the existence of the above-mentioned external causes leads to the appearance of additional forces of the described structure, whose destructive character is well known in mechanics of gyroscopic systems, structural mechanics (in particular, in mechanics of bridges), and other adjacent fields; it is described in the literature [8\u201310]. Even if there forces do not actually lead to the vehicle instability, they still significantly change the frequencies of its vibrations in the vertical plane, and this fact cannot be ignored when solving the corresponding problems. These facts were already considered earlier by the authors in the paper [11]." + ] + }, + { + "image_filename": "designv11_33_0003752_etfa.2019.8869400-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003752_etfa.2019.8869400-Figure1-1.png", + "caption": "Fig. 1. Considered coordinate frames for the skid-steering tracked robot.", + "texts": [ + " Trajectory feasibility is guaranteed by recurring to set-based arguments taking into account constraints on control inputs, trajectory tracking error and network delay. 978-1-7281-0303-7/19/$31.00 \u00a92019 IEEE 933 In Section II mathematical modelling of skid-steering network-controlled mobile robot is proposed. In Section III the trajectory tracking control design for a network controlled robot is tackled. In Section IV the feasible motion planning algorithm is introduced. Finally, in Section V numerical simulations are discussed in order to show the effectiveness of the proposed algorithm. Given the inertial reference frame E, see Fig.1, let q \u201c \u201c x y \u03b8 \u2030T be the robot pose vector. The nominal kinematic first-order model is 9q \u201c Gpqq \u00a8 u\u0303 (1) with Gpqq \u201c \u00bb \u2013 cos \u03b8 0 sin \u03b8 0 0 1 fi fl (2) being u\u0303 \u201c \u201c V\u0303 \u03c9\u0303 \u2030T where V\u0303 and \u03c9\u0303 are the effective forward and rotational robot velocities respectively. The relationship between u\u0303 and the vector of effective angular velocities of the track sprockets w\u0303 \u201c \u201c w\u0303R w\u0303L \u2030T can be written in the following form u\u0303 \u201c J \u00a8 w\u0303 (3) being J \u201c \u201e R{2 R{2 R{D \u00b4R{D (4) where D is the distance between tracks and R is the radius of track sprockets", + " The relationship between w\u0303 and the controlled track sprockets angular velocities w \u201c \u201c wR wL \u2030T is shown below w\u0303 \u201c H \u00a8 w (5) where H \u201c \u201e \u00b5R 0 0 \u00b5L (6) Be u \u201c \u201c V \u03c9 \u2030T (7) the vector where V and \u03c9 denote the forward and rotational control velocities respectively. The kinematic relationship between (3) and (7) can be expressed as u\u0303 \u201c J \u00a8H \u00a8 J\u00b41 \u00a8 u (8) Then, (1) can be rewritten into the following form 9q \u201c Gpqq \u00a8 J \u00a8H \u00a8 J\u00b41 \u00a8 u (9) Let L be a reference system whose origin is located in M with xL directed as the MN segment as shown in Fig.1. Robot pose expressed in L has the following form qL \u201c RLEp\u03b80q \u00a8 pq \u00b4 q0q (10) being RLEp\u03b80q \u201c \u00bb \u2013 cosp\u03b80q sinp\u03b80q 0 \u00b4sinp\u03b80q cosp\u03b80q 0 0 0 1 fi fl the roto-translation matrix from E to L whereas q0 \u201c \u201c x0 y0 \u03b80 \u2030T represents the center of L expressed in E. Be TDL p\u00a8q \u201c qDL p\u00a8q , u D L p\u00a8q ( the desired trajectory expressed in L. TDL p\u00a8q has the form of couples pose and control inputs such that qDL p\u00a8q is compliant with (9) over the horizon r0, t\u0302s by assuming uDL p\u00a8q \u201c \u201c V DL \u03c9DL \u2030T (11) being V DL and \u03c9DL the nominal mobile robot forward and rotational velocities respectively where \u03c9DL \u201c 2V DL D \u00b5DR \u00b4 \u00b5 D L \u00b5DR ` \u00b5 D L being \u00b5DR and \u00b5DL nominal values of friction coefficients" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002941_icelmach.2018.8506824-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002941_icelmach.2018.8506824-Figure1-1.png", + "caption": "Fig. 1. Geometry of investigated machine stator.", + "texts": [ + " Additionally, a simplified radiation model is utilized and combined with the natural convection model for a more realistic boundary definition. The air flow through the cooling channels of the stator is measured for the respective operating points in order to determine the actual exit speed. This way, the influence of partial blockage of specific cooling ducts due to the terminal box is taken into account. The methodology is applied for the thermal characterization of a frameless Medium Voltage (MV) Interior Permanent Magnet (IPM) traction motor prototype. The modelled geometry of the machine stator is illustrated in Fig. 1. It includes the insulation sheets between the two layers of the winding, non-magnetic wedges and cooling channels. The windings are modelled as bulk copper regions with covering insulation, having mass values equal to the real one [16]. End winding regions are neglected to simplify the model. The actual winding configuration and rotor geometry are not illustrated due to confidentiality reasons. Heat sources related to friction are not considered. The local heat sources in the stator coils are defined as follows: 2 20 , 3 1 293 oph C loc w Cu w I R p T T K V (1) where Rph|20 o C is the measured resistance of the winding at 20oC ambient temperature, Vw is the volume of the winding and \u03b1Cu is the thermal coefficient of copper" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003770_1059712319882105-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003770_1059712319882105-Figure1-1.png", + "caption": "Figure 1. Body-referenced co-ordinate system on an AUV.", + "texts": [ + " The non-linear depth and yaw controller was proposed to enhance stability in the closed-loop system with variable saturation. However, it needs improvements in the performance with respect to the steady state. Thus, these limitations were considered for the enhancement of adaptive second-order sliding mode control effectively in the upcoming research works. The dynamic form of a slender body AUV contains two frames namely inertia-referenced frame and bodyreferenced frame as in Cui et al. (2016). These frames can be used to illustrate the motion of an AUV. As revealed in Figure 1, the AUV is centered at the center of buoyancy (CB) and thus has the body-referenced xaxis forward, y-axis to up, and z-axis to the port. Further, from the bridge of the body, the roll (u) about the x-axis is in positive counterclockwise direction, yaw (c) about the y-axis is in positive turning toward the left, and pitch (u) about the z-axis is positive toward bow-up direction. Generally, the 3D model can be divided into three interacting subsystems for the horizontal plane, vertical plane, and roll" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003893_0954406219888240-Figure27-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003893_0954406219888240-Figure27-1.png", + "caption": "Figure 27. 3D model of the Tricept parallel manipulator.", + "texts": [ + " If the pure contact force along the w0-axis, the deformation is the double of the result of equation (33), and it is \u00bc 8F E \u00f0LU c7\u00de \u00f042\u00de If the constraint is a pure moment and around w0-axis, the contact moment is \u00bc 9M2 8L5 U\u00f0E \u00de 2 \u00f0d2U c2dU\u00de 1=2 !1=3 \u00f043\u00de Therefore, the maximum contact deformation of the universal joint is Finally, translating the contact deformation models into the local coordinate system and taking the result and equation (40) into the equation (31), the maximum contact deformation model of the combined spherical joint can be obtained. Application case: A 1PU\u00fe 3UPS parallel manipulator As shown in Figure 27, the 1PU\u00fe 3UPS parallel manipulator is chosen to show the effects of different spherical joints on the parallel manipulator, where P stands for prismatic pair, U stands for universal joint, and S stands for combined spherical joint.26 The 1PU\u00fe 3UPS parallel manipulator is constituted of a moving platform and a static platform. Two platforms are connected by three identical active UPS Pcon R \u00bc \" 4Fu E LR , 4Fv E LR , Fw 2E ffiffiffiffiffiffi AR r :, 9M2 u 8L5 R\u00f0E \u00de 2 \u00f0d2R c2dR\u00de 1=2 !1=3 , 9M2 v 8L5 R\u00f0E \u00de 2 \u00f0d2R c2dR\u00de 1=2 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003341_s42417-019-00128-x-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003341_s42417-019-00128-x-Figure1-1.png", + "caption": "Fig. 1 Bump foil bearing with co-ordinate system", + "texts": [ + " Influence of foil bearing design parameters viz. compliance, clearance and speed on threshold rotor mass are investigated. In the foil bearing, there is a compliant structure integrated with bump foil and top foil. The hydrodynamic pressure develops between top foil and shaft. The foil deflection and relative sliding takes place due to pressure [7, 10]. The sagging effect in the top foil of meso-scale bearing is relatively small due to less value of bump pitch and hence neglected [4]. The coordinate system of foil bearing is shown in Fig.\u00a01. For the isothermal condition, the non-dimensional form compressible unsteady Reynolds equation can be written as [7, 16] The non-dimensional terms in Eq.\u00a0(1) are (1) \ud835\udf15 \ud835\udf15\ud835\udf03 ( p\u0304h\u03043 \ud835\udf15p\u0304 \ud835\udf15\ud835\udf03 ) + \ud835\udf15 \ud835\udf15z\u0304 ( p\u0304h\u03043 \ud835\udf15p\u0304 \ud835\udf15z\u0304 ) = \ud835\udeec \ud835\udf15 \ud835\udf15\ud835\udf03 ( p\u0304h\u0304 ) + 2\ud835\udeec\ud835\udf08 \ud835\udf15 \ud835\udf15\ud835\udf0f ( p\u0304h\u0304 ) . (2) p\u0304 = p pa , h\u0304 = h C , z\u0304 = z R , \ud835\udf0f = \ud835\udf14t, \ud835\udf08 = \ud835\udf14s \ud835\udf14 , \ud835\udeec = 6\ud835\udf07\ud835\udf14 pa ( R C )2 , where p\u0304 , h\u0304 , z\u0304 , , and are pressure, film thickness, bearing axial length, time, speed ratio and bearing number, respectively. The film thickness is In Eq.\u00a0(3), is eccentricity ratio and subscript represents X and Y direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000199_intlec.2009.5351813-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000199_intlec.2009.5351813-Figure5-1.png", + "caption": "Figure 5. The reactive-power flow ofDFIG", + "texts": [ + " It becomes a secondary excitation generator which rotates with super-synchronous speed. The power flow at this time is shown in Fig. 6. Here, the secondary loss PLOSS is neglected. B. Reactive power flow Next, the reactive power flow of an induction machine is described. Set excitation reactive power of an induction machine to Qm' and let conversion of the reactive power from Ql and a rotation child to a stator be Q21 and the rotation child reactive power Q2 for stator reactive power. (Direction makes delay reactive power positive and expresses it in Figure 5) It will be set to Ql =0 (power-factor I), if reactive power is poured in from the rotation child side so that it may be now set to Q21 =Qm =SQ2 \u00b7 Since the control of reactive power can output big reactive power from the stator side with reactive power with smaller controlling from the low rotation child side of frequency, it is efficient. In the case of a slip s =0 , the reactive power of infmite size can be supplied from the stator side. Therefore, the synchronous generator which operate direct-current excitation can output big reactive power by controlling field system voltage" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002248_s00170-017-1085-4-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002248_s00170-017-1085-4-Figure2-1.png", + "caption": "Fig. 2 A multiple robot system with coordinated motion", + "texts": [ + " Therefore, each robot arm is subjected to the coupling of other manipulators, and the performance of the stiffness of the entire systemwill differ from a single robot. In order to improve the position accuracy and structural performance of a coordinated multiple robot system, it is necessary to investigate the system stiffness. In order tomake the definitionmore accurate, the following assumptions are made: (1) the robot\u2019s end effector holds the object tightly with no slip movement; (2) all robot joints are active and can be controlled; (3) the entire system is in a static equilibrium state when in operation such as drilling as shown in Fig. 2. To model the system stiffness, the deflection of a point that is attached to the workpiece is selected as the observation point as shown in Fig. 2 for this definition. Then, the stiffness of a coordinated multiple robot system is defined as the value of the required force to produce a unit deflection at the selected point. This can be shown in matrix form. If the external force vector acted on the workpiece is expressed as [fx, fy, fz, nx, ny, nz], i.e., three force components and three moment components, correspondingly, the infinitesimal displacement vector at the selected point is expressed as [dx, dy, dz, \u03b4x, \u03b4y, \u03b4z], i.e., three linear deformation components and three angular deformation components, then the stiffness matrix is used to relate the force vector to the displacement vector as below" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001898_j.proeng.2017.01.158-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001898_j.proeng.2017.01.158-Figure4-1.png", + "caption": "Fig. 4. The partition of blade surface.", + "texts": [ + " However, due to the adjacent blade basin, blade back and the flow channel need to be processed at one consistent path, and a blade cannot be processed completely, so the blade should be divided into two symmetrical areas, the division is based on the highest and lowest point of blade in the direction of blisk height. If the central axis of the blisk is Z-axis, the highest and lowest points can be obtained by utilizing a plane which is perpendicular to Z-axis and tangent to the U-direction iso-parametric line of the blade root, and the tangent points are the extreme points. The V-direction iso-parametric lines where the extreme points lie are the dividing lines. At this time a complete blade surface is divided into six areas, shown in Fig. 4. After the machining area is divided, it is necessary to plan the cutter location points on each area. In this paper, the most commonly used ball-end cutter which treats the cutter center point as the cutter location point is used to process the tool path. [8]When machining a surface with a ball-end cutter, the cutter center point is constrained to the surface which was offset a tool radius R by the machined surface. The cutter center point coordinates can be obtained by offsetting cutter contact point by a tool ball radius along the surface normal vector direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002367_s12206-017-1114-0-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002367_s12206-017-1114-0-Figure8-1.png", + "caption": "Fig. 8. An illustration of physical interference between a workpiece and the tool of DED process in the previous stepwise multi-axis slicing algorithm.", + "texts": [ + " 7 shows three different computer generated images of sliced layers each of which is generated from corresponding slicing methods; 2-D slicing method, multi-axis stepwise slicing method using planar layers, and multi-axis slicing method using an SCL model with relative to an example .STL model. The previous approach of multi-axis stepwise slicing algorithms first decomposes a part into sub-volumes, each of which can be completely built along a certain direction by simply rotating the slicing direction when needed, which, however, prone to unwanted physical interferences between the laser-nozzle assembly and the workpiece (or platform plate) when moving with each other as illustrated in Fig. 8. Even an SCL model, naturally provides an interference-free 5- axis tool-part movement along the spherical surface when building the part, is also suffering a problem of tool-plate interference in the beginning of the build process. This problem of physical interference can be avoided by measuring 2-D motion space of the platform plate tilting & rotating and 2-D tool space translating continuously checked and controlled during the build. In this study, the motion space of the circular platform plate tilting only can be modeled as a parametric circle of which the radius R is corresponding to maximum distance from the plate origin O to any tip of the platform plate (2-D motion space of the platform plate)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002503_1350650117753915-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002503_1350650117753915-Figure4-1.png", + "caption": "Figure 4. CFD results at 10 kr/min shaft speed. (a) Pressure distribution (b) Temperature distribution and (c) Air entrainment.", + "texts": [ + " With this model, the displacements applied on domain boundaries or in subdomains are diffused to other mesh points by solving the equation: r dispr \u00bc 0 \u00f018\u00de where is the displacement relative to the previous mesh locations and disp is the mesh stiffness, which determines the degree to which regions of nodes move together. This equation is solved at the start of each outer iteration or time step for steady state or transient simulations, respectively. Experiment1 is conducted using VG22 as lubricant. The dynamic viscosity of oil is expressed as eff T \u00bc 16537T 1:788 \u00f019\u00de where T is the temperature in C. Equation (19) is for oil VG22 at 40\u201390 C. Figure 4 shows the steady CFD calculation result when shaft speed is 10 kr/min. As shown in the figure, pressure, temperature, and air volume fraction distribution is presented in detail. The shaft and ring rotate around Z axis negative direction. And rotation direction remains the same in the following figure. These results show one of the strengths of the CFD. This method is capable of capturing the details of fluid field. Although no experiment result is available on the pressure and temperature distribution in the literature,1 the pressure and temperature distribution is reasonable according to author\u2019s experience. Furthermore, Figure 4(c) shows the air entrainment results. Air flows into the fluid domain in large scale when shaft speed is 10 kr/min. This finding means that air is an important part of the inner film when shaft speed is 10 kr/min. The validation of the air entrainment is presented subsequently. Hatakenaka and Yanai1 reported that the leveling off of the ring speed is related to the subsynchronous vibration. To validate the effect of air entrainment, the calculation results on the rotation speed of the floating ring are selected as the parameter to be compared" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002497_978-3-319-69480-1_13-Figure13.5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002497_978-3-319-69480-1_13-Figure13.5-1.png", + "caption": "Fig. 13.5 Finite element model", + "texts": [ + " In order to minimize computational cost, an axisymmetric modeling technique has been adopted, in which only a cross section of the cylinder is modeled. Also, due to the axial symmetry of geometry, loading, and boundary conditions, a symmetry boundary condition was introduced at the middle plane of the cylinder so that only half the cylinder is needed to be modeled. These simplifications provided significant improvements to the computational efficiency without any sacrifice of the accuracy of the results. Figure 13.5 shows a comparison between the theoretical model and the actual FEA model. The material model presented in Sect. 13.2.3.3 and the material characteristics presented in Sect. 13.3.1 have been used in the numerical simulation. The material characteristic curve used in the theoretical analysis is discretized and introduced as an array to the model. Axisymmetric solid elements are utilized. In order to avoid shear locking and hourglass effects, second-order quadratic elements were selected for the analysis, which provides parabolic distribution of the deformations and results in high degree of accuracy of the model stiffness" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002938_icelmach.2018.8507253-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002938_icelmach.2018.8507253-Figure2-1.png", + "caption": "Fig. 2. Structural von-Mises stress distribution at maximum speed 60,000 rpm considering the different rotor structures.", + "texts": [ + " It can be seen that the machine M1 can provide the maximum torque with lower current than the other machines. This is due to a higher PM flux linkage. However, this becomes a disadvantage at high speed where M1 needs higher d-axis current for flux weakening operation. Nevertheless, the performances of the machines are similar at high speed operating point (Joule losses and PM losses), except the core losses that are higher for M1. The machines have been analyzed using structural FEA. The equivalent stress distributions for the maximum speed (60,000 rpm) are shown in Fig. 2. It can be seen that M2 is less resilient against the centrifugal force than the other machines. This is due to gap existing between poles which creates a local stress concentration in the sleeve. However, this can be reduced by using a pole gap filler that has the same mechanical and thermal properties as those of the PMs. In this study, a glass fiber has been used and the stress has been reduced by 268 MPa. Inset PM machine has also been considered. In this case, the pole gap filler is the iron, which has similar properties to those of SmCo magnets (density, thermal dilatation," + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001319_icelmach.2012.6350265-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001319_icelmach.2012.6350265-Figure5-1.png", + "caption": "Fig. 5. Current flow FE model for brush-segment contact resistance", + "texts": [ + " Generally, the voltage drop between a brush and a segment is modeled by a variable resistance due to two phenomena: the resistance of the brush body and the voltage drop at the sliding interface. For the brush body voltage drop, we can consider in a first approach this internal resistance as an inverse function of the brush-segment surface contact written \u0394S. However, this approach is not sufficient because the current flow is not homogeneous in the brush and the segments. For more accuracy, a field model using current flow equations can be used to identify the equivalent contact resistance (Fig. 5). Due to the double-layered brushes composition, a finite element model seems more suitable. The purpose of this second layer (\u03c12 >> \u03c11) is the increase of resistance at the end of contact and thus the acceleration of the extinction of the current. The voltage drop at the sliding interface is often studied separately by using an empirical equation [2]. It is due to the fact that electrical contact is not perfect but the current passes throw many micro-structures, causing locally very high current densities. This contact surface quality depends on the pressure of springs; the mechanical and physical characteristics of materials used and of course film quality. We propose to study those two phenomena in a same model by adding a highly resistive thin layer at the contact (\u03c14>>\u03c13) as showed Fig.5 to take into account this phenomenon. Thus, the distribution of current density in the brush is complex and a Finite element model seems more relevant. We propose to identify this resistance through using a finite element calculation. The FE computation of the current flow is based on a scalar potential formulation. The current density J (shown Fig. 6) is obtained from differentiation of the field of potential. The two surfaces represented with black lines in Fig. 5 are considered as two Dirichlet boundary conditions, respectively with V1 and V2 voltages. The V1 potential is related to the potential of the brush cable and the V2 potential at the bottom of a segment. Resistance is finally deduced from the Finite Element calculation of the electrical power W using the definition: iiibr JAlW 2 (1) where: brl is the axial length iJ is the current density of the element i iiA and are the surface and the electric resistivity of The model is therefore based on some assumptions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001841_0278364916679719-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001841_0278364916679719-Figure8-1.png", + "caption": "Fig. 8. Two experimental setups of a legged mechanism with a movable center of mass.", + "texts": [ + " The region R\u0303 associated with such contacts, shrinks to a single line segment, shown with a dashed line at the center of Figure 7(b). This section describes experiments that measure the feasible equilibrium region of a legged mechanism prototype having a movable center of mass, which is supported on a frictional terrain against gravity. The objective of these experiments is to validate the analytical characterization of the feasible equilibrium region R\u0303 of tame stances supported by multiple frictional contacts. As shown in Figure 8, the legged mechanism consists of a rigid central ring made of aluminum, extendable legs made of steel, and a movable heavy steel cylinder that determines the location of the mechanism\u2019s center of mass. The central ring has a diameter of 240 mm and is drilled with 36 holes that allow attachment of three, four, or five legs. Each leg can be assembled in three different lengths of 30 mm, 60 mm, and 90 mm. Each leg ends with a spherical footpad made of steel, which maintains a frictional point contact with the supporting terrain. The supporting terrain consists of stainless steel plates whose slope angles can be adjusted with a lead screw mounted against a horizontal support plate (Figure 8). The heavy steel cylinder which determines the mechanism\u2019s center of mass, slides on a linear guide mounted on top of the central ring (Figure 8). The linear guide can be attached at 15\u25e6 orientation increments using screws that can be inserted in 24 holes drilled into the central ring. The mass of each extendable leg is 0.82 kg, the central ring and linear guide combined mass is 6.17 kg, and the movable cylinder mass is 4.25 kg. As a preliminary stage, the coefficient of static friction between the footpads and supporting plates, \u03bc, was experimentally measured. A horizontal force was applied to the legged mechanism while its footpads were supported by horizontal plates" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000261_iros.2009.5354810-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000261_iros.2009.5354810-Figure1-1.png", + "caption": "Fig. 1. A planar two-link space-based robot system.", + "texts": [ + " The remainder of paper is organized as follow: In section II, the dynamics of a space-based robot system is formulated. Section III presents the Jacobi relation of the system. The augmentation approach is introduced in section IV. A robust adaptive composite control of space robot system is proposed in section V. To show the feasibility of control scheme, The simulation studies are presented in section VI, which are followed by conclusions given in section VII. II. DYNAMICS OF THE SYSTEM Without a loss of generality, a planar two-link space-based robot system with payload is considered here, Fig. 1. The S 978-1-4244-3804-4/09/$25.00 \u00a92009 IEEE 2353 system consists of the base 0B , 1B (link 1) and 2B (link 2) and the payload P . We assume the end-effector hold the payload rigidly. 0O coincides with the mass center 0CO of 0B , iO )2,1( =i is the rotational center of the revolute joint between 1\u2212iB and iB , 1CO is the mass center of 1B , 2CO is the mass center of combination 2B and P , ix is the symmetrical axis of each link. The other symbols are defined as follows 0l Distance from joint 0O to 1O ; il Length of link i )2,1( =i ; ia Distance from joint iO to the mass center CiO )2,1( =i ; im Mass of iB )2,1,0( =i ; Pm Mass of P ; M Total mass of the entire system; iJ Inertial moment of iB )2,1,0( =i with respect to its mass center; PJ Inertial moment of P with respect to its mass center CPO ; )( XYO \u2212 Inertial coordinate frame of the system; )( iii yxO \u2212 Local coordinate frame of iB )2,1,0( =i ; C Mass center of the entire system; ier Unit vector pointing along with ix )2,1,0( =i ; ir r Position vector of mass center of iB )1,0( =i ; 2r r Position vector of mass center of combination 2B and P ; Cr r Position vector of the entire mass center C ; Pr r Position vector of the end-effector P ; 0\u03b8 Attitude angle of the base, which is the angle between the Y axis and the 0x axis; i\u03b8 Rotational angle of joint iO )2,1( =i , i", + " Since the matrix D is symmetric positive-definite, from (20), we know that q~ , s&\u0302 and 1\u03a6 are bounded. Considering (14) and (18), q&~ , s&&\u0302 and \u2217\u03a6 are bounded too. Based on the results, we have 0lim = \u221e\u2192 V& t , which means 0~lim = \u221e\u2192 q t , 0\u02c6lim = \u221e\u2192 s& t . Therefore, the control law (17), the adaptation law (19) and the definition of \u03a8u and du given in (24) can asymptotically stabilize the space robot system to track the desired trajectory described in terms of PdX . VI. NUMERICAL SIMULATION For illustrative purposes, a planar two-link space-based robot system shown in Fig. 1 is considered in our simulation study. The actual plant parameters of the system are as follows kg400 =m , kg22 21 == mm , kg5.2=Pm , m5.10 =l , m32 11 == al , m32 22 == al , 2 0 mkg17.34 \u22c5=J 2 21 mkg5.12 \u22c5== JJ , 2mkg5.1 \u22c5=PJ . In this simulation, the parameter vector \u2217\u03a6 and 1\u03a6 are T 654321 )( \u2217\u2217\u2217\u2217\u2217\u2217\u2217 = \u03c6\u03c6\u03c6\u03c6\u03c6\u03c6\u03a6 , T 2101 )( PPP LLL=\u03a6 . Where, 2 202 2 101 2 00001 LmLmLmJ P+++=\u2217\u03c6 , 2 212 2 111 2 01012 LmLmLmJ P+++=\u2217\u03c6 , 2 222 2 121 2 02023 LmLmLmJ PP +++=\u2217\u03c6 , 2120211101010004 LLmLLmLLm P++=\u2217\u03c6 , 2220212101020005 LLmLLmLLm P++=\u2217\u03c6 , 2221212111020106 LLmLLmLLm P++=\u2217\u03c6 , 0000 lLLP += , 1011 lLLP += , 2022 lLLP += " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001898_j.proeng.2017.01.158-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001898_j.proeng.2017.01.158-Figure5-1.png", + "caption": "Fig. 5. Processing areas in combined surface.", + "texts": [ + " The formula for calculating the interval L is: 22 2L R (1) Where R is the radius of the ball-end cutter and \u03b5 is the scallop height. And the cutting step is calculated by the equal chord deviation method. For a given maximum chord deviation max and the minimum radius of curvature min on the curve, the minimum cutting step length can be calculated as follows: max2min minl (2) Three processing areas, namely, blade back, blade basin and flow channel, need to be dealt with in a process, as shown in Fig. 5. The machining method proposed in this paper is a combined surface processing method, which needs to make the three regions of the same number of processing paths, while the arc length of the blade basin root curve is greater than that of the blade back root curve, which will result in the number of processing paths of the blade basin surface is more than that of the blade back surface under the same processing error. In order to ensure that the two surfaces have the same number of processing paths, first calculate the cutter location point on the blade basin surface, according to its total number of path lines then back to calculate the number of path lines of the blade back surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002711_978-981-13-0305-0_23-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002711_978-981-13-0305-0_23-Figure2-1.png", + "caption": "Fig. 2 Laser-based powder bed fusion additive manufacturing [20]", + "texts": [ + " Afterwards, using the energy source (Laser or electron beam) melting of the metal powder takes place tracking the outline of each layer. The foundation layer is melted and then solidified into a solid state. When manufacturing of one layer has been completed, there is a downward motion of the piston in building chamber and upward movement of the piston in powder chamber. The upward movement is equivalent to thickness of the layer. A fresh powder layer is spread over the whole working area and whole process is recurred again until the complete part is built. In laser based powder bed fusion additive manufacturing (Fig. 2), the powder is spread over the working platform with the help of blade or roller and this layer is melted with a laser beam. Then the lowering of platform equal to the thickness of the layer is done and again the powder layer is spread. This procedure is repeated till the whole component is manufactured. Generally Nd:YAG-laser and fiber laser is used for sintering of the powder. In electron beam melting (Fig. 3), high power electron beam is used which is controlled by electromagnetic coils. The electron beam heats the whole powder bed for each layer build" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003400_978-3-030-19648-6_41-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003400_978-3-030-19648-6_41-Figure4-1.png", + "caption": "Fig. 4. Free-form surface load on a tooth of a gear: (a) a gear which is constructed by SGLibrary, (b) a free-form load on a tooth of the gear, (c) feature surfaces of the gear which are calculated by PDE Toolbox", + "texts": [ + " The edges between those neighboring facets will be detected as the boundary of the feature surfaces. However, the functionality of this concept is very limited. Figure 3 shows the feature surfaces of a box and a sphere. The surface of a sphere is much smoother than the box so that only one feature surface is detected. In this case, we have no chance to apply loads on a free-form surface of the sphere if we use the feature surface concept of PDE Toolbox. Generally, if we have a PDEModel with a very complicated geometry like the gear in Fig. 4a, we are not able to apply any free-form surface load like the green strip in Fig. 4b since they cannot be detected by PDE Toolbox as single feature surface. This is a remarkable drawback of PDE Toolbox which may limit the potential of our bionic structural optimization methods since the geometry of our model is always changing during the shape optimization and the indices of the feature surfaces are also inconstant. In this paper, we introduce the concept of overlapping region to determine the area of a free-form surface for applying boundary conditions, which enables the PDE Toolbox to apply any free-form surface loads independent from the feature surfaces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000525_978-3-642-23244-2_54-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000525_978-3-642-23244-2_54-Figure6-1.png", + "caption": "Fig. 6. 3D-model of DELTA robot", + "texts": [], + "surrounding_texts": [ + "Based on the described pr is currently being subject of the accuracy and dyna individual components. These experiments hav cluding the Beckhoff com garding the working space The obtained results w the AMPER international Another task within the of the end effector. This seems to be essential for interface for control of th M. Opl et a onstructional variants of the joint couplings oposals, the DELTA robot prototype has been built th to intensive experimental tests focused on improvemen mics of the mechanisms and also on durability of th e sufficiently verified the control system functions in pany library, as well as the theoretical knowledge re size and singular positions of the DELTA robot. ere presented in an exposition of the Dyger company fair trade which took place in Brno in spring 2011. project, which is still being worked on, will be rotatio additional construction component of the DELTA robo most of the industrial applications. Also, a suitable use e robot functions will be supplemented in cooperatio l. at t e - - at n t r n DELTA - Robot with Parallel Kinematics 451 with the Dyger company. The objective is to design a marketable product which could be offered for applications in industry (see Fig. 8). Acknowledgemet This work is supported by Brno University of Technology, Faculty of Mechanical Engineering project NO. FSI-S-10-27, CZ.1.07/2.3.00/09.0162 and by company DYGER, s.r.o." + ] + }, + { + "image_filename": "designv11_33_0000041_cae.20391-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000041_cae.20391-Figure5-1.png", + "caption": "Figure 5 A 3D model of the identification station.", + "texts": [ + " There are four main stations within this training unit. The detailed definitions of the stations are as follows: This station is the first station of the MTMPS unit and includes a part magazine and two sensors as optical (B6) and inductive (B7) to differentiate material types. Inserted materials are detected by means of an optical and an inductive sensor. If signal received from B6 and B7 sensors, the part is metal. If signal received only from B6 optical sensor, the part is plastic. Testing station is shown in Figure 5. The MTMPS is equipped with a three-axis handling device. Transfer cylinder carries parts by means of vacuum handling and a pressure sensor also controls the vacuum. Part handling by means of vacuum between stations modules was realized via Cartesian axes movements and could be called as a Cartesian Robot. X- and Y-axis are obtained by using two DC gear motors (D.AS-DC-G-24-2) and two slide units, spindle drive (D.ERKSP-250) and Z axes movement is given by the pneumatic transfer cylinder [11]. Processing Station consists of four modules, turntable module, marking module, drilling module, and cleaning module as shown in Figure 6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001626_s0219878910002117-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001626_s0219878910002117-Figure2-1.png", + "caption": "Fig. 2. Difference of surface conditions in the grasp of two objects (i.e., curvature and friction at contact point).", + "texts": [ + "w or ld sc ie nt if ic .c om by N A T IO N A L U N IV E R SI T Y O F SI N G A PO R E o n 11 /2 4/ 15 . F or p er so na l u se o nl y. To improve the dexterity of the hands, Yamada et al. [2005, 2010] investigated grasp stability of two objects in two dimensions. In this case, the stability depends not only on contact position, contact curvature, contact force, and contact friction between each object and each finger but also on contact position, contact curvature, and contact friction between two objects as shown in Fig. 2. Stiffness matrices of the grasp including these effects were explicitly derived. Grasp stability is evaluated by positive definiteness of the matrices. 1.3. Approach of this paper This paper attacks grasp stability of multiple planar objects. Each finger is replaced with a 2D spring model, and the grasp stability is analyzed from the viewpoint of potential energy (Appendix A). Grasp stiffness matrices are derived with consideration of contact friction property. Grasp stability is evaluated by eigenvalues and eigenvectors of the matrices" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001308_s1068798x12020189-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001308_s1068798x12020189-Figure1-1.png", + "caption": "Fig. 1. Equipment for biplanar dynamic balancing: (1) rotor of electric motor; (2) stator of electric motor; (3) vibroacoustic sensor of second bearing (second chan nel); (4) flywheel; (5) laser marker; (6) motor shaft; (7) motionless base; (8) pulley; (9) reflective tape; (10) vibroacoustic sensor of first bearing (first channel).", + "texts": [ + " In addition, the lower speed limit should be no less than 10 Hz, so as to prevent incorrect readings of the vibroacoustic sensors employed. We consider final balancing, in which the first har monic of the vibrosignal is less than the permissible value (ideally, the null value). For example, in the 3A110 grinding machine with a high speed drive, vibrations that impair product quality are observed. The main source of vibrations is the AIR132 induction motor. To study the influence of vibration on the prod uct quality, we investigate the dynamic characteristics of the system by means of a vibrodynamic test bench (Fig. 1). The electric motor is started at a rotor speed of 600 rpm (frequency f = 100 Hz). In what follows, we consider time scans of prelimi nary startup that illustrate important aspects of fre quency analysis (Fig. 2a). They identify some typical characteristics of vibration spectra. The sinusoidal vibration contains several frequency components. It is evident from Fig. 2a that two main harmonics are present: a low frequency harmonic (20 Hz); and a high frequency component (100 Hz; Fig. 2a, graphs 1 and 2)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003754_j.triboint.2019.106023-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003754_j.triboint.2019.106023-Figure1-1.png", + "caption": "Fig. 1. Schematic diagram of the hydrogen uptake apparatus.", + "texts": [ + " The focus of this investigation is to understand the complex interplay between the lubricant composition and the tribofilm effect on hydrogen permeation rate. The newly developed technique, described in our previous paper [14], is used to determine whether lubricant formulation provides any improvement with regards to hydrogen uptake into the steel. The influences of two commonly used additives, zinc dialkyldithiophosphate (ZDDP) and molybdenum dithiocarbamate (MoDTC), as well as water contamination on hydrogen permeation rate were examined in this research. Fig. 1 shows a schematic illustration of the developed apparatus that enables quantitative measurement of hydrogen permeation from a lubricated contact. This setup is made of a lubricated tribological part and an electrochemical hydrogen detection part. These two parts are separated by a thin steel membrane fastened and sealed at the bottom of the oil bath. Hydrogen is generated from the rubbing contact between the steel membrane and a metal counterpart. The larger amounts of hydrogen atoms on the surface of the steel leads to the diffusion of an unknown portion of the atomic hydrogen into the steel membrane" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003466_j.mechmachtheory.2019.06.017-Figure21-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003466_j.mechmachtheory.2019.06.017-Figure21-1.png", + "caption": "Fig. 21. Coordinate systems applied for TCA.", + "texts": [], + "surrounding_texts": [ + "The tooth surfaces p of tapered involute pinion and the normal vector to the surfaces can be expressed in coordinate system S p by: r p ( u p , l p , \u03c8 p ) , f rp ( u p , l p , \u03c8 p ) = 0 n p ( u p , l p , \u03c8 p ) (55) The face-gear surfaces 2 and the normal vector are expressed in coordinate system S 2 by: r 2 ( u s , l s , \u03c8 s ) , f s 2 ( u s , l s , \u03c8 s ) = 0 n 2 ( u s , l s , \u03c8 s ) (56) With the assistance of the coordinate transformation matrices M fp and M f 2 , the continuous tangency of the face-gear drive can be expressed in fixed coordinate system S f: r ( p ) f ( u p , l p , \u03c8 p , \u03c6p ) \u2212 r ( 2 ) f ( u s , l s , \u03c8 s , \u03c62 ) = 0 , n ( p ) f ( u p , l p , \u03c8 p , \u03c6p ) \u2212 n ( 2 ) f ( u s , l s , \u03c8 s , \u03c62 ) = 0 , f rp ( u p , l p , \u03c8 p ) = 0 , f s 2 ( u s , l s , \u03c8 s ) = 0 (57) The system indicated by Eq. (57) contains eight unknowns. Considering the condition | n f (p) | = | n f ( 2 ) | = 1, only seven inde- pendent nonlinear scalar equations are included. Typically, the bearing contact on the face-gear surfaces can be localized in the condition that N s - N p = 1 \u223c3. N p repre- sents tooth number of the tapered involute pinion. As shown in Fig. 22 , asymmetric bearing contact is detected on the concave side and convex side of the face-gear. The distribution of the contact points and the orientation of the major axis of the contact ellipse on the convex side are more beneficial for avoidance of the edge contact. Thus, the transformation on the convex side is preferred. This means that during the design of a helical face-gear drive, a proper spiral direction of the face-gear tooth must be determined with the direction of rotation into consideration." + ] + }, + { + "image_filename": "designv11_33_0000345_6.2009-5890-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000345_6.2009-5890-Figure8-1.png", + "caption": "Figure 8. NextGen wing structural components.", + "texts": [ + " 9 of 25 American Institute of Aeronautics and Astronautics The sweep dynamics are mostly affected by forces induced by the aerodynamic load whereas the chi dynamics are mostly affected by forces induced by skin stretching. The following sections describe the mechanical and actuation characteristics of these mechanisms and the resulting model of the kinematics, actuation, constraints, and low-level control. Chi Mechanism The morphing wing structure consists of a main center spar (Spar 2 ) with one forward spar and two aft spars connected by four forward and four aft pivoting ribs (Figure 8). A carbon fiber leading and trailing edge as well as aileron assembly are attached to leading and trailing edge (Spars 1 and 4 ) spars, respectively. The flexible skin/rib structure is suspended between the spars. By design, the acute angle of the members is limited to a range of 29 to 76 degrees. The opening and closing of this angle represents the chi degree-of-freedom of the wing. The forces encountered by the chi mechanism can be attributed to two primary sources: flight loads and skin forces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000151_iemdc.2009.5075258-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000151_iemdc.2009.5075258-Figure6-1.png", + "caption": "Fig. 6. Schematic view of impression of homopolar flux", + "texts": [ + " Geometry faults or deviations like the tooth shape and parameters or the concentricity of the diameters and the width of the air gap can be studied. Theoretically, the 2D model provides a good basis for an optimized machine design (shape of the lamination and most geometric parameters). Until now no 2D model of the step motor seems to exist which can combine the unipolar effect with rotation of the rotor and excited coils. A major challenge is the impression of homopolar flux (see dotted radial arrows fig. 6). The best tool would be a magnetic voltage source coupled via external circuit to the inner and outer radius. However, we do not have software in which the feature is implemented. It can be seen by the following expressions, that either a vortex field can be impressed by putting positive an negative potential on two edge points of the geometry or use flux between two lines what shall produce the same effect (see fig. 6). rot A = B (9) Phi = \u222b\u222b A V d a = \u222e l A d s (10) For a full description of the problem the definition of boundary condition is necessary. Typically two basic types are available: the Dirichlet boundary and the Neumann boundary conditions which are at the nodes. Dirichlet indicates the value and Neumann the derivation of the value in the node (see fig. 5a). For electromagnetic fields, the boundary is set for field or flux density. One can, for example, imagine that in an infinite distance to the object the potential is (A) = 0 (Calibration)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003469_j.engfailanal.2019.06.024-Figure15-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003469_j.engfailanal.2019.06.024-Figure15-1.png", + "caption": "Fig. 15. The deformation and Mises stress of the cage.", + "texts": [], + "surrounding_texts": [ + "Considering the effects of mesh quality, especially the contact between needle roller and cage, the influence of different mesh sizes on the maximum stress of cage and needle roller were analyzed with the fillet radius is 0.5 mm, the results under different grids are shown in Fig. 13, and the calculated errors under different grids on the basis of minimum grid stress are shown in Table 2. Considering the calculation time and error value, the element size of cage is selected to 0.5mm, the element size of needle is selected to 0.3mm in subsequent analysis,\n6. Experiments and result analysis\n6.1. Experimental verification\nIn order to verify the theoretical and FE results, the laboratory simulation experiments are carried out on the dynamic performance test bench of downhole tools (as shown in Fig. 3), the specific parameters areTable 1.\nSince the experiments are based on existing sensors of the tool, the verification should depend on the deflection angle of the drill bit, which is considered to the deflection angle of the lower end of the main shaft (x=0) [30]. Taking maximum offset distance of the eccentric mechanism as the value of 6mm and other parameters into the above equation, it can be obtained that the offset force P= 12,409 N, the deflection angle of the lower end of the main shaft is 0.690\u00b0(x=0), and the supporting reaction force of the selfaligning bearing is R1= 7899 N. The experimental results show that the tool deflection angle is 0.6389\u00b0 [31], which is very close to the theoretical results. In order to further analyze the correctness of the results, the deflection angles of drill bits at different offset distance are compared as shown in Fig. 14. From the results we can know that the error between the results of theory, FE and experimental is small. The reliability of theoretical analysis and FE analysis are illustrated by experiments. The value of FE results is less than experimental and theoretical values, since the finite element model is simplified, the influence of shell and other factors is not considered, therefore, it is necessary to consider the stiffness of the shell in the theoretical model.", + "The maximum total deformation and stress of needle roller and cage are obtained by simulating the same size of the original invalid bearing by FE model, fillet radius of needle roller is set to 0.25mm. As shown in Figs. 15 and 16, the maximum von Mises stress is at one end of the needle, almost the same as the actual failure figure of the needle roller. Stress concentration occurs at the edge of cage, which is also consistent with the failure position of the cage. From the results we can know the maximum Mises stress of the needle is 1332MPa, which isalmost equal to the ultimate contact stress [p]= 1500MPa. The maximum Mises stress of the cage is 292.46MPa, which is less than yield stress of cage material \u03c3b= 325 MPa.\nTable 2 Relative error of the maximum stress.\nElement size of cage 2.0mm 1.0 mm 0.8mm 0.5 mm 0.3mm\nError of needle roller (%) \u221223.663 \u22126.979 \u22122.183 \u22120.202 0 Error of cage (%) \u221241.224 \u221221.251 \u221212.9018 \u22122.153 0 Element size of cage 1.0mm 0.8 mm 0.5mm 0.3 mm 0.2mm Error of needle roller (%) \u221212.156 \u22125.853 \u22121.969 \u22120.505 0 Error of cage (%) \u221233.029 \u221222.950 \u221213.318 \u22122.801 0\nan gl\ne/ \u00b0", + "According to the deformation diagram of the needle roller, the maximum inclination angle of the needle roller is produced:\n= \u2212 \u03b2 \u03b3 \u03b3\nL arctan\nb\nmax min\n(20)\nwhere is the maximum inclination angle of the needle roller,\u00b0; \u03b3max, the maximum deformation at one end of needle roller, m; \u03b3min, the minimum deformation at the other end of needle roller, m.\nIt can be obtained \u03b2=0.0524\u00b0. According to the relative fatigue life of the roller bearing at different inclined angles [32\u201333], if the fatigue life of the bearing without inclined angle is set to 1, the relative fatigue life of needle roller bearings is only 0.02 while\u03b2=0.05\u00b0.\nFrom the above analysis we can know the wear of the right end of the needle roller is obviously larger than the left side, which is consistent with the actual failure form of roller needle and the theoretical mechanical law.The needle and cage edge produce a stress concentration, that gradually wear the roller needle and cage. The inclined angle seriously affects the use of needle roller bearings and accelerates the failure of the needle and cage, which eventually leads to the failure of the cantilever bearing.\n6.2.2. Influence of fillet radius on needle roller Another problem with the failure needle roller is that there is no fillet or the fillet radiusare too smallon the both ends of needle, the initial fillet of needle roller is shown in Fig. 17. Based on the FE model, the failure of the cantilever bearing with different fillet radius (r=0mm, 0.25mm, 0.5 mm, 0.75mm, 1mm, 1.25mm, 1.5 mm) was studied. The simulation results of rolling needle and cage are shown in Fig. 18 and Fig. 19, while under different fillet radius of needle rollers.\nFigs. 18 and 19 show that it caneffectively reduce the maximumvon Mises stress, total deformation and inclination \u03b2 while appropriately increasing the fillet radius. For example, the maximum stress of the needle roller and the cage are 684.36MPa, 98.878Mpa, and the inclination is 0.0272\u00b0while the fillet radius is 0.75mm. The maximum stress, deformation and inclination reduced by 48.62%, 66.19% and 48.09% respectively while compared to fillet radius is 0.25mm. The results show that the fillet radius of needle roller have obvious influence on the maximum stress, deformationand inclination of the needle roller and the cage. By increasing the fillet angle to about 0.75mm, the maximum stress and deformation can be reduced." + ] + }, + { + "image_filename": "designv11_33_0002779_1464419318789185-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002779_1464419318789185-Figure5-1.png", + "caption": "Figure 5. Simulation results by the FE model with 134.09 N m to the gear and 2000 r/min to the pinion: (a) displacement distribution of the assembly, (b) Mises distribution of the pinion, and (c) CPRESS distribution of the gear.", + "texts": [ + " The FE model of the pair is solved using implicit integration algorithm, and three meshing periods under steady-state conditions are investigated for the impact force calculation. Under steady-state conditions with a load torque 134.09N m applied to the gear (equivalent to applying a 50N m input torque on the pinion) and a rotational speed 2000 r/min given to the pinion, the simulation results of the FE model are successfully obtained. Due to space limitation, the primary results at a point during the analysis are shown in Figure 5. As shown in Figure 5(a), the pinion correctly engaged with the gear due to their interaction. This outcome demonstrates that the boundary conditions, the load, and the contact relationship defined in the model are appropriate. In addition, the clear contact traces of the pinion and gear are acquired as shown in Figure 5(b) and (c). This outcome indicates that the structured mesh models of the teeth are valid to guarantee the computational accuracy. In sum, the FEM model is acceptable to obtain the results of the meshing impact force. The basic mechanism of meshing impact out of the transverse contact path is illustrated in Figure 6. First, there are two tooth pairs engaging correctly along the transverse contact path while another impact tooth pair is about to engage (Figure 6(a)). Then, as the pinion rotates, the impact tooth pair contacts at the initial impact point (Figure 6(b))" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000793_17452759.2010.527010-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000793_17452759.2010.527010-Figure2-1.png", + "caption": "Figure 2. Definition of cusp height in an error triangle.", + "texts": [ + " For this reason, the number of possible layer thicknesses should be limited to, e.g. four, which leads to four different parameter sets (laser power, scan speed, hatching space, etc.). To emphasize the advantage of a simple algorithm compared with uniform slicing, an illustrative part will be sliced with the cusp height concept. 3.1 Methodology for error quantification Cusp height concept is one of the first approaches for adaptive slicing and still one of the most effective ones. Hence, it will be adopted in this paper. Figure 2 illustrates the definition of cusp height in an error triangle. According to Figure 2 cusp height c is given by the layer thickness t and the inclination angle g between the corresponding facet and the vertical (\u00a7 build direction): c t sin g (1) By observing the normal vector n of a facet, it is also obvious that sin g nz \u00bdn\u00bd ; n (nx; ny; nz): (2) Taking into account a predefined maximum value for an allowable deviation cmax, cusp height vector c has to fulfil the following inequation: \u00bdc\u00bd c5cmaxBt; c cn (3) D ow nl oa de d by [ D al ho us ie U ni ve rs ity ] at 0 6: 43 3 1 D ec em be r 20 14 Equations (1) (3) can be used to quantify and evaluate the staircase error when uniformly slicing a part", + " 2009, Mangan and Whitaker 1999) it is possible to define separate threshold values for allowable deviation or rather allowable cusp height cmax. To evaluate surface quality with regard to the affecting stair stepping effect Equation (3) is checked for validity for each facet in a layer. As a rule, for an arbitrary part orientation this results in maybe completely different cusp height values. For this reason the upper mentioned approach is extended to additional parameters for quantification. Therefore, an average cusp height cL is determined with the help of the number of faces f: cL 1 f Xf 1 i 0 ci (6) According to Figure 2 the occurring cusp height c is at most equal to layer thickness t. Considering the overall existent maximum cusp height cocc,max, an evaluation factor rSurf for surface quality characterisation is given as rSurf c (cocc;max c)2 t ; c 1 n Xn 1 i 0 cL;i; cocc;max]c (7) where n denotes the number of layers. Including the orientation of the facet normal vector (nz positive or negative) one can define different threshold values for shape deviation with different inclusion properties of the layer model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001432_iros.2011.6094612-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001432_iros.2011.6094612-Figure3-1.png", + "caption": "Fig. 3. The tested IMU comprised of two tri-axis accelerometers separated by 120mm and one uni-axis rate gyro.", + "texts": [ + " The constants A, B, and C in (15) are obtained by a least-squares fitting of a quadratic function to the static solution variance data for dynamic magnitudes up to and including the threshold value. \u03c32 \u03b8 = { Ad2 +Bd+ C if d < dThreshold \u03c32 dynamic if d \u2265 dThreshold (15) The sensor fusion technique discussed above was tested on an IMU constructed with two ADXL-325 \u00b1 3g tri-axis accelerometers and one ADXRS610 \u00b1 300deg/sec rate gyro from Analog Devices1. The two accelerometers are mounted 120 mm apart with the rate gyro mounted in the middle. The device is shown in Fig. 3. The IMU is mounted to a Quanser2 SRV-02 rotational link servo and subjected to a number of trajectories to determine the characteristics of the IMU and filter methodology. A series of constant angular velocity inputs and sinusoidal inputs are given to the SRV-02 hardware under close loop control. The inclination 1www.analog.com 2www.quanser.com estimate from the IMU and filter is compared to the encoder angle measurement from the SRV-02 to characterize the performance of the algorithm discussed above" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002077_ffe.12654-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002077_ffe.12654-Figure9-1.png", + "caption": "FIGURE 9 A, The critical section of the axle pin with the position and siz region", + "texts": [ + "2 The structural integrity assessment procedure is based on 3 sets of input data. The first set represents the mechanical properties and the fracture toughness of the cast material, the second set consists of the loading distribution, and the third set is the position and size of the flaws. During inspection, the defects are detected within the region of the geometrical change of the axle pin. The position and size of the detected crack in the critical section of the axle pin are schematically shown in Figure 9A. The idealized unique surface crack in the most critical region, due to conservatism, can be considered by the FITNET2 procedure, as shown in Figure 9B. Per the FITNET2 procedure, it is allowable to assume and perform the assessment for the worst\u2010case crack geometry. This conservative approach leads to conservative results. An idealization of the crack geometry is shown in Figure 9B, where a = 20 mm is the crack depth and 150\u00b0 is the radial surface crack length. The wind\u2010power turbine's axle pin is loaded by bending, as 2 bearings are mounted onto the pin, shown schematically in Figure 1. Therefore, the bearings e of the crack. B, The idealized unique surface crack in the most critical on the pin ensure rotation of the hub with a direct drive to the direct current generator, while on the axle pin, the stator generator was fixed, as shown in Figure 1. The stress intensity factor is calculated via Equation 2 below: KI \u00bc \u03c3\u22c5 ffiffiffiffiffiffiffi \u03c0\u22c5a p \u22c5Y a T ; \u03b2 ; (2) where T is the thickness of the axle pin, Y(a/T,\u03b2) is the stress intensity function, a/T is the crack depth ratio, \u03b2 is the surface crack length in radians, and \u03c3 is the principal opening load, obtained by FE modelling" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000009_cae.20327-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000009_cae.20327-Figure1-1.png", + "caption": "Figure 1 Elements of the oscillating group.", + "texts": [ + " The paper is structured as follows: second section gives an explanation of the elements of the oscillating group of the washing machine and of the assumptions made to define the mathematical model that has been used to analyze its movement; third section details the structure of the software, and fourth section describes the simulation conducted by the students. Finally, the advantages of this simulation software are revealed in fifth section. The elements of the oscillating group of a horizontal axis washing machine are shown in Figure 1. It is formed by a plastic tub, a non-rotating component that contains a steel drum where the laundry is placed. Both elements are joined through two bearings that enable the drum to turn; thanks to a pulley belt system driven by a motor joined to the lower part of the tub. There are a number of concrete counterweights attached to the tub for balancing the group. Finally, a metallic frame or cabinet is the support where the tub is joined by two springs at its upper part and by two telescopic friction dampers at the bottom" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002933_icelmach.2018.8507125-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002933_icelmach.2018.8507125-Figure2-1.png", + "caption": "Fig. 2. Electromagnetic model", + "texts": [ + " 2 F F F k F F F k F F F k \u03bd \u03bd \u03bd \u03bd \u03bd \u03bd \u03bd \u03bd \u03bd \u03bd \u03bd \u03bd \u03a6 \u03a6 \u03a6 = = \u2212 = + = = = = + = = = = \u2212 = (16) Thus, we can come to a preliminary judgment that the amplitudes of (2\u03bdp,0f) and (0,2f) decrease after MMF phase compensation according to (12)-(14). The modification of harmonic contents under MMF phase compensation will be analyzed in the following section. IV. FEA OF RADIAL FORCE DENSITY Finite element analysis is an important way for machine analysis and optimization. This section analyzes the radial force density distributions and harmonic contents under different conditions, including normal, one-phase open and MMF phase compensation. The electromagnetic model of a 16p24s modular PMSM is built and illustrated in Fig. 2. Here, we still assume that phase A in Unit 1 (hereafter called phase A1) is suffering from OC fault. The simulation results of air-gap flux density and harmonic contents are shown in Fig. 3. It can be clearly seen that the flat-top value of the radial flux density in the faulty Unit 1 decreases in fault condition. OC fault in one unit causes asymmetrical field distribution in the whole machine and thus results in a few subharmonics. After MMF phase compensation, the amplitudes of the main harmonic contents slightly increase, and the amplitudes of subharmonics are also lower than those before compensation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003183_0954406219843954-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003183_0954406219843954-Figure3-1.png", + "caption": "Figure 3. Definition of misalignments.", + "texts": [ + " In this arrangement, the electrical drive unit solely injects the power losses induced by the components of the hypoid back-to-back test rig. The test pinion is fixed with tapered roller bearings by a pot construction in a mechanically rigid gearbox. This pot construction allows an axial adjustment of the pinion. The wheel can be adjusted in axial direction as well. Thus, the contact pattern and the backlash of the test gear set can be adjusted optimally and a variation of the J- and H-values is possible. The definition of the misalignments of a bevel gear set is shown exemplary in Figure 3. The load application is realized by twisting both halves of the load clutch against each other and fixing them with several screws (4). The load level is measured by strain gauges that are mounted on the torsional shaft (5). Prior to the test run, the measurement technology for the load torque is calibrated by the use of predefined standard weights. The lubrication of the test gearbox is designed as oil dip lubrication. The oil level is adjusted to about 10mm below the wheel axle. The mineral oil \u2018\u2018FVA 3\u2019\u2019 was used with different contents (0 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003882_s11071-019-05343-5-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003882_s11071-019-05343-5-Figure5-1.png", + "caption": "Fig. 5 Schematic diagram of forces acting on outer ring", + "texts": [ + " The differential equation [24] for the vibration of the inner ring can be described as Fix + N\u2211 j=1 [(F\u03b7i j \u2212 T\u03b7i j ) cos\u03b1i j \u2212 Qi j sin \u03b1i j ] = mi x\u0308i Fiy + N\u2211 j=1 [(T\u03be i j \u2212 F\u03be i j ) cos\u03d5i j \u2212 (T\u03b7i j \u2212 F\u03b7i j ) sin \u03b1i j sin \u03d5i j +Qi j cos\u03b1i j sin \u03d5i j ] = mi y\u0308i Fiz + N\u2211 j=1 [(T\u03be i j \u2212 F\u03be i j ) sin \u03d5i j + (T\u03b7i j \u2212 F\u03b7i j ) sin \u03b1i j cos\u03d5i j \u2212 Qi j cos\u03b1i j cos\u03d5i j ] = mi z\u0308i Mix + N\u2211 j=1 (F\u03be i j \u2212 T\u03be i j )ri j = Iix \u03c9\u0307ix \u2212 (Iiy \u2212 Iiz)\u03c9iy\u03c9iz Miy + N\u2211 j=1 [Qi j sin \u03b1i j + (T\u03b7i j \u2212 F\u03b7i j ) cos\u03b1i j ]ri j cos\u03d5i j \u2212 N\u2211 j=1 [ (T\u03be i j \u2212 F\u03be i j ) DW 2 ki sin \u03b1i j sin \u03d5i j ] = Iiy\u03c9\u0307iy \u2212 (Iiz \u2212 Iix )\u03c9iz\u03c9ix Miz + N\u2211 j=1 [Qi j sin \u03b1i j + (T\u03b7i j \u2212 F\u03b7i j ) cos\u03b1i j ]ri j sin \u03d5i j + N\u2211 j=1 [ (T\u03be i j \u2212 F\u03be i j ) DW 2 ki sin \u03b1i j cos\u03d5i j ] = Iiz\u03c9\u0307iz \u2212 (Iix \u2212 Iiy)\u03c9ix\u03c9iy (3) where mi is the mass of the inner ring; Fix , Fiy , and Fiz are the external loads acting on the inner ring; Mix , Miy , and Miz are the external torques acting on the inner ring; x\u0308i, y\u0308i, and z\u0308i are the displacement accelerations of the barycentre of the inner ring along the direction of each coordinate axis in the coordinate system {O; X,Y, Z};\u03c9ix , \u03c9iy, and \u03c9iz are the angular velocities of the inner ring in the coordinate system {O; X,Y, Z}; \u03c9\u0307ix , \u03c9\u0307iy , and \u03c9\u0307iz are the angular accelerations of the inner ring in the coordinate system {O; X,Y, Z}; Iix , Iiy , and Iiz are the moments of inertia of the inner ring in the coordinate system {O; X,Y, Z}; ki is the inner ring raceway curvature radius coefficient; and ri j is the raceway radius of the inner ring and can be calculated by ri j = dm 2 \u2212 DW 2 ki cos\u03b1i j (4) 2.4 Differential equations of vibrations on outer ring Since the outer ring of the bearing is installed inside the bearing pedestal and does not rotate with the rotating inner ring, the outer ring is subject to the force of the bearing pedestal. The forces acting on the outer ring comprise the contact force and dynamic friction forces caused by the interaction between outer ring and balls. The forces of the j th ceramic ball on the outer ring of the bearing are shown in Fig. 5. In Fig. 5, \u03c6o j is the position angle of the j th ceramic ball relative to the outer ring. The differential equation for the vibration of the outer ring can be described as Fox + N\u2211 j=1 [(T\u03b7o j \u2212 F\u03b7o j ) cos\u03b1o j + Qo j sin \u03b1o j ] = mo x\u0308o Foy \u2212 N\u2211 j=1 [ (T\u03beo j \u2212 F\u03beo j ) cos\u03d5o j \u2212 (T\u03b7o j \u2212 F\u03b7o j ) sin \u03b1o j sin \u03d5o j + Qo j cos\u03b1o j sin \u03d5o j ] = mo y\u0308o Foz \u2212 N\u2211 j=1 [( T\u03beo j \u2212 F\u03beo j ) sin \u03d5o j + ( T\u03b7o j \u2212 F\u03b7o j ) sin \u03b1o j cos\u03d5o j \u2212 Qo j cos\u03b1o j cos\u03d5o j ] = mo z\u0308o Mox + N\u2211 j=1 (T\u03beo j \u2212 F\u03beo j )ro j = Iox \u03c9\u0307ox \u2212 ( Ioy \u2212 Ioz) \u03c9oy\u03c9oz Moy + N\u2211 j=1 [ Qo j sin \u03b1o j + ( T\u03b7o j \u2212 F\u03b7o j ) cos\u03b1o j ] ro j cos\u03d5o j + N\u2211 j=1 [ (T\u03beo j \u2212 F\u03beo j ) DW 2 ko sin \u03b1o j sin \u03d5o j ] = Ioy\u03c9\u0307oy \u2212 (Ioz \u2212 Iox )\u03c9oz\u03c9ox Moz + N\u2211 j=1 [ Qo j sin \u03b1o j + ( T\u03b7o j \u2212 F\u03b7o j ) cos\u03b1o j ] ro j sin \u03d5o j \u2212 N\u2211 j=1 [ (T\u03beo j \u2212 F\u03beo j ) DW 2 ko sin \u03b1o j cos\u03d5o j ] = Ioz\u03c9\u0307oz \u2212 (Iox \u2212 Ioy)\u03c9ox\u03c9oy (5) where mo is the mass of the outer ring; Fox , Foy , and Foz are the external loads acting on the outer ring; and Mox , Moy , and Moz are the external torques acting on the outer ring" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003174_s12008-019-00545-y-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003174_s12008-019-00545-y-Figure1-1.png", + "caption": "Fig. 1 Schematic diagram of a scotch yoke mechanism", + "texts": [ + " As the scotch yoke mechanism involves rolling-sliding contact and fluctuated load, it is vital to maintain an efficient way of lubrication for the moving elements. In this paper, elastohydrodynamic lubrication (EHL) analysis is carried out on this mechanism to analyze its dynamics under this light. In this section, the velocity of the contacting surfaces relative to the point of contact is determined plus the load between them. The dynamic analysis is developed over the cycle of operation of the mechanism. Accordingly, particular positions are selected for the EHL analysis. Figure\u00a01 shows a schematic diagram of a scotch yoke mechanism; the velocity of the contacting surfaces is determined as following: The velocity of a point on the pin adjacent to the point of contact in the x and y directions are, respectively: (v1)x = Rp?\u0307? sin (\ud835\udefc + \ud835\udf03) (v1)y = Rp?\u0307? cos (\ud835\udefc + \ud835\udf03) * Oday I. Abdullah oday.abdullah@tuhh.de 1 Mechanical Engineering Department, University of\u00a0Kerbala, Kerbala, Iraq 2 Energy Engineering Department, University of\u00a0Baghdad, Baghdad, Iraq 3 Hamburg University of\u00a0Technology, Hamburg, Germany 4 Department of\u00a0Industrial Engineering, Nanomates Centre, University of\u00a0Salerno, Fisciano, Italy 1 3 where Rp is the radial distance between the point of contact and the center of rotation of the rotating disk and " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002926_0142331218788115-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002926_0142331218788115-Figure1-1.png", + "caption": "Figure 1. The quadrotor system.", + "texts": [ + ", 2014) \u20acxi =(cosuisinuicosci +sinuisinci) ui4 mi +Di1, 1 \u20acyi =(cosuisinuisinci sinuicosci) ui4 mi +Di1, 2 \u20aczi =cosui cosui ui4 mi g +Di3, 2 \u20acui = bi1ui1 +Di2, 1 \u20acui = bi2ui2 +Di2, 2 \u20acci = bi3ui3 +Di3, 1, \u00f05\u00de where (xi, yi, zi) T is the mass center position of the ith quadrotor system, (fi, ui,ci) T is the roll-pitch-yaw of the ith quadrotor system, mi is the mass of the ith quadrotor system, g is gravity acceleration, bi1 = li 2 ffiffi 2 p Iixx , bi2 = li 2 ffiffi 2 p Iiyy , bi3 = li Iizz , li is the length of the ith quadrotor frame, and Iixx, Iiyy, Iizz are the total inertia moment of x-, y-, z-axis of the ith quadrotor system, respectively. Obviously, from Figure 1 and equation (5), the control inputs uij, (j=1, 2, 3, 4) will be the combination of fij, (j=1, 2, 3, 4), where fij, (j=1, 2, 3, 4) is the lifting force of the jth rotor of the ith quadrotor. Therefore, we have control inputs of the ith quadrotor as ui1 = kip(fi1 fi2 fi3 + fi4), ui2 = kip(fi1 + fi2 fi3 fi4), and ui3 = kid(fi1 fi2 + fi3 fi4) for the roll, pitch and yaw, respectively. kip is a thrust coefficient of the ith quadrotor, kid is a drag coefficient of the ith quadrotor and the control input of the ith quadrotor for the altitude is with the form of ui4 = kip(fi1 + fi2 + fi3 + fi4). Dij, l, (j= 1, 2, 3, l= 1, 2) is the bounded uncertainties of the ith quadrotor, consisting of unknown modelling errors, parameter uncertainties, and/or external disturbances. The ith quadrotor system is shown in Figure 1. Assumption 2: Bounded uncertainties: There exists a positive constant kD such that jjDij, ljj\u0142 kD, for i= 1, 2, . . . , n 1, j= 1, 2, 3, l = 1, 2. Let xi1 = \u00bdxiyi T , xi2 = \u00bd _xi _yi T , xi3 = \u00bduiui T , xi4 = \u00bd _ui _ui T , xi5 = \u00bdcizi T , and xi6 = \u00bd _ci _zi T . _xi1 =xi2 _xi2 =Gi0(xi5, ui4)ui(xi3)+Di1 _xi3 =xi4 _xi4 =Gi1ui, 12 +Di2 _xi5 =xi6 _xi6 = gi1 +Gi2(xi3)ui, 34 +Di3 \u00f06\u00de where Gi0(xi5, ui4)= ui4 mi sinci cosci cosci sinci , u(xi3)= sinui cosui sin ui , Gi1 = bi1 0 0 bi2 , ui, 12 = ui1 ui2 , gi1 = 0 g , ui, 34 = ui3 ui4 , Gi2(xi3)= bi3 0 0 1 mi cosui cos ui , Dij = Dij, 1 Dij, 2 , i= 1, 2, " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001750_ieejias.132.426-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001750_ieejias.132.426-Figure1-1.png", + "caption": "Fig. 1. Relationship between d\u2013q and \u03b3\u2013\u03b4 reference frames", + "texts": [], + "surrounding_texts": [ + "\u96fb\u6c17\u5b66\u4f1a\u8ad6\u6587\u8a8c D\uff08\u7523\u696d\u5fdc\u7528\u90e8\u9580\u8a8c\uff09 IEEJ Transactions on Industry Applications Vol.132 No.3 pp.426\u2013436 DOI: 10.1541/ieejias.132.426\n\u8ad6 \u6587\n\u30e2\u30c7\u30eb\u898f\u7bc4\u9069\u5fdc\u30b7\u30b9\u30c6\u30e0\u306b\u57fa\u3065\u304f\u6c38\u4e45\u78c1\u77f3\u30e2\u30fc\u30bf\u306e \u78c1\u6975\u4f4d\u7f6e\u30bb\u30f3\u30b5\u30ec\u30b9\u5236\u5fa1\u3068\u30d1\u30e9\u30e1\u30fc\u30bf\u611f\u5ea6\n\u6b63 \u54e1 \u5c0f\u539f \u6b63\u6a39\u2217,\u2217\u2217 \u6b63 \u54e1 \u91ce\u53e3 \u5b63\u5f66\u2217\nRotor Position Sensorless Control and Its Parameter Sensitivity of Permanent Magnet Motor Based on Model Reference Adaptive System\nMasaki Ohara\u2217,\u2217\u2217, Member, Toshihiko Noguchi\u2217, Member\n\uff082011\u5e746\u67083\u65e5\u53d7\u4ed8\uff0c2011\u5e7410\u67083\u65e5\u518d\u53d7\u4ed8\uff09\nThis paper describes a new method for a rotor position sensorless control of a surface permanent magnet synchronous motor based on a model reference adaptive system (MRAS). This method features the MRAS in a current control loop to estimate a rotor speed and position by using only current sensors. This method as well as almost all the conventional methods incorporates a mathematical model of the motor, which consists of parameters such as winding resistances, inductances, and an induced voltage constant. Hence, the important thing is to investigate how the deviation of these parameters affects the estimated rotor position. First, this paper proposes a structure of the sensorless control applied in the current control loop. Next, it proves the stability of the proposed method when motor parameters deviate from the nominal values, and derives the relationship between the estimated position and the deviation of the parameters in a steady state. Finally, some experimental results are presented to show performance and effectiveness of the proposed method.\n\u30ad\u30fc\u30ef\u30fc\u30c9\uff1a\u30e2\u30c7\u30eb\u898f\u7bc4\u9069\u5fdc\u30b7\u30b9\u30c6\u30e0\uff0c\u6c38\u4e45\u78c1\u77f3\u540c\u671f\u30e2\u30fc\u30bf\uff0c\u78c1\u6975\u4f4d\u7f6e\u30bb\u30f3\u30b5\u30ec\u30b9\u5236\u5fa1\uff0c\u30d1\u30e9\u30e1\u30fc\u30bf\u611f\u5ea6 Keywords: model reference adaptive system, permanent magnet synchronous motor, rotor position sensorless control, parameters sensitivity\n1. \u306f\u3058\u3081\u306b\n\u6628\u4eca\u306e\u6c38\u4e45\u78c1\u77f3\u540c\u671f\u30e2\u30fc\u30bf\uff08\u4ee5\u4e0b PMSM\uff09\u306f\u5c0f\u5f62\uff0c\u9ad8\u52b9 \u7387\uff0c\u5bb9\u91cf\u62e1\u5927\u306b\u3088\u308a\uff0c\u305d\u306e\u9069\u7528\u7bc4\u56f2\u306f\u7523\u696d\u5206\u91ce\u3060\u3051\u3067\u306a\u304f 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\u30b9\u5236\u5fa1\u306e\u65b9\u5f0f\u3068\u3057\u3066\uff0c\u63a8\u5b9a\u56de\u8ee2\u5ea7\u6a19\u3067\u8868\u3055\u308c\u308b\u8a98\u8d77\u96fb\u5727\u6210 \u5206\u304b\u3089\u6f14\u7b97\u3067\u6c42\u3081\u308b\u65b9\u5f0f\uff0c\u78c1\u675f\u3092\u63a8\u5b9a\u3059\u308b\u305f\u3081\u56fa\u5b9a\u5b50\u5ea7\u6a19\n\u2217 \u9759\u5ca1\u5927\u5b66 \u3012432-8561 \u6d5c\u677e\u5e02\u4e2d\u533a\u57ce\u5317 3-5-1 Shizuoka University 3-5-1, Johoku, Naka-ku, Hamamatsu 432-8561, Japan\n\u2217\u2217\uff08\u682a\uff09\u7af9\u4e2d\u88fd\u4f5c\u6240 \u3012578-0984 \u6771\u5927\u962a\u5e02\u83f1\u6c5f 6-4-35 Takenaka Seisakusho Co., Ltd. 6-4-35, Hishie, Higashi-Osaka 578-0984, Japan\n\u3042\u308b\u3044\u306f\u63a8\u5b9a\u56de\u8ee2\u5ea7\u6a19\u3067\u56db\u6b21\u5143\u306e\u9069\u5fdc\u30aa\u30d6\u30b6\u30fc\u30d0\u3092\u69cb\u6210\u3057\uff0c \u5f97\u3089\u308c\u305f\u78c1\u675f\u304b\u3089\u78c1\u6975\u4f4d\u7f6e\u30fb\u901f\u5ea6\u3092\u63a8\u5b9a\u3059\u308b\u65b9\u5f0f\uff0c\u30e2\u30fc\u30bf 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\u307e\u305f\uff0c\u5b89\u5b9a\u6027\u3084\u30d1\u30e9\u30e1\u30fc\u30bf\u5909\u52d5\u306b\u95a2\u3059\u308b\u89e3\u6790\u3082\u4e00\u90e8\u793a\u3055\u308c \u3066\u3044\u308b\u304c\u5341\u5206\u3067\u3042\u308b\u3068\u306f\u8a00\u3048\u306a\u3044\u3002\u7279\u306b\u8fd1\u5e74\u306e\u30bb\u30f3\u30b5\u30ec\u30b9 \u5236\u5fa1\u3078\u306e\u9ad8\u6027\u80fd\u5316\u8981\u6c42\u3092\u6e80\u8db3\u3059\u308b\u305f\u3081\u306b\u306f\uff0c\u30d1\u30e9\u30e1\u30fc\u30bf\u5909 \u52d5\u6642\u306e\u5236\u5fa1\u306e\u5b89\u5b9a\u6027\u3084\u78c1\u6975\u4f4d\u7f6e\u63a8\u5b9a\u5024\u306b\u5bfe\u3059\u308b\u5f71\u97ff\u3092\u660e\u78ba \u306b\u3057\uff0c\u4f55\u3089\u304b\u306e\u5bfe\u7b56\u624b\u6bb5\u3092\u7528\u3044\u3066\u30d1\u30e9\u30e1\u30fc\u30bf\u5909\u52d5\u306e\u5f71\u97ff\u3092\u6975 \u529b\u5c11\u306a\u304f\u3059\u308b\u3053\u3068\u304c\u91cd\u8981\u306a\u8ab2\u984c\u3067\u3042\u308b\u3002\u3053\u306e\u305f\u3081\uff0c\u5404\u7a2e\u30d1 \u30e9\u30e1\u30fc\u30bf\u3092\u505c\u6b62\u4e2d\u53ca\u3073\u904b\u8ee2\u4e2d\u306b\u8a08\u6e2c\u3059\u308b\u65b9\u5f0f\uff0c\u30a4\u30f3\u30c0\u30af\u30bf\u30f3 \u30b9\u5909\u52d5\u306b\u30ed\u30d0\u30b9\u30c8\u306a\u5236\u5fa1\u65b9\u5f0f\u306a\u3069\u306e\u7814\u7a76\u304c\u3055\u308c\u3066\u3044\u308b (11) (12)\u3002 \u3057\u304b\u3057\u306a\u304c\u3089\uff0c\u524d\u8005\u306e\u65b9\u6cd5\u3067\u306f\u672c\u6765\u306e\u76ee\u7684\u306e\u30bb\u30f3\u30b5\u30ec\u30b9\u5236 \u5fa1\u4ee5\u5916\u306b\u30d1\u30e9\u30e1\u30fc\u30bf\u3060\u3051\u3092\u63a8\u5b9a\u3059\u308b\u624b\u6bb5\u304c\u5225\u306b\u5fc5\u8981\u3068\u306a\u308a\uff0c \u5f8c\u8005\u306e\u65b9\u6cd5\u3067\u3082\u30a4\u30f3\u30c0\u30af\u30bf\u30f3\u30b9\u4ee5\u5916\u306e\u30d1\u30e9\u30e1\u30fc\u30bf\u5909\u52d5\u306b\u306f\nc\u00a9 2012 The Institute of Electrical Engineers of Japan. 426", + "\u5225\u306e\u88dc\u511f\u624b\u6bb5\u304c\u5fc5\u8981\u3068\u306a\u308b\u305f\u3081\uff0c\u5236\u5fa1\u88c5\u7f6e\u306e\u69cb\u6210\u304c\u8907\u96d1\u306b \u306a\u308b\u8ab2\u984c\u304c\u5b58\u5728\u3059\u308b\u3002 \u4ee5\u4e0a\u306e\u8b70\u8ad6\u304b\u3089\uff0c\u7b46\u8005\u3089\u306f\u30d1\u30e9\u30e1\u30fc\u30bf\u540c\u5b9a\u306b\u3082\u5fdc\u7528\u53ef\u80fd \u306a\u30e2\u30c7\u30eb\u898f\u7bc4\u9069\u5fdc\u30b7\u30b9\u30c6\u30e0\uff08\u4ee5\u4e0bMRAS\uff09\u3092\u7528\u3044\u305f\u8868\u9762\u78c1 \u77f3\u5f62\u6c38\u4e45\u78c1\u77f3\u540c\u671f\u30e2\u30fc\u30bf\uff08\u4ee5\u4e0b SPMSM\uff09\u306e\u78c1\u6975\u4f4d\u7f6e\u30bb\u30f3\u30b5 \u30ec\u30b9\u5236\u5fa1\u6cd5\u3092\u63d0\u6848\u3057\uff0c\u305d\u306e\u6709\u52b9\u6027\u3092\u30b7\u30df\u30e5\u30ec\u30fc\u30b7\u30e7\u30f3\u304a\u3088 \u3073\u5b9f\u9a13\u306b\u3088\u3063\u3066\u78ba\u8a8d\u3057\u3066\u3044\u308b (13) (14)\u3002\u3053\u306e\u65b9\u5f0f\u306fMRAS\u306e \u4e00\u65b9\u5f0f\u3067\u3042\u308b\u30d1\u30e9\u30e1\u30fc\u30bf\u540c\u5b9a\u3092\u5229\u7528\u3057\u3066\u304a\u308a\uff0c\u78c1\u6975\u4f4d\u7f6e\u30fb\u901f \u5ea6\u3092\u30d1\u30e9\u30e1\u30fc\u30bf\u5909\u52d5\u3068\u898b\u306a\u3057\u5404\u3005\u3092\u63a8\u5b9a\u3057\u3066\u3044\u308b\u3002\u3053\u306e\u305f \u3081\uff0c\u30e2\u30fc\u30bf\u30d1\u30e9\u30e1\u30fc\u30bf\u5909\u52d5\u6642\u306e\u30d1\u30e9\u30e1\u30fc\u30bf\u63a8\u5b9a\u88dc\u511f\u3092\u5bb9\u6613\u306b \u8ffd\u52a0\u62e1\u5f35\u3067\u304d\u308b\u3053\u3068\u304c\u7279\u9577\u3067\u3042\u308b\u3002\u307e\u305f\uff0c\u65e2\u63d0\u6848\u306e\u30bb\u30f3\u30b5 \u30ec\u30b9\u5236\u5fa1\u6cd5\u306b\u304a\u3044\u3066\u3082\u7279\u306b\u4f4e\u901f\u57df\u3067\u306e\u9ad8\u6027\u80fd\u5316\u306e\u5b9f\u73fe\u306b\u306f\uff0c \u30d1\u30e9\u30e1\u30fc\u30bf\u5909\u52d5\u88dc\u511f\u304c\u6709\u52b9\u306a\u624b\u6bb5\u3067\u3042\u308b\u305f\u3081\uff0c\u3042\u3089\u304b\u3058\u3081 \u30d1\u30e9\u30e1\u30fc\u30bf\u5909\u52d5\u6642\u306e\u5236\u5fa1\u7cfb\u306e\u6319\u52d5\u3092\u8abf\u3079\u308b\u5fc5\u8981\u304c\u3042\u308b\u3002\u672c \u8ad6\u6587\u3067\u306f\uff0c\u30d1\u30e9\u30e1\u30fc\u30bf\u63a8\u5b9a\u88dc\u511f\u3092\u9069\u7528\u3059\u308b\u306b\u5f53\u305f\u308a\uff0cMRAS \u3067\u306e\u30bb\u30f3\u30b5\u30ec\u30b9\u5236\u5fa1\u306e\u5b89\u5b9a\u6027\u3068\u78c1\u6975\u4f4d\u7f6e\u63a8\u5b9a\u8aa4\u5dee\u306e\u89e3\u6790\u304a \u3088\u3073\u305d\u306e\u59a5\u5f53\u6027\u306e\u5b9f\u9a13\u691c\u8a3c\u3092\u4e2d\u5fc3\u306b\u5831\u544a\u3059\u308b\u3002\u306f\u3058\u3081\u306b\u96fb \u6d41\u5236\u5fa1\u30eb\u30fc\u30d7\u5185\u306bMRAS\u3092\u7528\u3044\u305f\u30bb\u30f3\u30b5\u30ec\u30b9\u5236\u5fa1\u306e\u69cb\u6210 \u3092\u8ff0\u3079\u308b\u3002\u3053\u306e\u5834\u5408\uff0c\u898f\u7bc4\u30e2\u30c7\u30eb\u306b\u7406\u60f3\u30e2\u30c7\u30eb\uff08SPMSM \u306e d-q\u56de\u8ee2\u5ea7\u6a19\u306b\u304a\u3051\u308b\u96fb\u5727\u96fb\u6d41\u65b9\u7a0b\u5f0f\u3067\uff0c\u5404\u8ef8\u306e\u96fb\u5727\u3068 \u96fb\u6d41\u306e\u95a2\u4fc2\u304c\u4e00\u6b21\u9045\u308c\u306b\u306a\u308b\uff09\u3092\u63a1\u7528\u3057\u3066\u4e26\u5217\u5f62MRAS\u3092 \u69cb\u6210\u3059\u308b\u3068\uff0c\u7d50\u679c\u7684\u306b\u96fb\u6d41\u5236\u5fa1\u7cfb\u304c\u975e\u5e72\u6e09\u5316\u3055\u308c\u305f\u30bb\u30f3\u30b5 \u30ec\u30b9\u5236\u5fa1\u3068\u306a\u308b\u3002\u6b21\u306b\u5404\u7a2e\u306e\u30d1\u30e9\u30e1\u30fc\u30bf\u5909\u52d5\u306b\u5bfe\u3059\u308b\u30bb\u30f3 \u30b5\u30ec\u30b9\u5236\u5fa1\u7cfb\u306e\u5b89\u5b9a\u6027\u3068\u78c1\u6975\u4f4d\u7f6e\u306e\u63a8\u5b9a\u7279\u6027\u306b\u3064\u3044\u3066\u8ff0\u3079 \u308b\u3002\u6700\u5f8c\u306b\u30d1\u30e9\u30e1\u30fc\u30bf\u5909\u52d5\u6642\u306e\u5b9a\u5e38\u7279\u6027\u3068\u904e\u5ea6\u7279\u6027\u3092\u5b9f\u6a5f \u5b9f\u9a13\u306b\u3088\u308a\u793a\u3057\uff0c\u63d0\u6848\u6cd5\u306e\u30d1\u30e9\u30e1\u30fc\u30bf\u5909\u52d5\u6642\u306b\u304a\u3051\u308b\u5b89\u5b9a \u6027\u3068\u78c1\u6975\u4f4d\u7f6e\u63a8\u5b9a\u8aa4\u5dee\u306e\u6319\u52d5\u304c\u89e3\u6790\u3068\u4e00\u81f4\u3059\u308b\u3053\u3068\u3092\u691c\u8a3c \u3059\u308b\u3002\n2. MRAS\u306e\u69cb\u6210\n\u78c1\u6975\u4f4d\u7f6e\u3092 d \u8ef8\u3068\u3057\u305f\u56de\u8ee2\u5ea7\u6a19\u8ef8\uff08d-q\u8ef8\uff09\u306b\u304a\u3051\u308b\u96fb \u6d41\u5236\u5fa1\u30eb\u30fc\u30d7\u5185\u3067\uff0cMRAS\u3092\u7528\u3044\u3066\u975e\u5e72\u6e09\u30bb\u30f3\u30b5\u30ec\u30b9\u5236\u5fa1 \u3092\u69cb\u6210\u3059\u308b\u65b9\u6cd5\u306b\u3064\u3044\u3066\u8ff0\u3079\u308b\u3002\u63a8\u5b9a\u78c1\u6975\u8ef8\u3092 \u03b3\u8ef8\u3068\u3057\u305f \u03b3-\u03b4\u5ea7\u6a19\u3068 d-q\u5ea7\u6a19\u3068\u306e\u95a2\u4fc2\u3092 Fig. 1\u306b\u793a\u3059\u3002\n\u30082\u30fb1\u3009 \u03b3-\u03b4\u5ea7\u6a19\u306b\u304a\u3051\u308b\u72b6\u614b\u65b9\u7a0b\u5f0f SPMSM\u306e \u03b3-\u03b4\n\u63a8\u5b9a\u56de\u8ee2\u5ea7\u6a19\u306b\u304a\u3051\u308b\u96fb\u5727\u96fb\u6d41\u65b9\u7a0b\u5f0f\u3092 (1)\u5f0f\u306b\u793a\u3059\u3002\u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3v\u03b3v\u03b4 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 = \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3R + pL \u2212\u03c9\u0302L \u03c9\u0302L R + pL \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3i\u03b3i\u03b4 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 + \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3\u2212\u03c9\u03c6 sin(\u03b8 \u2212 \u03b8\u0302) \u03c9\u03c6 cos(\u03b8 \u2212 \u03b8\u0302) \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 (1)\n\u3053\u3053\u3067 v\u03b3\uff0cv\u03b4\uff0ci\u03b3\uff0ci\u03b4\uff1a\u63a8\u5b9a\u5ea7\u6a19\u4e0a\u306e\u96fb\u5727\uff0c\u96fb\u6d41\uff0c\u03c9\uff1a\u56de\u8ee2 \u901f\u5ea6\u771f\u5024\uff0c\u03c9\u0302\uff1a\u56de\u8ee2\u901f\u5ea6\u63a8\u5b9a\u5024\uff0c\u03c6\uff1a\u8a98\u8d77\u96fb\u5727\u5b9a\u6570\uff0c\u03b8\uff1a\u78c1\u6975 \u4f4d\u7f6e\u771f\u5024\uff0c\u03b8\u0302\uff1a\u78c1\u6975\u4f4d\u7f6e\u63a8\u5b9a\u5024\uff0cR\uff1a\u5dfb\u7dda\u62b5\u6297\uff0cL\uff1a\u540c\u671f\u30a4\u30f3 \u30c0\u30af\u30bf\u30f3\u30b9\u3067\u3042\u308b\u3002(1)\u5f0f\u3092\u5909\u5f62\u3057\u3066\u96fb\u6d41\u3092\u72b6\u614b\u5909\u6570\u304a\u3088\u3073 \u51fa\u529b\u5909\u6570\uff0c\u96fb\u5727\u3092\u5165\u529b\u5909\u6570\u3068\u3059\u308b PMSM\u306e \u03b3-\u03b4\u63a8\u5b9a\u56de\u8ee2\u5ea7 \u6a19\u306b\u304a\u3051\u308b\u72b6\u614b\u65b9\u7a0b\u5f0f\u3092\u6c42\u3081\u308b\u3068 (2)\u5f0f\u3068\u306a\u308b\u3002\nx\u0307 = Ax \u2212 \u03c9\u0302Jx + Bu \u2212 Be\u03b3\u03b4 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 (2a)\ny = Cx \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 (2b)\n\u3053\u3053\u3067\uff0cx = y = [ i\u03b3 i\u03b4 ]T \uff0cu = [ v\u03b3 v\u03b4 ]T e\u03b3\u03b4 = [ e\u03b3 e\u03b4 ]T \uff1a \u6c38\u4e45\u78c1\u77f3\u306b\u3088\u308b\u8a98\u8d77\u96fb\u5727\uff0c e\u03b3 = \u2212\u03c9\u03c6 sin(\u03b8 \u2212 \u03b8\u0302)\uff0ce\u03b4 = \u03c9\u03c6 cos(\u03b8 \u2212 \u03b8\u0302)\uff0c A = \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3\u2212R/L 0\n0 \u2212R/L\n\u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6\uff0cB = \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a31/L 0\n0 1/L\n\u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6\uff0cC = \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a31 0\n0 1 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6\uff0c J = \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a30 \u22121\n1 0 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 \u30082\u30fb2\u3009 \u56de\u8ee2\u901f\u5ea6\u3068\u78c1\u6975\u4f4d\u7f6e\u63a8\u5b9a\u5024\u306e\u5c0e\u51fa \u901f\u5ea6\u3068\u4f4d\u7f6e \u3092\u63a8\u5b9a\u3059\u308b\u305f\u3081\u306b\u4e26\u5217\u5f62\u306eMRAS\u3092\u69cb\u7bc9\u3059\u308b\u3002MRAS\u306b \u304a\u3044\u3066\u6f38\u8fd1\u5b89\u5b9a\u306a\u7cfb\u3092\u69cb\u7bc9\u3059\u308b\u305f\u3081\u306b\u306f\uff0c\u898f\u7bc4\u30e2\u30c7\u30eb\u3068\u3057 \u3066\u53b3\u5bc6\u306b\u30d7\u30ed\u30d1\u30fc\u306a\u5f37\u6b63\u5b9f\u95a2\u6570\u3067\u3042\u308b\u5fc5\u8981\u304c\u3042\u308b\u3002PMSM \u306e \u03b3-\u03b4\u56de\u8ee2\u5ea7\u6a19\u306b\u304a\u3051\u308b\u72b6\u614b\u65b9\u7a0b\u5f0f (2a)\u306b\u304a\u3044\u3066 x\u306e\u975e\u5e72 \u6e09\u9805\u3092\u9664\u3044\u305f\u5f0f\u306f\u5404\u8ef8\u306e\u5165\u529b\uff08\u96fb\u5727\u3068\u9006\u8d77\u96fb\u529b\u306e\u548c\uff09\u3068\u51fa \u529b\uff08\u96fb\u6d41\uff09\u306e\u95a2\u4fc2\u304c\u4e00\u6b21\u9045\u308c\u3068\u306a\u308b\u305f\u3081\uff0c\u3053\u306e\u5f0f\u3092\u898f\u7bc4\u30e2 \u30c7\u30eb\u3068\u3059\u308b\u3068\u53b3\u5bc6\u306b\u30d7\u30ed\u30d1\u30fc\u3067\u5f37\u6b63\u5b9f\u95a2\u6570\u3068\u306a\u308a\u6700\u9069\u3067\u3042 \u308b\u3002\u3057\u305f\u304c\u3063\u3066\uff0c\u898f\u7bc4\u30e2\u30c7\u30eb\u3092 (3)\u5f0f\u306e\u3088\u3046\u306b\u8868\u73fe\u3059\u308b\u3002\n\u02d9\u0302x = Ax\u0302 + Br \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 (3)\n\u3053\u3053\u3067\uff0cx\u0302 = [ i\u0302d i\u0302q ]T \uff1a\u30e2\u30c7\u30eb\u96fb\u6d41\uff0cr = [ v\u2217\u03b3 v\u2217\u03b4 ]T \uff1a\u96fb\u5727\u6307 \u4ee4\u3067\u3042\u308b\u3002\u6b21\u306b\uff0c\u898f\u7bc4\u30e2\u30c7\u30eb\u3068\u5b9f\u969b\u5024\u3068\u306e\u8aa4\u5dee \u03b5\u03b3\uff0c\u03b5\u03b4\u3092\u4ee5 \u4e0b\u306b\u5b9a\u7fa9\u3059\u308b\u3002\n\u03b5\u03b3 = i\u0302d \u2212 i\u03b3, \u03b5\u03b4 = i\u0302q \u2212 i\u03b4, \u03b5 = [ \u03b5\u03b3 \u03b5\u03b4 ]T \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 (4)\n(2a)\u5f0f\u3068 (3)\u5f0f\u304b\u3089\u62b5\u6297\uff0c\u30a4\u30f3\u30c0\u30af\u30bf\u30f3\u30b9\uff0c\u8a98\u8d77\u96fb\u5727\u5b9a\u6570\u306a \u3069\u306e\u30d1\u30e9\u30e1\u30fc\u30bf\u306f\u65e2\u77e5\u3068\u3057\uff0c\u56de\u8ee2\u901f\u5ea6\u306e\u307f\u3092\u672a\u77e5\u30d1\u30e9\u30e1\u30fc \u30bf\u3068\u3059\u308b\u8aa4\u5dee\u65b9\u7a0b\u5f0f\u3092\u6c42\u3081\u308b\u3002(3)\u5f0f\u304b\u3089 (2a)\u5f0f\u3092\u3072\u3044\u3066 \u6574\u7406\u3059\u308b\u3068\n\u03b5\u0307 = A\u03b5 \u2212 B(u \u2212 r \u2212 \u03c9\u0302B\u22121Jx \u2212 e\u03b3\u03b4) \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 (5)\n\u3068\u306a\u308b\u3002PMSM\u306b\u52a0\u3048\u308b\u5236\u5fa1\u5247\uff08\u96fb\u5727\u5165\u529b\uff09\u3092\nu = r + \u03c9\u0302B\u22121Jx + e\u03b3\u03b4 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 (6)\n\u3067\u4e0e\u3048\u308b\u3068 (5)\u5f0f\u306f\u6b21\u5f0f\u3068\u306a\u308a\uff0c\n\u03b5\u0307 = A\u03b5 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 (7)\n427 IEEJ Trans. IA, Vol.132, No.3, 2012", + "\u8aa4\u5dee \u03b5\u306f\u96f6\u306b\u6f38\u8fd1\u53ce\u675f\u3059\u308b\u3002\u3057\u304b\u3057\uff0c\u5b9f\u969b\u306b\u306f (6)\u5f0f\u306f\u4e0e \u3048\u3089\u308c\u306a\u3044\u306e\u3067\uff0c\u78ba\u5b9a\u7684\u7b49\u4fa1\u539f\u7406\uff08CE\u539f\u7406\uff09\u3092\u7528\u3044\u308b\u3068\uff0c \u5165\u529b\u96fb\u5727\u306f (8)\u5f0f\u3068\u306a\u308b (16)\u3002\nu = r + \u03c9\u0302B\u22121Jx + e\u0302dq \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 (8)\n\u3053\u3053\u3067\uff0ce\u0302dq = [ 0 \u03c9\u0302\u03c6 ]T \uff1a\u6c38\u4e45\u78c1\u77f3\u306b\u3088\u308b\u63a8\u5b9a\u8a98\u8d77\u96fb\u5727\n(8)\u5f0f\u3092 (5)\u5f0f\u306b\u4ee3\u5165\u3059\u308b\u3068\uff0c\u8aa4\u5dee\u65b9\u7a0b\u5f0f\u3068\u3057\u3066\u8a98\u8d77\u96fb\u5727\n\u3068\u63a8\u5b9a\u8a98\u8d77\u96fb\u5727\u3092\u542b\u3093\u3060 (9)\u5f0f\u3092\u6c42\u3081\u308b\u3053\u3068\u304c\u3067\u304d\u308b\u3002\n\u03b5\u0307 = A\u03b5 \u2212 B(e\u0302dq \u2212 e\u03b3\u03b4) \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 (9)\n(9)\u5f0f\u3067\u63a8\u5b9a\u5ea7\u6a19\u306e\u8a98\u8d77\u96fb\u5727 e\u03b3\u03b4 = [ e\u03b3 e\u03b4 ]T \u3092\u305d\u308c\u305e\u308c\u4ee5 \u4e0b\u306e\u3088\u3046\u306b\u8fd1\u4f3c\u3067\u304d\u308b\u306e\u3067\ne\u03b3 = \u2212\u03c9\u03c6 sin(\u03b8 \u2212 \u03b8\u0302) \u2212\u03c9\u03c6(\u03b8 \u2212 \u03b8\u0302) \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 (10)\ne\u03b4 = \u03c9\u03c6 cos(\u03b8 \u2212 \u03b8\u0302) \u03c9\u03c6 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 (11)\n\u3055\u3089\u306b\uff0c\u03b8 \u2212 \u03b8\u0302 \u2192 0\u306f \u03c9 \u2212 \u03c9\u0302\u2192 0\u306b\u5bfe\u3059\u308b\u5341\u5206\u6761\u4ef6\u3067\u3042\u308b \u305f\u3081\uff0c\u03c9 =\u4e00\u5b9a\u306e\u5834\u5408 |\u03c9|\u306f\u5b9a\u6570\u3068\u306a\u308a \u03c9 \u2212 \u03c9\u0302\u3092 |\u03c9|(\u03b8 \u2212 \u03b8\u0302) \u306b\u7f6e\u304d\u63db\u3048\u3066\u3082\u3053\u306e\u7cfb\u306e\u5b89\u5b9a\u6027\u306b\u306f\u5f71\u97ff\u3057\u306a\u3044\u3002\u3053\u306e\u6761\u4ef6 \u3068\uff0ce\u0302dq = [ 0 \u03c9\u0302\u03c6 ]T \uff0c(10)\u5f0f\u3068 (11)\u5f0f\u3092 (9)\u5f0f\u306b\u4ee3\u5165\u3057\u6574 \u7406\u3059\u308b\u3068 (14)\uff0c\n\u03b5\u0307 A\u03b5 \u2212 B\u03c6(\u03b8 \u2212 \u03b8\u0302)|\u03c9| \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3sgn\u03c9\n\u22121\n\u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 (12)\n\u3068\u306a\u308b\u3002\u3053\u3053\u3067\uff0csgn\u03c9\uff1a\u56de\u8ee2\u901f\u5ea6\u306e\u6975\u6027\u3067\u3042\u308b\u3002 \u6700\u5f8c\u306b\uff0c\u3053\u306e (12)\u5f0f\u306b\u30dd\u30dd\u30d5\u306e\u8d85\u5b89\u5b9a\u8ad6\u3092\u9069\u7528\u3059\u308b\u3068\uff0c MRAS\u3092\u5b89\u5b9a\u306b\u3059\u308b\u56de\u8ee2\u901f\u5ea6\u63a8\u5b9a\u5024 \u03c9\u0302\u3068\u78c1\u6975\u4f4d\u7f6e\u63a8\u5b9a\u5024 \u03b8\u0302 \u304c\u305d\u308c\u305e\u308c (13)\uff0c(14)\u5f0f\u3068\u3057\u3066\u6c42\u307e\u308b (15)\u3002\n\u03c9\u0302 = r1(\u03b5\u03b4 \u2212 \u03b5\u03b3 sgn \u03c9\u0302) + r2 \u222b (\u03b5\u03b4 \u2212 \u03b5\u03b3 sgn \u03c9\u0302)dt\n\u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7(13)\n\u03b8\u0302 = \u222b \u03c9\u0302dt \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7(14)\n\u305f\u3060\u3057\uff0cr1\uff0cr2 \u306f\u63a8\u5b9a\u30a2\u30eb\u30b4\u30ea\u30ba\u30e0\u306e\u6bd4\u4f8b\u30b2\u30a4\u30f3\u3068\u7a4d\u5206\u30b2 \u30a4\u30f3\u3067\u3042\u308b\u3002 \u30082\u30fb3\u3009 MRAS\u306e\u69cb\u6210 \u672c\u8ad6\u6587\u3067\u691c\u8a0e\u3059\u308b\u78c1\u6975\u4f4d\u7f6e\u30bb\u30f3 \u30b5\u30ec\u30b9\u5236\u5fa1\u30b7\u30b9\u30c6\u30e0\u306e\u69cb\u6210\u3092Fig. 2\u306b\u793a\u3059\u3002\u5f93\u6765\u306e PMSM \u975e\u5e72\u6e09\u30d9\u30af\u30c8\u30eb\u5236\u5fa1\u30b7\u30b9\u30c6\u30e0\u306bMRAS\u306e\u8981\u7d20\u3067\u3042\u308b\u9069\u5fdc\u30e2 \u30c7\u30eb\u3068\u901f\u5ea6\u30fb\u4f4d\u7f6e\u63a8\u5b9a\u30d6\u30ed\u30c3\u30af\u3092\u4ed8\u52a0\u3057\u305f\u69cb\u6210\u3068\u306a\u3063\u3066\u3044 \u308b\u3002\u9069\u5fdc\u30e2\u30c7\u30eb\uff0c\u901f\u5ea6\u30fb\u4f4d\u7f6e\u63a8\u5b9a\uff0c\u975e\u5e72\u6e09\u6f14\u7b97\u306e\u5404\u30d6\u30ed\u30c3\u30af \u3067\u5b9f\u884c\u3055\u308c\u308b\u6f14\u7b97\u5f0f\u306f\u305d\u308c\u305e\u308c (3)\uff0c(13)\uff0c(14)\uff0c(8)\u5f0f\u3067\u3042 \u308a\uff0c\u305d\u306e\u4e2d\u3067\u975e\u5e72\u6e09\u6f14\u7b97\u30d6\u30ed\u30c3\u30af\u306f\u901a\u5e38\u306e\u30bb\u30f3\u30b5\u4ed8\u304d\u30d9\u30af \u30c8\u30eb\u5236\u5fa1\u3067\u3082\u5fc5\u8981\u306a\u6f14\u7b97\u3067\u3042\u308b\u3002\u3053\u306e\u3053\u3068\u304b\u3089\uff0c\u672c\u8ad6\u6587\u3067 \u306e\u975e\u5e72\u6e09\u30bb\u30f3\u30b5\u30ec\u30b9\u5236\u5fa1\u306f\u5f93\u6765\u306e\u30d9\u30af\u30c8\u30eb\u5236\u5fa1\u306b (3)\uff0c(13)\uff0c (14)\u5f0f\u306e\u7c21\u5358\u306a\u6f14\u7b97\u3092\u8ffd\u52a0\u3059\u308b\u3060\u3051\u3067\u5b9f\u73fe\u3067\u304d\uff0c\u6f14\u7b97\u8ca0\u8377 \u306f\u308f\u305a\u304b\u3067\u3042\u308b\u3002\u3057\u304b\u3082\uff0c\u96fb\u6d41\u5236\u5fa1\u306e\u51fa\u529b\u4fe1\u53f7\u3092MRAS\u306e \u5165\u529b\u4fe1\u53f7\u3068\u3057\u305f\u305f\u3081\uff0c\u5236\u5fa1\u52d5\u4f5c\u306b\u4f34\u3063\u3066\u767a\u751f\u3059\u308b\u6301\u7d9a\u7684\u306a\u5909 \u52d5\u3092\u542b\u3093\u3060\u4fe1\u53f7\u304c\u5165\u529b\u3055\u308c\u308b\u3053\u3068\u306b\u3088\u308a\uff0cMRAS\u306e\u78c1\u6975\u4f4d \u7f6e\u63a8\u5b9a\u306b\u8981\u6c42\u3055\u308c\u308b\u5165\u529b\u4fe1\u53f7\u306e PE\uff08Persistently Exciting\uff09 \u6027\u3092\u6e80\u8db3\u3067\u304d\u308b\u306e\u304c\u7279\u9577\u3068\u306a\u3063\u3066\u3044\u308b\u3002\n3. \u30d1\u30e9\u30e1\u30fc\u30bf\u5909\u52d5\u306b\u5bfe\u3059\u308b\u78c1\u6975\u4f4d\u7f6e\u306e\u5f71\u97ff\nPMSM\u306e\u30d1\u30e9\u30e1\u30fc\u30bf\uff08\u5dfb\u7dda\u62b5\u6297\uff0c\u540c\u671f\u30a4\u30f3\u30c0\u30af\u30bf\u30f3\u30b9\uff0c \u8a98\u8d77\u96fb\u5727\u5b9a\u6570\uff09\u304c\u5909\u52d5\u3057\u305f\u5834\u5408\u306b\u304a\u3044\u3066\uff0c\u63d0\u6848\u3059\u308bMRAS \u3092\u7528\u3044\u305f\u78c1\u6975\u4f4d\u7f6e\u30bb\u30f3\u30b5\u30ec\u30b9\u5236\u5fa1\u306e\u6027\u80fd\u3092\u660e\u78ba\u5316\u3059\u308b\u305f\u3081\uff0c \u305d\u308c\u305e\u308c\u306e\u30d1\u30e9\u30e1\u30fc\u30bf\u5909\u52d5\u306b\u5bfe\u3059\u308b\u5b89\u5b9a\u6027\u3068\u78c1\u6975\u4f4d\u7f6e\u63a8\u5b9a \u8aa4\u5dee\u3078\u306e\u5f71\u97ff\u3092\u8ff0\u3079\u308b\u3002 \u30083\u30fb1\u3009 \u5dfb\u7dda\u62b5\u6297\u5909\u52d5\u306e\u5f71\u97ff\n\u30083\u30fb1\u30fb1\u3009 \u5dfb\u7dda\u62b5\u6297\u5909\u52d5\u6642\u306e\u8aa4\u5dee\u65b9\u7a0b\u5f0f \u898f\u7bc4\u30e2\u30c7\u30eb \u306e\u5dfb\u7dda\u62b5\u6297\u3092 Rm\uff0c\u540c\u671f\u30a4\u30f3\u30c0\u30af\u30bf\u30f3\u30b9\u3092 Lm \u3068\u304a\u304f\u3068\u898f\u7bc4 \u30e2\u30c7\u30eb\u306f (15)\u5f0f\u3068\u306a\u308b\u3002\n\u02d9\u0302x = Am x\u0302 + Bmr \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 (15)\n\u3053\u3053\u3067\uff0cx\u0302 = [ i\u0302d i\u0302q ]T \uff1a\u30e2\u30c7\u30eb\u96fb\u6d41\uff0cr = [ v\u2217\u03b3 v\u2217\u03b4 ]T \uff1a\u96fb\u5727\n\u6307\u4ee4\uff0cAm = \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3\u2212Rm/Lm 0\n0 \u2212Rm/Lm\n\u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6\uff0cBm = \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a31/Lm 0\n0 1/Lm \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6\u3067 \u3042\u308b\u3002 \u03b3-\u03b4\u63a8\u5b9a\u56de\u8ee2\u5ea7\u6a19\u306b\u304a\u3051\u308b (2a)\u5f0f\u3067\u62b5\u6297\u5024\u304c\u30e2\u30c7\u30eb\u5024 Rm \u3068\u305a\u308c\u3066 R\u3068\u306a\u308a\uff0c\u540c\u671f\u30a4\u30f3\u30c0\u30af\u30bf\u30f3\u30b9\u3068\u8a98\u8d77\u96fb\u5727\u5b9a\u6570\u304c \u5909\u5316\u305b\u305a\u306b\u305d\u308c\u305e\u308c\u306e\u30e2\u30c7\u30eb\u5024 Lm\uff0c\u03c6m \u3068\u540c\u4e00\u306e\u5834\u5408\u3092\u6c42 \u3081\u308b\u3068 (16)\u5f0f\u3068\u306a\u308b\u3002\nx\u0307 = ARx \u2212 \u03c9\u0302Jx + Bmu \u2212 Bme\u03b3\u03b4m \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 (16)\n\u3053\u3053\u3067\uff0ce\u03b3\u03b4m = [e\u03b3m e\u03b4m]T\uff1a\u6c38\u4e45\u78c1\u77f3\u306b\u3088\u308b\u8a98\u8d77\u96fb\u5727\uff0c e\u03b3m = \u2212\u03c9\u03c6m sin(\u03b8 \u2212 \u03b8\u0302)\uff0ce\u03b4m = \u03c9\u03c6m cos(\u03b8 \u2212 \u03b8\u0302)\uff0cAR =\u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3\u2212R/Lm 0\n0 \u2212R/Lm \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 \u540c\u69d8\u306b\u3057\u3066\uff0c\u5165\u529b\u96fb\u5727 u \u3092 (8) \u5f0f\u304b\u3089\u6c42\u3081\u308b\u3068 (17) \u5f0f\u3068 \u306a\u308b\u3002\nu = r + \u03c9\u0302B\u22121 m Jx + e\u0302dqm \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 (17)\n\u3053\u3053\u3067\uff0ce\u0302dqm = [ 0 \u03c9\u0302\u03c6m ]T \uff1a\u6c38\u4e45\u78c1\u77f3\u306b\u3088\u308b\u63a8\u5b9a\u8a98\u8d77\u96fb\u5727\n(17)\u5f0f\u3092 (16)\u5f0f\u306b\u4ee3\u5165\u3057\u305f\u5f0f\u3068 (15)\u5f0f\u304b\u3089\uff0c\u5dfb\u7dda\u62b5\u6297 \u5909\u52d5\u6642\u306e\u8aa4\u5dee\u65b9\u7a0b\u5f0f\u3092\u6c42\u3081\u308b\u3068\n\u03b5\u0307 = Am\u03b5 + (Am \u2212 AR)x \u2212 Bm(e\u0302dqm \u2212 e\u03b3\u03b4m) \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 (18)\n\u3068\u306a\u308b\u3002(18)\u5f0f\u306f (9)\u5f0f\u3068\u540c\u69d8\u306b\n428 IEEJ Trans. IA, Vol.132, No.3, 2012" + ] + }, + { + "image_filename": "designv11_33_0002358_1077546317743121-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002358_1077546317743121-Figure1-1.png", + "caption": "Figure 1. Drum shearer: (a) factual picture; and (b) working process.", + "texts": [ + " Simulation results indicate that the proposed method exhibits a good control effect in suppressing the dynamic load and an improved adaptive ability to the complex and changeable working conditions. The proposed method also can be useful for the electromechanical transmission systems with similar structure applied in other areas. Keywords Active torque control, dynamic load suppression, electromechanical transmission system, impact load The drum shearer is one of the major parts of the fully mechanized coal mining outfit equipment, and plays a decisive role in improving the cutting efficiency of the fully mechanized coal face, as shown in Figure 1. Additionally, the drum shearer cutting unit is an electromechanical coupling system that mainly consists of a driving motor, gear transmission system, and drum, and consumes 80\u201390% of the entire installed power of the drum shearer (Shu et al., 2015). However, a coal seam is always non-homogenous with hard parcels and rock intercalation, and causes various concerns, including strong impact, heavy loads, and large fluctuations to the drum. Therefore, the cutting unit is one of the least reliable components of the drum shearer (Liu C et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002079_1350650117719602-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002079_1350650117719602-Figure1-1.png", + "caption": "Figure 1. (a) Schematic diagram of noncircular bearing; (b) schematic diagram of two-lobe journal bearing.", + "texts": [ + " In the present work, a faster and more accurate method has been suggested to compute the dynamic characteristic parameters of a gas bearing system. Present work also illustrates the dynamic behavior of the rotor bearing assembly by considering two degree-of-freedom systems for the two-lobe selfacting gas journal bearing. The outcome of the numerical simulated results is expected to be quite useful for the innovative developments in the design of the gas bearing system. Analysis A schematic of the gas journal bearing has been shown in Figure 1. The governing Reynolds equation for a gas-lubricated journal bearing is expressed as follows10,16,19 @ @x h3 12 p @p @x \u00fe @ @y h3 12 p @p @y \u00bc U 2 @ ph\u00f0 \u00de @x \u00fe @ ph\u00f0 \u00de @t \u00f01\u00de where p is the gas film pressure, h is the gas film thickness, and U is relative sliding velocity between two surfaces. By using the following nondimensional parameters, equation (1) is expressed in the nondimensional form as follows p \u00bc p ps , h \u00bc h Cr , u \u00bc R! Pa R C 2 , \u00bc x R , \u00bc y R , \u00bc !t @ @ h3 12 p @ p @ ! \u00fe @ @ h3 12 p @ p @ " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002655_jifs-169558-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002655_jifs-169558-Figure1-1.png", + "caption": "Fig. 1. A simulated crack on one tooth of the sun gear.", + "texts": [ + " [29] compared the responses caused by crack paths of a straight line and a parabola and stated that the latter is correct to simulate a crack. Pandya et al. [30] revealed the effects of the backup radio and pressure angle on the crack propagation path. Based on these results, the tooth crack in this paper is set on the tooth root of the sun gear in the first stage of the gearbox. To analyze the effects of different levels of the crack, the crack is simulated from its initial position and then propagates along both the tooth width and depth simultaneously, as shown in Fig. 1. For the modeled sun gear, its specification is: the tooth number is 20, the module is 1 mm, the tooth width is 10 mm, the contact ratio is 1.635, and the pressure angle is 20\u25e6. Refer to Refs. [18, 31], the crack begins at the root circle. Assume that its initial position is marked as Q shown in Fig. 2(a). The crack grows along two directions. First, along the crack depth, its propagation path is simulated by a parabola. To control the curvature of the path, two angles are set. The intersection angle of the tangent line and the gear tooth centerline is , and the point P is the vertex of the parabola representing the crack propagation path along the crack depth" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000988_s00542-010-1189-3-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000988_s00542-010-1189-3-Figure3-1.png", + "caption": "Fig. 3 Free body diagram of the HDD spindle system", + "texts": [ + "2 Nonlinear equations of motion The proposed stability analysis assumes that the rotating disk-spindle system supported by the coupled journal and thrust bearings is linear and that the dynamic coefficients of the coupled journal and thrust bearings are time-invariant, i.e., constant. However, as the disk-spindle system rotates, the dynamic coefficients change depending on the relative positions of the rotating shaft and the sleeve. This paper derives the nonlinear equations of a rotating disk-spindle system supported by coupled journal and thrust bearings to show that the proposed stability analysis method is also effective for a nonlinear rotating disk-spindle system supported by coupled journal and thrust bearings. Figure 3 shows the free body diagram of the HDD spindle system, including the bearing reaction force, rotor weight, and centrifugal force due to the mass unbalance. The motion of the rigid rotor of the HDD spindle system can be described in terms of the five degrees of freedom, i.e., the three translational displacements in the x, y and z directions and two tilting displacements in the hx and hy directions, so that the five nonlinear differential equations can be derived, as follows: ma\u20aceX \u00bc FxJ1 \u00fe FxJ2 mueu _h2 z cos / c2 \u00fe FzT1 \u00fe FzT2\u00f0 \u00des2 \u00feWX \u00f07\u00de ma\u20aceY \u00bc FxJ1 \u00fe FxJ2 mueu _h2 z cos / s1s2 \u00fe FyJ1 \u00fe FyJ2 mueu _h2 z sin / c1 FzT1 \u00fe FzT2\u00f0 \u00des1c2 \u00feWY \u00f08\u00de ma\u20aceZ \u00bc FxJ1 \u00fe FxJ2\u00f0 \u00dec1s2 \u00fe FyJ1 \u00fe FyJ2 s1 \u00fe FzT1 \u00fe FzT2\u00f0 \u00dec1c2 \u00fe mag\u00fe mug\u00f0 \u00de mueu _h2 z sin /s1 cos /c1s2\u00f0 \u00de \u00feWZ \u00f09\u00de Ixc2 \u20achx \u00bc Ix \u00fe Iy Iz s2 _hx _hy Iz _hy _hz \u00feMxJ1 \u00feMxJ2 \u00feMxT1 \u00feMxT2 zu mugs1 mueu _h2 z sin / mugc1c2eu sin / WY Lf \u00f010\u00de Iy \u20achy \u00bc Ix Iz\u00f0 \u00dec2s2 _hx _hx \u00fe Izc2 _hz _hx \u00feMyJ1 \u00feMyJ2 \u00feMyT1 \u00feMyT2 \u00fe zu mugc1s2 mueu _h2 z cos / \u00fe mugc1c2eu cos /\u00feWXLf \u00f011\u00de where c1, s1, c2 and s2 are coshx, sinhx, coshy and sinhy, respectively, and mu, eu, zu and / are the mass unbalance of the rotor of the HDD spindle system and the positions of the mass unbalance from the mass center G, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001678_icacc.2011.6016394-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001678_icacc.2011.6016394-Figure2-1.png", + "caption": "Figure 2. Frames and force acting on the helicopter", + "texts": [ + " Vector, p = [x y z y represents the translational displacements, hV = [u v w)' is the translational velocities of the helicopter,8=[\u00a2 () Vft is the Euler angles describing helicopter attitudes, and (J) = [p q r r is the angular velocity vector. In this section, we concentrate on the force and moment acting on the helicopter, and the six degrees-of-freedom (6 DOF). For more information about the helicopter dynamic models can refer to references [I] [2] [10] [11]. To better describing helicopter's locations and attitudes, two three dimensional right handed Cartesian frames are introduced [11], shown in Fig. 2. They are Body Frame (BF) and Earth Inertial Frame (EF). The origin of BF is the COG in helicopter body, with the x-axis pointing through the body head, y-axis pointing to the right seen from above and z-axis pointing downwards and perpendicular to x-y plane. The EF called the NED frame as well is used as a reference frame in order to apply the Newtonian mechanics. The x-axis is defined to point north, z-axis pointing downwards vertically and y-axis perpendicular to both. An orthonormal transformation matrix can be given as (1), describing the rotation from the BF to the EF [11]. [ cos Bcos \\'! sin \u00a2sin Beos 1jI- cos \u00a2sin IjI cos\u00a2sinBcos\\,!+sin\u00a2sin\\,! ] (1) It is well known that the force acting on the helicopter is very complex. By neglecting minor force, we can simplify the force to main rotor thrust ( T mr ), tail rotor thrust (T,r ) and gravity of helicopter [11], as shown in Fig.2. The force acting on the COG can be written as (2). bF=[:;:j=[ \ufffd:\ufffd\u00b7S:\ufffd.b\ufffd:' l+[1\ufffd.l+mg[Si\ufffd; i\ufffdo\ufffd8l F -Tm. cosal< cosb], 0 cos!/!sm8 (2) Where r/J is the roll Euler angle, \u00b0 is the pitch Euler angle, als is the longitudinal flapping angle of main rotor, and bls is the lateral one. According to Newton-Euler equations we can get the helicopter rigid body equations of motion as b V = \ufffd. b F _ m x bV m riJ = rl('r - m x (I\u00b7m)) (3) where m is the mass of the helicopter, and I is the moment of inertia matrix of the helicopter defined as below" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003253_012016-Figure11-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003253_012016-Figure11-1.png", + "caption": "Figure 11. Principle of actuator integration (a), manufactured hip stem with embedded actuator system (b), computer tomography scan with detailed view of the actuator system (c) [2].", + "texts": [ + " However, these are not regular examinations and an implant loosening is often detected too late. A routine monitoring by means of an implant-embedded system would be useful. This paper describes the integration of a materially bonded, hermetically encapsulated actuator in a hip implant for a defined vibrational excitation within the implant and the surrounding bone for monitoring implant loosening and structural health. Variations in the implantbone-interface lead to a shift in the natural frequencies, which can be detected by wireless data transmission. The hip stem (Figure 11) was additively manufactured by LBM from Ti6Al4V titanium alloy powder. For this material, a heat treatment is necessary for adjusting micro-structural and mechanical properties and to relieve residual stresses. However, conventional post-process heat treatments cannot be applied due to the embedded temperature-sensitive actuator system. Therefore, in a first step only the base body of the hip stem including a cavity for incorporation of the actuator system is manufactured, cleaned and subsequently heat-treated. After the heat treatment, the alloy contains alpha and beta phases. The encapsulated actuator is then placed into the cavity flush to the top surface. Afterwards, the part is positioned in the LBM machine and the additive process is continued till the part is finished. After support removal, the cone for fitting the ball head and a mounting hole for the impactor was machined and the unmachined areas are shot-blasted. The finished part is shown in Figure 11b. Computer tomography verifies the material bond between the actuator system and the hip stem (Figure 11c). Static properties of the completed hip stem correspond to the properties of the volume fractions of the heat-treated and the as-build parts. Dynamic measurements experimentally validated the functionality of the embedded actuator. A numerical modal analysis using a finite element model is utilized to identify the initially unknown natural frequencies. The first two modes for lateral and vertical bending of the hip stem are located at 2180 Hz and 3488 Hz, resp. (Figure 12). These frequency values are used for the excitation for the WTK IOP Conf" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002284_ssrr.2017.8088167-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002284_ssrr.2017.8088167-Figure6-1.png", + "caption": "Fig. 6 Axis of support area and section division", + "texts": [ + " 0hhS max where h0 and hmax are obtained using the following expressions. coscossin0 hdh cos)( 2/122 max hdh Here, in order to make the gait generation method easier to understand, some specific examples will be given. To apply stability control to the motion, each edge of the robot's torso that is grounded is defined as follows. The supporting edge at the front of the robot is defined as Axis 1, and the remaining edges are defined as Axis 2, Axis 3, and Axis 4 in the counterclockwise direction. Also, as seen in Fig. 6, the endeffectors are defined as Leg 1, Leg 2, Leg 3, and Leg 4 in the same manner. Next, we describe the gait generation method for the crawling motion with stability control. First, while the robot is still in its initial posture, a NESM is calculated for each axis. After that, the discriminant, Di, of the NESM for each axis is calculated using the following equation. 1_thii SSD i =1,2,3,4 ) where Si is the NESM for ith axis and Si_th is its threshold value. While performing the crawling motion on the flat surface, the threshold value of the NESM is determined based on the magnitude of the center of mass motion with reference to the robot's coordinate system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003010_iccais.2018.8570323-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003010_iccais.2018.8570323-Figure1-1.png", + "caption": "Fig. 1. Block diagram of Linear Feedback and LQR Control", + "texts": [ + "1489 -8.0453e-26 A 0.0001 0.0003 -0.2560 -0.0044 -0.0133 -0.0292 0 0 B 1 0 0 0 0 1 0 0 0 0 1 0 0 0 0 1 C 0 0 0 0 0 0 0 0 D III. LINEAR FEEDBACK In this control method, the pole placement method is presented. If the system declares fully controllable, the poles of the closed-loop system can be placed at any desired location by means of state feedback via a suitable state feedback gain matrix [14]. The block diagram of feedback control system is shown in \u201cFig. 1\u201d. The control signal is produced as a function of the system error, the reference roll or yaw angle and the actual output roll or yaw angle is compared, then the value of controller gain (K) is adjusted to control the input signal for the desired output roll or yaw angle. The final computed value of gain matrix K 8.6544 -36.0379 10.5165 -86.3719 631.3773 90.1786 -327.6410 303.6114 K IV. LINEAR QUADRATIC REGULATOR LQR is a controller that is alike to pole placement method, but instead of using pole placement method to select the poles location, the value of feedback gain matrix K is calculated by minimizing the cost function to obtain the design requirement", + "00 \u00a92018 IEEE 368 The cost function for LQR is defined as 0 T TJ x Qx u Ru dt Where Q is the weighting function of states and R is the weighting function of control variables [14]. The values of Q and R, are given below which is used to calculate the gain matrix K. The weight matrix Q and control matrix R 10 0 0 0 0 10 0 0 0 0 10 0 0 0 0 10 Q 1 0 0 1 R The final computed value of gain matrix K 4.1742 -2.0060 -6.1360 -2.5386 1.6638 -0.4535 -5.4706 -0.9007 K Block diagram for LQR and linear feedback are same as shown in \u201cFig. 1\u201d, with the only variation in the values of gain matrix K. V. SIMULATION AND RESULTS The Simulink diagram of linear feedback control and LQR control is same for the lateral control of an aircraft as shown in \u201cFig. 2\u201d, with the only difference in finding the values of gain matrix K. The gain of linear feedback controller has been chosen through the pole placement method, while the gain of LQR controller is selected through the weighting Q and control R matrices. In this paper, we considered aileron and rudder deflection as control inputs for lateral control" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002709_978-1-84996-220-9_5-Figure5.14-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002709_978-1-84996-220-9_5-Figure5.14-1.png", + "caption": "Figure 5.14 Target position of the planning task: (a) distance view and (b) narrow view", + "texts": [ + " The robot model has 43 DoF and for each limb there are two 3D models, one for visualization and one simplified model for collision checking purposes. The robot is operating in a kitchen environment, which is also modeled with full and reduced resolution for visualization and collision checking as described in Section 5.2. The starting position of the robot is located outside the kitchen and a trajectory for grasping an object is searched. In this experiment the target object is placed inside the fridge (Figure 5.14(a)). For this task, the planner uses the subsystems Platform, Torso, Right arm, Right Wrist and Right Hand. In our test setup the subsystem for the right hand consists of six instead of eight joints because the two middle and the two index finger joints are coupled and thus are counted as one DoF. The overall number of joints used for planning, and therefore the dimensionality of the C-space, is 19. We make the assumption that a higher level task planning module has already calculated a goal position for grasping the object, thus cgoal , the target in C-space, is known" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001452_1464419311408949-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001452_1464419311408949-Figure5-1.png", + "caption": "Fig. 5 The test rig", + "texts": [ + " During the analysis, the angular acceleration is specified as the input from the motor. Analyses are performed for various levels of unbalance and defect conditions. Finally, fast Fourier transform and short-time Fourier transform are processed and Campbell diagrams are obtained using simulated data. In what follows, an experimental procedure is described and then the numerical and experimental results are presented and compared. The test rig designed and developed for the validation process is shown in Fig. 5. A heavy rigid block is used as a bench and it is isolated from the ground using soft supports. A separate alternating current (AC) motor is used for driving the rotor. The motor provides rotational excitation input to the system and it is accelerated until 1000 r/min. A pulley\u2013belt mechanism is used for power transmission from motor to the rotor. The test rig is designed such that the ball bearing, shaft, and disc components can be disassembled easily. This allowed replacing individual components quickly and made it possible to experiment with various components with different characteristics" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002377_978-3-319-68826-8_7-Figure7.69-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002377_978-3-319-68826-8_7-Figure7.69-1.png", + "caption": "Fig. 7.69 Schematic of WMR Pioneer 2-DX with laser rangefinder mounted", + "texts": [ + " The implementation of such a task is possible by way of combination of GS and OA behavioural controls, along with a relevant switching function. Another examples of local methods of generation of robots\u2019 motion trajectories are e.g. potential fields method [1, 60, 62] or elastic band approach [8]. The sensory systemofWMRPioneer 2-DX,whose design is discussed inSect. 2.1, was equipped with an additional Banner LT3PU laser rangefinder, installed on the upper panel of the WMR frame and marked as sL in Fig. 7.69. The axis of the laser rangefinder beam is parallel to the symmetry axis of theWMR frame, \u03c9Lr = 0\u25e6, and its dimensions are, respectively, 90\u00d7 70\u00d7 35 [mm]. The laser rangefinder was calibrated to measure distance within the limits of dLr \u2208 \u30080.4, 4.0\u3009 [m]. It is powered with 12 [V], the output signal UL \u2208 \u30080, 10\u3009 [V] is converted by a 12-bit A/D converter of the digital signal processing board into adequately scaled information about the distance of the selected WMR point to the obstacle. Distance to obstacles is measured by the sensory system of WMR Pioneer 2- DX and consists in cyclical activation of measurements carried out by individual ultrasonic sensors su1, " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002033_elma.2017.7955470-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002033_elma.2017.7955470-Figure1-1.png", + "caption": "Fig. 1. General arrangement drawing of the magnetic gear", + "texts": [ + " The harmonic analysis has been performed using Fast Fourier Transform (FFT) method. The modelling of the magnetic gear\u2019s field is implemented by ANSYS Maxwell [11] and the results from the harmonic analysis are obtained with MATLAB software [12]. II. MAGNETIC GEAR\u2019S CONSTRUCTIONS In this section two constructions of a coaxial magnetic gear have been considered. The dimensions of the gear for both constructions are the same. The general arrangement drawing of a coaxial magnetic gear is depicted in Fig. 1. The first construction has 4 inner permanent magnet pole pairs, 22 outer permanent magnet pole pairs and 26 steel segments. The second one has 6 inner permanent magnet pole pairs, 18 outer permanent magnet pole pairs and 24 steel segments. The sketch with the dimensions of the magnetic gear\u2019s construction is depicted in Fig. 2 and the dimensions are shown in Table I. The permanent magnets, which are used in the permanent magnet rotors, are made of NdFeB35 alloy [13]. The material used for the steel segments is a low carbon steel AISI 1008 [14]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003008_icems.2018.8549405-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003008_icems.2018.8549405-Figure4-1.png", + "caption": "Fig. 4. MMF produced by permanent magnet. (a). The equivalent magnetic circuit. (b). MMF distribution", + "texts": [ + "78125 2 r r r r r r r r r r m rr r b g bm bm m g m bm \u03bc\u03b8 \u03b8 \u03b2 \u03c4 \u03c4\u03bc \u03c0\u03b2 \u03c0 \u03b8 \u03b8 \u03c0 \u03c4 \u03c4 \u2265 \u039b = \u2212 + \u2212 + \u2212 \u2212 (13) Where br and r are the length of slot opening and slot pitch of rotor, respectively. Hence, the permeance of double-salient air gap structure can be obtained by combining (1), (12) with (13). B. The Flux Density Produced by Permanent Magnet By means of the Carter coefficient, the effects of stator and rotor slot can be compensated to obtain the MMF distribution in air gap produced by permanent magnets. The equivalent smooth rotor and stator, and the corresponding MMF distribution are shown in Fig.4(a) and Fig.4(b), respectively. Hence, the MMF produced by permanent in at air gap can be deduced in Fourier series form by ( ) ( ) ( )( )2 21 2 1 sin sin 6 2 1 82 1 8 m pm n nF F n n \u03c0 \u03b8 \u03c0\u2265 \u2212 = \u2212 \u2212 (14) Where 2 1 22 1 2 2 pm m c m pmpm pm R R RF H h R RR R R R R = + + + (15) And 1 1 2 0 1 0 0 = , = , =cs cr m m pm t m m k k g h h R R R w w\u03c4 \u03bc \u03bc \u03bc (16) Where t is the tangential length of the U-shaped magnetic core at air gap side, and kcs and kcr are the carter coefficient of stator and rotor, respectively. Therefore, combine (1) with (14), the flux density produced by permanent magnet can be calculated by C" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000945_ssd.2010.5585533-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000945_ssd.2010.5585533-Figure1-1.png", + "caption": "Figure 1. Stator of CPAES: Location of the three phase armature winding.", + "texts": [ + "00 \u00a920 1 0 IEEE The CPEAS concept presents a claw pole topology where the DC-excitation winding is located in the stator rather than in the rotor as in conventional claw pole machines. As a result, the brush-ring system and the associated mainte nance problem have been discarded, which represents cru cial cost and availability benefits. The machine is equipped by a three phase armature winding. In the manner of conventional claw pole alterna tors, the stator is made up of a laminated cylindrical mag netic circuit as shown in figure 1. The field winding is simply wound in a ring shape. Figure 2 shows the photo of one half of the stator field winding. The two halves of the field winding are inserted in both sides of the machine between the armature end-windings and the housing as illustrated in figure 3. They are con nected in series in such a way to produce additive fluxes. Following the transfer of the DC-excitation winding from rotor to stator, appropriate changes of the magnetic circuit have been introduced. These concerned mainly the rotor where the two iron plates with overlapped claws facing the air gap tum to be magnetically decoupled" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003768_s105261881905008x-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003768_s105261881905008x-Figure2-1.png", + "caption": "Fig. 2. Three most widespread types of profiled uniform-strength beams to loading with end force: triangular beam with constant thickness (1); parabolic beam with constant width (2); constarea beam with constant cross-section area (3).", + "texts": [ + " The second case of a uniformly distributed load is not much more complex, and it corresponds to loading of bridges, building structures, or wind loads on the limb of a tree. The third case explains the contradictions arising in the calculation of deflection according to beam theory, which were analyzed in Section 4, in order to turn to the total energy analysis of the possibility of increasing the weight efficiency of profiled springs. TYPE I: Loading with Concentrated Force (One Upper Stroke) (Fig. 1, I) Index i in expressions (1) and (2) and later denotes the type of beam (Fig. 2): i = 0 for a beam with constant cross-section sizes; i = 1 for a beam with constant thickness; i = 2 for a beam with constant width; i = 3 for a beam with constant cross-section area (constarea beam). The power dependences (1) and (2) with fixed length l contain five design parameters: initial cross-section sizes and , indices and in the laws of their changes (1) and the length of end part To define them unambiguously, the beam design must satisfy five conditions: (1) Specified stiffness C, i.e" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003710_tro.2019.2938348-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003710_tro.2019.2938348-Figure1-1.png", + "caption": "Fig. 1. (a) Parallel mechanism composed of a C\u2212(x)\u2212R(y)\u2212 C+(x) subchain and an M4 subchain. (b) M4 subchain is realized by three pairs of symmetric U \u2212 U chains interconnected through cylindrical joints as proposed in [10]. (c) Only two pairs of symmetric U \u2212 U chains are required if we employ additional interconnection with the C\u2212(x)\u2212R(y)\u2212 C+(x) subchain.", + "texts": [ + " Consider the maximal inscribing symmetric subspace Mmax and the minimal covering symmetric subspace Mmin with Mmax \u2282 Q \u2282Mmin. SQ1 (Q2) can be obtained by compressing Mmin or expanding Mmax. A. Compressing Mmin Some reflective-type adjoint-invariant submanifolds can be synthesized by assembling Q1 \u00b7Q2 \u00b7Q1 chains with a Mmin generator in parallel. Example 4 (KG for SC(x)(R(y))): Notice that SC(x)(R(y)) \u2282 C\u2212(x) \u00b7 R(y) \u00b7 C+(x) which is generated by cascading a pair of symmetric cylindrical joints (C\u2212(x), C+(x)) about the xy plane with a R(y) joint in the middle, as shown in Fig. 1(a). On the other hand the minimal covering symmetric subspace Mmin of SC(x)(R(y)) is M4. Then, we show that SC(x)(R(y)) = (C\u2212(x) \u00b7 R(y) \u00b7 C+(x)) \u2229M4. First, SC(x)(R(y)) belongs to both C\u2212(x) \u00b7 R(y) \u00b7 C+(x) and M4. Second, at home configuration e, the constraint forces of M4 is {e3, e6}, while that of C\u2212(x) \u00b7 R(y) \u00b7 C+(x) is {e2}. The feasible tangent space of the parallel mechanism at home configuration is simply {e1, e4, e5}. According to[5, Position 6], the parallel mechanism consisting of a M4 KG and a C\u2212(x)\u2212R(y)\u2212 C+(x) subchain is the KG for SC(x)(R(y)), as shown in Fig. 1(b) (only one subchain of the M4 KG is drawn here for clarity) . A practical mechanism can be derived by replacing the full M4 generator (e.g., [10, Example 5]) by its subchains M j 4 , and interconnecting M j 4 as well as the C\u2212(x)\u2212R(y)\u2212 C+(x) chain in a similar manner. This reduce the number of M j 4 subchains from 3 as required in the full M4 generator to 2, as shown in Fig. 1(c). Example 5 (KG for SC(x)(Iez\u0302\u03c0/2(M2A))): It is easy to see that SC(x)(Iez\u0302\u03c0/2(M2A)) is equivalent to SC(y)(M2A) up to the conjugation map Ie\u2212z\u0302\u03c0/2 . Recall SC(y)(M2A) = SR(y)(ST (y)(M2A)) \u2282 SR(y)(PL(x)) \u2282 R\u2212(y) \u00b7 PL(x) \u00b7 R+(y), where PL(x) is realized by cascading three revolute joints parallel to x, and (R\u2212(y),R+(y)) are a pair of symmetric revolute joints about the xy plane. Combine the two distal revolute joints into a U(x, y\u2212) pair and a U(x, y+) pair (y\u2212 and y+ in theU pairs are used to show that they are symmetric about the xy plane)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003832_s11665-019-04449-6-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003832_s11665-019-04449-6-Figure4-1.png", + "caption": "Fig. 4 Fracture toughness specimen dimensions for this study, in inches (1 in. = 2.54 cm)", + "texts": [ + "03B-30, MTS Fracture Toughness Clevis-640 series, and MTS Multipurpose Elite Fracture Toughness testing software. The specimen dimensions were based on the width as measured from the back of the specimen to the center of the loading pins. The thickness is half the width. This is a critical dimension when choosing what size specimen to make. It must be sufficiently thick enough to satisfy ASTM E399 require- ments (Ref 15). The yield strength and Young s modulus of the material need to be determined. As a result, tensile testing was performed according to ASTM E8 (Ref 16). Figure 4 shows the specimen size used for this study. The standard specimen dimensions were chosen in which W = 2.54 cm (1 in.). Based on this value, the thickness (B) was 1.3 cm (0.5 in.). The notch geometry was modified to hold the crack-opening displacement (COD) at the base of the notch as allowed by the ASTM Standard (Ref 15). There were several validity checks that were needed to make sure that the fracture toughness test was considered a valid plane strain KIC (Ref 15). By comparing the tensile properties of the longitudinal sample data from Tables 3 and 4, the percent error was 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001347_s10846-012-9725-2-Figure19-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001347_s10846-012-9725-2-Figure19-1.png", + "caption": "Fig. 19 The recommended shape of the spincopter\u2019s wing", + "texts": [ + " Due to the nonlinearities of the airfoil characteristics, increasing the angle of attack over 13\u25e6\u2013 15\u25e6, depending on the used airfoil, may cause certain issues (i.e. decreasing L/D ratio) [20]. Applying the results on a real spincopter model, an increased length and a narrowed width of the wing are suggested. Also, an angle of attack of 12 \u2212 13\u25e6 is recommended since it avoids the fore mentioned nonlinearities. Lowering the parameters TC1 and RC2 also has a positive influence on most of the physical indicators, but it reduces the robustness of the aircraft. Hence, decreasing those parameters for 30\u201340[%] does not cause any real weaknesses (Fig. 19). Spincopter controller is implemented using an arduino based autopilot board for multirotor crafts. The arduino compatibility makes it easy to program and adapt to spincopter construction and the onboard sensors (i.e. IMU, compass and sonar) provide feedback for the controller. An indirect form of Kalman filter is used to fuse the information from different sensors. Spincopter is controlled through a joystick, connected to a stan- dard PC platform. The communication between the PC and the autopilot board is implemented using standard XBee modules" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002459_robio.2017.8324729-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002459_robio.2017.8324729-Figure7-1.png", + "caption": "Fig. 7: Vertical obstacle avoidance flight", + "texts": [ + " (13) Then, the desired position along x axis, x\u2032 r, can be determined by observing the relative motion of the obstacle and unique solution can be calculated from eq. (12). During the process, Catenrary number is numerically calculated to estimate the shape of the cable and proper shape will be obtained to avoid contact between the cable and obstacle. 2) Vertical obstacle avoidance flihgt: As same as the above approach, when the movable area in the horizontal direction is limited, the vertical position of one UAV is changed as shown in Fig. 7. By changing the vertical position of a UAV, the lowest point of the hanged cable can be moved higher position on vertical direction compared to the highest position of the obstacle. The condition of this constraint is expressed as: z\u2032r > zo x\u2032 p = xp = const. (14) In this case, the desired position along z axis, z\u2032r, can be determined by observing the relative motion of the obstacle and unique solution can be calculated from eq. (12). This section describes flight experiments for the proposed shape estimation method explained in Section III and obstacle avoidance flight proposed in Section IV" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000751_011-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000751_011-Figure2-1.png", + "caption": "Figure 2. The scheme of the streetboard.", + "texts": [ + " (8) Solving the system of (8) together with (4) we can obtain, in principle, both the motion of the system and the constraint forces. The streetboard (sometimes called snakeboard; see [3]) is a variant of the skateboard in which the passive wheel assemblies can pivot freely around the vertical axis. By coupling the twisting of the human torso with appropriate turning of the wheels (where the turning is controlled by the rider\u2019s foot movement), the rider can generate a snake-like locomotion pattern without having to push off the ground. A simplified model is shown in figure 2, where the streetboard is composed of a rotor (representing the human torso), board and wheel axles. We assume that the front and rear wheel axles move through equal and opposite rotations. This is based on observations of human streetboard riders who use roughly the same phase relationship. A momentum wheel rotates about a vertical axis through the centre of mass, simulating the motion of a human torso. The physical parameters for the system are the mass of the board and the body m, the inertia of the rotor Jr, the inertia of the wheels about the vertical axes Jk, the inertia of the board J and the half-length of the board l. There are five degrees of freedom of the streetboard\u2013rider system: two for the position of the centre of mass in the x\u2013y-plane, one for the rotation of the board, one for the rotation of the rider\u2019s body and finally one degree of freedom corresponds to the rotation of the wheel axes. We denote the corresponding coordinates by (x, y, \u03b8, \u03c8, \u03b2). In figure 2 we can see that the rotation angle \u03b8 is defined with respect to the x-axis, while the rotation angles \u03c8 and \u03b2 are defined with respect to the board. Note that although other definitions of coordinates are possible, this one keeps the model as simple as possible. The Lagrangian of the streetboard mechanical system is given by L = 1 2m(x\u03072 + y\u03072) + 1 2Jk(\u03b2\u0307 + \u03b8\u0307 )2 + 1 2Jk(\u2212\u03b2\u0307 + \u03b8\u0307 )2 + 1 2Jr(\u03c8\u0307 + \u03b8\u0307 )2 + 1 2J \u03b8\u03072, (9) which corresponds to the unconstrained situation when there is no relation between the rotation of the wheel axles and the rotation of the board", + " Hence, the Lagrangian takes the form L = 1 2m(x\u03072 + y\u03072) + Jkb\u0307(t)2 + JrP\u0307 (t)\u03b8\u0307 + 1 2JrP\u0307 (t)2 + 1 2J \u03b8\u03072, (13) and the system of equations (11) reduces to a system of three equations for three unknown variables (x, y, \u03b8): \u2212mx\u0308 = 0, (14a) \u2212my\u0308 = 0, (14b) \u2212JrP\u0308 \u2212 J \u03b8\u0308 = 0. (14c) The mechanical model described in the previous section corresponds to a situation when the wheels are allowed to slip in the direction of their axles. In the real-world motion of the streetboard rider, this situation must not occur and therefore additional constraints must enter the problem. (i) The velocity v = (x\u0307, y\u0307) of the centre of mass of the streetboard must be parallel to the board, i.e. it must hold that (see figure 2) y\u0307 cos \u03b8 = x\u0307 sin \u03b8. (15) (ii) The velocities of both ends of the board must be vertical to the particular wheel axle. Since the board\u2013axle angles in the front \u2212\u03b2 and in the rear of the board \u03b2 are of the same size and opposite orientation with respect to the board, the ends of the board are moving on the osculating circle k1 (see figure 2). The centre of mass of the board is moving on the circle k2, which is concentric to k1. For the radii of both circles it holds that R2 = R1 cos \u03b2, R2 = l/tg \u03b2. (16) For the velocity of the centre of mass of the streetboard, it must hold that (see figure 2) v = (x\u0307, y\u0307) = \u2212 l\u03b8\u0307 tg \u03b2 (cos \u03b8, sin \u03b8). (17) Denote k = (cos(\u03b8 + \u03b2), sin(\u03b8 + \u03b2)), (18) the unit vector vertical to the rear axle (see figure 2). Definition (18) corresponds to the requirement that k forms angle \u03b2 with the board. Substituting, the scalar multiplication v \u00b7 k gives x\u0307 cos(\u03b8 + \u03b2) + y\u0307 sin(\u03b8 + \u03b2) = \u2212l\u03b8\u0307 cos2 \u03b2 sin \u03b2 . (19) The constraint (19) ensures that the velocities of both ends of the board are vertical to the particular wheel axle. Considering \u03b2 = b(t) a given function and substituting (15) into (19), we can rewrite nonholonomic constraints (15) and (19) in the form f 1 \u2261 y\u0307 cos \u03b8 \u2212 x\u0307 sin \u03b8 = 0, (20a) f 2 \u2261 \u03b8\u0307 l cos b(t) cos \u03b8 + x\u0307 sin b(t) = 0", + " Solving the system of differential equations is a nontrivial task even for simple problems. In this section we will show one possible solution procedure for the system (24) in the special case of streetboard motion. Setting the axles of the streetboard in a fixed position defined by the angle b(t) = \u03c0/4, the board is allowed to move only along a circle. This case corresponds to the situation when the streetboard rider does not move his feet with respect to the board. The centre of mass of the streetboard then moves along the circle with radius R2 = l/tg b = l (see figure 2). Let us solve the problem in the case when the rider is beginning to accelerate his torso rotation, i.e. P(t) = \u03b5t2/2. Let us assume the following initial conditions of the problem: x(0) = 0, y(0) = \u2212l, \u03b8(0) = 0, x\u0307(0) = 0. (26) By substituting b(t) = \u03c0/4, P(t) = \u03b5t2/2 into (24a) we get (x\u0308 \u2212 \u03b1x\u03072 tg \u03b8) ( m cos2 \u03b8 + J \u03b12 ) \u2212 Jr\u03b5\u03b1 = 0. (27) Taking into account that \u2212x\u0307\u03b1 = \u03b8\u0307 (see (24c)) and \u03b1 = 1 l cos \u03b8 (see (25)), equation (27) can be written in the form 1 cos \u03b8 (x\u0308 + x\u0307\u03b8\u0307 tg \u03b8) ( m + J l2 ) \u2212 Jr\u03b5 l = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003710_tro.2019.2938348-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003710_tro.2019.2938348-Figure2-1.png", + "caption": "Fig. 2. KG for SC(y)(M2A) composed of a U(x, y\u2212)\u2212R(x)\u2212 U(x, y+) subchain and an M5 subchain whose wrist plane (the plane passing through the three spherical joints in the wrist) is parallel to the xy plane. The y\u2212 and y+ axes of the pair of symmetric U pairs of the U(x, y\u2212)\u2212R(x)\u2212 U(x, y+) subchain intersect at a point Pxy in the xy plane.", + "texts": [ + " Example 5 (KG for SC(x)(Iez\u0302\u03c0/2(M2A))): It is easy to see that SC(x)(Iez\u0302\u03c0/2(M2A)) is equivalent to SC(y)(M2A) up to the conjugation map Ie\u2212z\u0302\u03c0/2 . Recall SC(y)(M2A) = SR(y)(ST (y)(M2A)) \u2282 SR(y)(PL(x)) \u2282 R\u2212(y) \u00b7 PL(x) \u00b7 R+(y), where PL(x) is realized by cascading three revolute joints parallel to x, and (R\u2212(y),R+(y)) are a pair of symmetric revolute joints about the xy plane. Combine the two distal revolute joints into a U(x, y\u2212) pair and a U(x, y+) pair (y\u2212 and y+ in theU pairs are used to show that they are symmetric about the xy plane). This yields a U(x, y\u2212)\u2212R(x)\u2212 U(x, y+) mechanism, as shown in the left subchain in Fig. 2. Its constraint force space is given by { [ xT , (Pxy \u00d7 x)T ]T }, where Pxy \u2208 R3 is a point in the xy plane. On the other hand SC(y)(M2A) \u2282Mmin =M5. It is realized as the Delta - Omni-wrist mechanism (the right subchain of Fig. 2), which contributes the constraint force space { [ 0, zT ]T }. The parallel mechanism formed by connecting these two subchains in parallel gives rise to the constraint force space {e1, e6}, and, therefore, it is a KG of SC(x)(Iez\u0302\u03c0/2(M2A)). B. Expanding Mmax The reflective-type adjoint-invariant submanifolds SM4 (R(z)) and SM4 (Hp(z)) can be synthesized by expanding the KG of its maximal inscribing symmetric subspace Mmax. Proposition 5: If SQ1 (Q2) is a reflective-type adjoint-invariant submanifold with Q1, a symmetric subspace ( =M5), and Q2 a Lie subgroup satisfying TeQ2 \u2282 h = [TeQ1, TeQ1] and ad\u03b7TeQ2 \u2282 TeQ2, \u2200\u03b7 \u2208 h, then a KG for SQ1 (Q2) could be synthesized by inserting aQ2 chain between each pair of symmetric sub-subchains in the KG for Q1, while reducing the corresponding DoFs in all interconnecting chains" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002306_tmi.2017.2776404-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002306_tmi.2017.2776404-Figure1-1.png", + "caption": "Fig. 1. Rotation in 2D and 3D space.", + "texts": [ + " Finally we present the algorithm to calculate the lower bound of the maximum TRE magnitude. Both the objective function and the constraint of the approximated problem take the form of TRE magnitude. Instead of handling with the rotation matrix R which should be orthogonal, we use the rotation axis n\u0302 and the rotation angle \u03b8 to represent the rotation in 3D space and investigate the isosurface of TRE magnitude. We start with the isoline of TRE magnitude in 2D space and then extend the conclusion to 3D space. 1) Isoline in 2D space: As shown in Figure 1(a), rotation in 2D space is to rotate the vector counterclockwise by angle \u03b8. The isoline of TRE magnitude in 2D space is defined as, IsoLine(C) = {p | \u2016Rp+ s\u2212 p\u2016 = C}. When the rotation angle is zero, the isoline is the whole Euclidean plane and C = \u2016s\u2016. When the rotation angle is not zero, we denote the rotation matrix R by [cos \u03b8,\u2212 sin \u03b8; sin \u03b8, cos \u03b8], denote the translation term s by (sx, sy), and denote point p by (px, py), then the points belonging to IsoLine(C) should satisfy, (px cos \u03b8 \u2212 py sin \u03b8 + sx \u2212 px)2 + (px sin \u03b8 + py cos \u03b8 + sy \u2212 py)2 = C2", + " Combining similar terms of px, py , we get a circle equation, (px \u2212 cx)2 + (py \u2212 cy)2 = C2/(4 sin2 \u03b8 2 ), in which (cx, cy) is the center, cx = 1 2 (sx \u2212 sy cot \u03b8 2 ), cy = 1 2 (sx cot \u03b8 2 + sy). Therefore the isoline of TRE magnitude in 2D space is a circle if the rotation angle is not zero. Denote the distance from the target location p to the circle center (cx, cy) by d, then we have \u2016TRE(p)\u2016 = 2d sin(\u03b8/2). (4) The case in which the rotation angle is zero could be regarded as a special case in which the circle center locates at infinity. 2) Isosurface in 3D space: As shown in Figure 1(b), the rotation in 3D space is to rotate the vector along axis n\u0302 = (nx, ny, nz) counterclockwise with angle \u03b8. According to Rodriguez formula [16], the rotation matrix could be calculated from the rotation axis n\u0302 and the rotation angle \u03b8. With respect to the rotation axis n\u0302, the vector from the coordinate origin to point p could be decomposed into the parallel part and the perpendicular part, denoted by p\u2016 and p\u22a5 respectively. p\u2016 is parallel to n\u0302 and p\u22a5 is perpendicular to n\u0302. Since the rotation matrix has no effect on the parallel part, we have, Rp\u2212 p = Rp\u22a5 \u2212 p\u22a5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003609_0954407019867491-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003609_0954407019867491-Figure4-1.png", + "caption": "Figure 4. The prototype vehicle with two axle motors prior to the differential.", + "texts": [ + " The sensor signals are transmitted to the brake- and motor-control modules having 5- and 1-ms sampling periods, respectively. For the details of the mathematical models used in the simulator, the reader is referred to studies of Bayar and colleagues23\u201325 for brevity. The parameters used for the simulator (sprung/ unsprung masses, drivetrain inertias, brake system parameters, etc.) belong to the hybrid electric vehicle mentioned by Bayar et al.23,24 A picture showing the overview of the drivetrain architecture for this vehicle is given in Figure 4. The data for the hydraulic friction brakes were not parametrized through experiments, rather the values were taken from another SUV of the same class, from dSPACE.26 Separate drivetrain models were built for each of the electric-vehicle drivetrain architectures shown in Figure 1. The electric motor and differential model parameters used for Architecture 1 are modified from the values used by Bayar et al.24 The electric motor parameters for on-board motors with side-shafts architecture are from Goggia et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003397_s00170-019-03894-w-Figure11-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003397_s00170-019-03894-w-Figure11-1.png", + "caption": "Fig. 11 Universal machine tool for the manufacturing of gear by freeform milling", + "texts": [ + " Once the tooth profile error of gear is given, the cutting pass number is decreasing as the tooth number increases, and so are the pressure angle and helix angle. But for the tooth module of gear, the bigger tooth module is, the more the cutting pass number is. Therefore, depending on the varying of workpiece parameters, the cutting pass number should be recalculated in view of the requirement of machining precision and efficiency. In order to verify the results of the proposed numerical analysis model, an experiment is performed by machining the typical herringbone gear on a large multiaxis machine tool. As shown in Fig. 11, the machine tool, which has four linear axes including X-axis, Y-axis, Z-axis, and U-axis, and two revolving axes including A-axis and C-axis, allows for the flexible machining of various gear types and sizes with the universal milling tool. Since the workpiece is a special herringbone gear without relief groove, as shown in Fig. 12, it is incapable for the traditional generating method or forming method to manufacture it. However, the free-form milling method is capable of machining it using the standard milling tools on the universal machine" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002633_s13738-018-1383-2-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002633_s13738-018-1383-2-Figure1-1.png", + "caption": "Fig. 1 Syringe carbon paste electrode (a). Small electric motor (b); model: N20; rated voltage: 3\u20136\u00a0 V DC; rated torque: 0.4\u00a0 kg cm; revolving speed: 30\u00a0rpm; motor size: 12 \u00d7 26\u00a0mm", + "texts": [ + " MTOAC was dissolved in bromoform as a proper carbon paste binder, where necessary. The electrode assembly including a tubular copper (14\u00a0cm length and 2\u00a0mm inner diameter) covered and insulated with a glass holder served as the electrode body and electrical contact. A plunger moved up and down to press the paste down when renewal of the electrode surface was needed. The small electric motor with an L-shape rotating blade slowly rotates to cut and remove excess paste and renewing the electrode surface after every measurement. Figure\u00a01 demonstrates the SCPE (left) and its schematic assembly (right). For the preparation of the SCPE, 10\u00a0mg of MWCNT, 64\u00a0mg of MTOAC and 246\u00a0mg of bromoform were homogenized in an ultrasonic bath for 15\u00a0min. Then, 90\u00a0mg of graphite powder was added to the mixture and homogenized again. The prepared paste was then packed into the copper pipe and pressed to release the captured air. A fresh surface was obtained by wiping, smoothing with the rotating blade and then used directly for voltammetric measurements" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000270_iemdc.2009.5075424-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000270_iemdc.2009.5075424-Figure8-1.png", + "caption": "Figure 8 Experimental setup to measure the air gap flux density of AFPMM.", + "texts": [ + " 7, the rotor was supported on a turntable so that it could rotate freely about a vertical axis. The flux-density probe of the gauss-meter was mounted on a platform with graduations to indicate the radial position of the probe with respect to the axis of the AFPMM. Axial positions of the field points were adjusted by inserting shims (each with thickness of 1 mm) between the magnetic pole and the probe. By turning the rotor to different angular positions, the flux density variations along a given radius R and axial position z can be measured. Fig. 8 shows the complete experimental setup for flux density measurements. Fig. 10 shows the variation of flux density along the circumferential direction at the mean radius (R = 0.0775 m). Agreement between computed and experimental results is satisfactory, thus verifying the validity of the proposed computation method. The experimental no-load line voltage waveform in Fig. 11 also confirms the low harmonic distortion in the AFPMM being studied. A semi-analytical method for computing the magnetic field of an axial-flux permanent-magnet machine with coreless armature has been presented" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003146_ichve.2018.8642003-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003146_ichve.2018.8642003-Figure1-1.png", + "caption": "Figure 1. Schematic of multiple wires", + "texts": [ + " In this paper, the partial capacitance matrix between conducting wires and insulated shield wires was firstly deduced by Maxwell's electrostatic equation. The formula for the induction voltage of insulated shield wires was obtained. Then the expression of the equivalent capacitance of power tap-off circuit was obtained by Thevenin\u2019s equivalence principle and the partial capacitance matrix. Finally, the actual line parameters were calculated and analyzed. II. THEORY AND ANALYSIS A n parallel conductors system was considered as shown in Fig. 1. Maxwell's electrostatic equation was used, the function between the conductor voltages with respect to ground u and the conductor per unit length charges q is defined as [4] QPU \u00d7= , (1) where U is the column vector of the u (kV); Q is the column vector of q (C/km); P is the potential coefficient matrix (km/F). The elements of P are computed by 978-1-5386-5086-8/18/$31.00 \u00a92018 IEEE i i i i ii r H r H P 2ln10182ln 2 1 6 0 \u00d7== , (2) ij ij ij ij ij d D d D P ln1018 2 ln 2 1 6 0 \u00d7== , (3) where Hi and Hj are average height above the ground-level of the ith and jth conductor (m); ri is the radius of conductor i (mm); dij is the mutual distance between i and j conductors (m); Dij is the distance between conductor i and the image of conductor j or vice versa (m); 0 is the vacuum permittivity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000946_ccdc.2012.6244274-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000946_ccdc.2012.6244274-Figure1-1.png", + "caption": "Fig 1The configuration of Propulsive Rocket Engines", + "texts": [], + "surrounding_texts": [ + "1 Instruction Recently, many countries have dedicated in developing heavy lifting launch vehicle (HLLV) technology to enhance their aviation capabilities. HLLV is aerodynamically unstable, structurally flexible, and strongly nonlinear, resulting a great challenge for the design of ascent flight control system. The main challenge from automatic engineer perspective is to guarantee high performance in the presence of model parameter perturbations, external disturbances and mismodeling dynamics. The launch control problem consist of automatic tracking the launch trajectory which was assumed to be optimally precalculated to achieve the desired control purpose. However, control analysis showed that depending solely on swinging central rocket engine can hardly meet the control requirements of large and medium-sized launch vehicle. so the combination of strap-on rocket boosters and central rocket engines is needed to control the attitude. At the same time, the rocket shows the characteristic of low, dense modal frequency, strongly coupled vibration and\nThis work is supported in part by the National Nature Science Foundation of China, No. 60674105, the Scientific Research Cultivation Project of Ministry of Education, No. 20081383 and the 2008 Spaceflight Support Foundation.\ncomplex partial deformation. what\u2019s more ,there exists a low-frequency resonance in the engine-servo loop. Thus a robust and reliable flight control system is necessarily demanded to meet the higher performance requirements. Facing to this kind of complex object, conventional ascent flight control system is based on the combination of variable gain and network approach. Nevertheless, this method is complex and difficult to design since the long parameter optimization process and repetitive works for the re-calibration networks. Besides, the control law switching moment will cause large transient tracking error. Consequently, many advanced guidance and control techniques were developed for dealing with the problems such as sliding mode control[1],neural network control[2], robust fault-tolerant control[3],etc\u2026Sliding mode control is robust with respect to matched internal and external disturbances, however ,undesired chattering produced by the high frequency switching of the control may be considered a problem for implementing the sliding mode control methods for some real applications[4]. In order to weaken the chattering, a lot of works have been done and some exciting research findings have been achieved in the 1980. Corradini, proposed the quasi-sliding mode approach, which includes continuous approximation one instead of sign function and the way to design the\n1713978-1-4577-2074-1/12/$26.00 c\u00a92012 IEEE", + "boundary layer[5]. Jian-Xin,in the use of internal model principle, designed a low-pass filter acting on switching function in the boundary layer resulting in a smooth signal[6]. Yip, P P developed a variable structure controller with a filter eliminating the chattering in control signal effectively[7]. In Kawamura.A \u2019s design, a conventional sliding mode control method with a new kind of disturbance observer is applied to servo drive system. Through the feed forward compensation, this method had reduced switching gain, as well as the chattering[8].Serrani,A made some changes to switching term to achieve a reduction in chattering[9]. Chen,Y proposed a fuzzy neural network adaptive switching gain algorithm and reduced the chattering greatly.\nAs for chattering problem, the literature [10] has proposed a novel sliding mode control methodology for MIMO system. Based on that paper, the authors in this paper developed the method and applied it to the HLLV attitude control system. This brief paper is organized as follows: the dynamic model of launch vehicle is briefly presented in section 2. Then, in section 3 the dynamic integral sliding mode control method is derived, followed by comparative simulation results in section 4. In section 5 the conclusion are proposed.\n2 Dynamical Model In this paper, the rocket-propelled segment, propelled by strap-on rocket boosters (SRBs) and central rocket engines (CREs), has been studied. The control torque includes contributions from the control input provided by the deflections of six engines, of which, the four SRBs make \u2019z\u2019 movement, and the two CREs make \u2019x\u2019 movement. According the control effect, the equivalent deflection angles ( ) ( ) ( ), ,z x z x z x\n\u03d5 \u03c8 \u03b3\u03b4 \u03b4 \u03b4 for pitch, yaw and roll control provided by SRBs and CREs are respectively given as:\n1\n1\n1\n1\n1 10 0 2 2\n1 10 0 2 2\n1 1 1 1 4 4 4 4\nz z\nz z z\nz z\nz\n\u03d5\n\u03c8\n\u03b3\n\u03b4 \u03b4\n\u03b4 \u03b4 \u03b4\n\u03b4 \u03b4\n\u03b4\n\u2212\n= = \u2212\n1\n1\n1\n1\n0 1 0 1 1 0 1 0\n1 1 1 1 2 2 2 2\nx x\nx x x\nx x\nx\n\u03d5\n\u03c8\n\u03b3\n\u03b4 \u03b4\n\u03b4 \u03b4 \u03b4\n\u03b4 \u03b4\n\u03b4\n\u2212\n= = \u2212\nThe configuration of Propulsive Rocket Engines (view from tail) is as following:\nThe Model, which discribe the motion of the vehicle, was desired to capture the physics charicters and to ensure the good results simultaneously. To illstrate the proplem, the rigid body equations of motion via small deviative linearization are as following :\n1 2 3 3 3 3 1\n1 2 3 3 3 3 2\n( )\n( )\nBYZ Z Z Z X X X X wp wq\nBZZ Z Z Z X X X X wp wq\nc c c c c c F c\nb b b b b b M b\n\u03d5 \u03d5 \u03d5 \u03d5 \u03d5 \u03d5 \u03d5 \u03d5 \u03d5 \u03d5 \u03d5\n\u03d5 \u03d5 \u03d5 \u03d5 \u03d5 \u03d5 \u03d5 \u03d5 \u03d5 \u03d5 \u03d5\n\u03b8 \u03b1 \u03b8 \u03b4 \u03b4 \u03b4 \u03b4 \u03b1 \u03b1\n\u03d5 \u03d5 \u03b1 \u03b4 \u03b4 \u03b4 \u03b4 \u03b1 \u03b1\n\u03d5 \u03b1 \u03b8\n\u2032\u2032 \u2032\u2032 \u2032\u0394 = \u0394 + \u0394 + + + + \u2212 + +\n\u2032\u2032 \u2032\u2032\u0394 + \u0394 + \u0394 + + + + = \u2212 +\n\u0394 = \u0394 + \u0394\n(1) '' '' '\n1 2 3 3 3 3 1 '' '' 1 2 3 3 3 3 2\n( )\n( ) z z z z x x x x BZ wp wq z z z z x x x x BY wp wq\nc c c c c c F c\nb b b b b b M b\n\u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8\n\u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8\n\u03c3 \u03b2 \u03c3 \u03b4 \u03b4 \u03b4 \u03b4 \u03b2 \u03b2\n\u03c8 \u03c8 \u03b2 \u03b4 \u03b4 \u03b4 \u03b4 \u03b2 \u03b2\n\u03c8 \u03b2 \u03c3\n\u0394 = \u0394 + \u0394 + + + + \u2212 + +\n\u0394 + \u0394 + \u0394 + + + + = \u2212 +\n\u0394 = \u0394 + \u0394\n(2) '' ''\n1 3 3 3 3z z z z x x x x BXd d d d d M\u03b3 \u03b3 \u03b3 \u03b3\u03b3 \u03b3 \u03b4 \u03b4 \u03b4 \u03b4\u0394 + \u0394 + + + + = (3)\n3 Dismc Design Due to rocket\u2019s axisymmetric property, pitch and yaw channels have the similar moving equations. Therefore, two channels are provided with exactly the same controller design process. Now we take pitch channel attitude control for example to illustrate the problem.\nThe control objective is to regulate the deviation angle to zero ( 0y\u03d5\u0394 = = ) to guarantee attitude angle tracking the given command signal. In order to facilite controller design, the engine inertia caused by the second derivative of control variables, disturbing power and moment and wind interference are together taken as the generalized disturbance. At the same time, the coefficients has been dealt with equivalently as following:\n1714 2012 24th Chinese Control and Decision Conference (CCDC)", + "3 3 3\n3 3 3\n3 3 3\n3 3 3\n2\n0.5\n0.5\nz x\nx z\nx x z z\nx z\nx x z z\nc c c c c c\nb b b b b b\n\u03d5 \u03d5 \u03d5\n\u03d5 \u03d5 \u03d5\n\u03d5 \u03d5 \u03d5 \u03d5 \u03d5 \u03d5\n\u03d5 \u03d5 \u03d5\n\u03d5 \u03d5 \u03d5 \u03d5 \u03d5 \u03d5\n\u03b4 \u03b4 \u03b4\n\u03b4 \u03b4 \u03b4\n\u03b4 \u03b4 \u03b4\n= =\n= +\n= +\n= +\n= +\nThus, the simplified system for controller design is as follow:\n1 2 3 3 3 1\n1 2 3 3 3 2\n( )\n( )\nBYZ z X X wp wq\nBZZ Z X X wp wq\nc c c c c F c\nb b b b b M b\n\u03d5 \u03d5 \u03d5 \u03d5 \u03d5 \u03d5 \u03d5 \u03d5\n\u03d5 \u03d5 \u03d5 \u03d5 \u03d5 \u03d5 \u03d5 \u03d5\n\u03d5\n\u03d5\n\u03b8 \u03b1 \u03b8 \u03b4 \u03b4 \u03b4 \u03b1 \u03b1\n\u03d5 \u03d5 \u03b1 \u03b4 \u03b4 \u03b4 \u03b1 \u03b1\n\u03d5 \u03b1 \u03b8\n\u2032\u2032 \u2032\u2032 \u2032\u0394 \u2212 \u0394 \u2212 \u0394 \u2212 \u2212 + +\n\u2032\u2032 \u2032\u2032\u0394 + \u0394 + \u0394 + =\u2212 \u2212 \u2212 +\n\u0394 = \u0394 + \u0394\n= +\n+\n(4) The system in locally equivalent differential Input-\nOutput form can be expressed in form as: 1 2\n2 3 * 3 \u02c6( , ) y y y y y f y u u\u03b3 \u03be = = = + +\n(5)\nwhere *\u03be is the generalized disturbance,and . disturbance. [ ] [ ]1 2 3\n2 1 1 3 1 2 1 1 2 2\n2 2 1 2 1 3 2 3\n3\n\u02c6\n\u02c6( , ) ( ) ( )\n[( ) ]\ny y y y\nf y u c c b y b c b c b y b c y c c b b c\nb\n\u03d5 \u03d5 \u03d5 \u03d5 \u03d5 \u03d5 \u03d5 \u03d5\n\u03d5 \u03d5 \u03d5 \u03d5 \u03d5 \u03d5 \u03d5 \u03d5\n\u03d5\n\u03d5 \u03d5 \u03d5\n\u03b4\n\u03b3\n= \u0394 \u0394 \u0394 =\n= \u2212 \u2212 + \u2212 \u2212\n+ + \u2212 +\n= \u2212\nHere, the controller, including control of dynamic continuous linear part and dynamic discontinuous nonlinear part, is designed as follows:\n1 2u u u= + (6) Where 1u is discontinuous in nature which is called dynamic integral control and 2u is continuous which stabilizes the system at the equilibrium point .In the next two subsections, the design of 1 2,u u is explored.\n3.1. Controller design for nonlinear part\nTo facilitate the design of 1u , system (5) can be expressed in an alternate form as follows:\n1 2\n2 3\n3 \u02c6( , , )\ny y y y y y u u u\u03c7 \u03be \u2217 = = = + +\n(7)\nWhere \u02c6 \u02c6( , , ) ( , ) ( 1)y u u f y u u\u03c7 \u03b3= + \u2212 and the In order to attain the desired performance and to compensate the uncertainties with reduced chattering , the integral sliding surface is designed in the following form (8).The integral sliding surface is designed in such way that the reaching phase is eliminated, which boosts the robustness against uncertainties from the very beginning.\n1 2\u03c3 \u03c3 \u03c3= + (8) Where 1\u03c3 is the conventional sliding surface and\n2\u03c3 is the integral term which can be determined. 1\u03c3 is defined by\n3\n1 3 1 , 1i i i c y c\u03c3 = = = (9)\nThe time derivative of (8) along (7) becomes: 1 2 2 3 1 2 2\u02c6( , , )c y c y y u u u u\u03c3 \u03c7 \u03c3 \u03be \u2217= + + + + + + Taking ( )2 1 2 2 3 2c y c y u\u03c3 = \u2212 + + (10)\nAnd the initial condition of 2 (0)\u03c3 is selected in such a way that satisfy the requirement (0) 0\u03c3 = Then,\n1\n2 1\n\u02c6( , , ) \u02c6( , ) ( 1) y u u u f y u u u \u03c3 \u03c7 \u03be \u03b3 \u03b3 \u03be \u2217 \u2217 = + + = + \u2212 + + (11)\nIt is known that a general sliding convergence condition: ( , , )k\u03c3 \u03bc \u03b5 \u03c3= \u2212\nwhich can make sure that the system is globally uniformly asymptotically stable.For this design procedure, it is defined as:\n( )sign k\u03b1 \u03b2 \u03c3 \u03b5 \u03c3 \u03c3 \u03c3 \u03c3= \u2212 \u2212 (12)\nwhere 0 , 1, 0, 0k\u03b1 \u03b2 \u03b5< < > > The new reaching law (12) has combined the fast approaching advantage of exponential reaching law and the smoothness into sliding surface advantage of power reaching law, which ensure the reaching speed and reduce the chattering at the same time. what\u2019s more, by identifying the distance between the system state and sliding surface, the approaching speed can be changed adaptively. By comparing (11) and (12), the expression of dynamic controller 1u is :\n( )1 2 1 \u02c6( , ) ( 1) ( )u y u u sign k\u03b1 \u03b2\n\u03d5 \u03b3 \u03b5 \u03c3 \u03c3 \u03c3 \u03c3 \u03b3\n= \u2212 + \u2212 + + (13)\nThe constant ,k\u03b5 are control gains which are selected according to uncertainty bounds; ,\u03b1 \u03b2 are adjustment factors which are selected according to the control performance. This control law enforce sliding mode along the manifold (8) from the very beginning. This completes the design of the discontinuous controller.\n3.2. Controller design of linear part It is known that when the system has reached the sliding mode along the manifold and stayed on it, the effect of controller 2u will stabilize the system at the equilibrium\npoint. Owing to the effect of 1u , the dynamics of original system on the manifold has been changed into the following form:\n1 2\n2 3\n3 2\ny y y y y u = = =\n(14)\nThis is a 3th subsystem which can be expressed in the following standard form:\n2y Ay Bu= +\nwhere 2 1 2 2 2 1\n1 1 1 2 1 1\n, O I O\nA B O O I\n\u00d7 \u00d7 \u00d7\n\u00d7 \u00d7 \u00d7\n= =\nand 3y R\u2208 is the state space vector of the subsystem.\n2012 24th Chinese Control and Decision Conference (CCDC) 1715" + ] + }, + { + "image_filename": "designv11_33_0001347_s10846-012-9725-2-Figure18-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001347_s10846-012-9725-2-Figure18-1.png", + "caption": "Fig. 18 Tested wing model; changing TC3", + "texts": [ + " Due to its smaller influence on the lift and drag, and the larger influence on the total mass and moment of inertia, for the large range of different materials, according to their density, these shapes might be extremely useful. Altering the conjunction position of the first and second segment, while keeping their sum unchanged, has an insignificant impact in the total lift and drag. At last, the changes of the tip chord length of the third segment are tested. For this reason, the third segment was modelled with a straight leading edge as shown in Fig. 18a. The changes in the lift and drag for different TC3 are given in Table 7. Another model with straight trailing edge (depicted in Fig. 18b) has caused a large decrease of the lift and an increase of the drag. As such, it won\u2019t have any useful impacts in the wing modelling for the spincopter. The length of the wing has evinced the largest influence on the total lift and drag. On the other hand, it causes a few negative contributions, such as an increase in mass and moment of inertia. To make the spincopter more controllable and due to the limitations of the actuators, it is useful to decrease the hovering angular velocity (derived in Section 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002455_1.4039395-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002455_1.4039395-Figure5-1.png", + "caption": "Fig. 5 Simulated optimal motion of the robots: (a) mobile manipulator with nonspherical wrist and (b) mobile manipulator with spherical wrist", + "texts": [ + " The obstacle avoidance added on average another 1 s for each task due to the two-phase planning, but bear in mind that this numerical simulations were performed in MATHEMATICA; it would be much faster if it was implemented for example in C. To prove this, we implement the algorithm in ROS written in C\u00fe\u00fe and use RViz to aid in the simulation. For the same set of task, it took on average only 45 ms to plan the optimal motion to approach each new task point without obstacles, while the obstacle avoidance added on average only another 45 ms for each task point. With such a short computational time, the approach has the capability to be applied as an online motion planner. Figure 5 shows the simulated motion obtained using the proposed SELH approach, where the robot employs redundancy to avoid obstacles and maximize manipulability. Figure 6 shows a snapshot of the ROS implementation, which includes a simple graphical user interface and simulation in RViz. Part of the ROS implementation includes the optimal fixed robot base placement capability described in Ref. [2] for completeness and can be found in ROS-I Asia Pacific. 5.2 Comparison With Existing Approaches. In this section, we qualitatively compare the results of our proposed SELH approach" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001446_1350650112460799-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001446_1350650112460799-Figure2-1.png", + "caption": "Figure 2. Schematic of the collapse of hemisphere: (a) deformable base case and (b) rigid base case.", + "texts": [ + " With the increase of interference, plastic deformation occurs and expands in the hemisphere. Plastic deformation results in larger contact area than the elastic Hertz solution. Therefore, equation (1) is the lower limit of overall contact area. The upper limit of contact area can be derived based on volume conservation. The volume of the hemisphere before contact is V1 \u00bc 2 3 R3 \u00f03\u00de When the interference is close to hemisphere radius, the whole hemisphere deforms plastically. And the hemisphere, either with deformable base or with rigid base, is collapsed, as shown in Figure 2. Since the top contact area is close to the base area, it is reasonable to assume the collapsed hemisphere to be a cylinder with height R ! and base area Au. The volume of the cylinder is V2 \u00bc Au R !\u00f0 \u00de \u00f04\u00de Based on volume conservation, the contact area of the collapsed hemisphere is Au \u00bc 2 R3 3 R !\u00f0 \u00de \u00f05\u00de As the collapsed hemisphere can never be a cylinder unless the interference reaches hemisphere radius, the contact area with cylinder hypothesis is the upper limit of overall contact area. The overall contact area of a hemisphere and a rigid plane is bounded by the two limits" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002110_whc.2017.7989929-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002110_whc.2017.7989929-Figure3-1.png", + "caption": "Figure 3. Deformation of variable-stiffness sheet containing piled-plastic films", + "texts": [ + " The thickness is determined using the number of piled fabrics, and is approximately 1 mm when 16 mesh sheets are piled and contained in a rubber bag. In a similar technology, the variable-stiffness sheets containing piled-plastic films have been used in haptic displays [21, 22]. Recently, these sheets have been employed in a membrane of soft robotic arms, which is referred to as layer jamming [23, 24]. The differences between the FJS and the layer-jamming sheet lie in their flexibility and stretchability. The piled-plastic films used in the layer-jamming sheet can bend only in one direction, and not in two directions simultaneously (Fig. 3). In contrast, the piled fabrics used in the FJS can fit on a spherical surface owing to their stretchability. The piled-plastic films can be made to fit on a spherical surface by employing a woven structure with strip-shaped films [25]. The variable-stiffness plates comprising plastic films capable of fitting on a spherical surface have been suggested [26]. However, the deformability of these mechanisms is limited by the size and nonstretchability of the films. In this paper, the structure and fundamental mechanical properties of the FJS are presented" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003469_j.engfailanal.2019.06.024-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003469_j.engfailanal.2019.06.024-Figure6-1.png", + "caption": "Fig. 6. Mechanical model of the main shaft while considering shell stiffness.", + "texts": [ + " = +e e e1 2 (1) \u23a1 \u23a3 \u23a4 \u23a6 \u23a1 \u23a3 \u23a4 \u23a6 = \u23a1 \u23a3 \u23a4 \u23a6 = \u23a1 \u23a3 \u23a4 \u23a6 \u03b1 \u03b1 \u03b1 \u03b1 e e \u03b8 \u03b8 e e e cos cos sin sin cos sin x y 1 2 1 2 1 2 (2) where \u03b11 is the angle between e1 and the x axis, \u03b12 is the angle between e2 and the x axis, \u03b8 is the angle between e and the x axis, e1, e2, eis the value of vector e1, e2, e, ex is the value of the e projected onto the x axis, ey is the value of the e projected onto the y axis. According to the geometric relations, we can know that. = + + \u2212e e e e e \u03b1 \u03b12 cos( )2 1 2 2 2 1 2 2 1 (3) = + + \u03b8 e \u03b1 e \u03b1 e \u03b1 e \u03b1 arctan sin sin cos cos 1 1 2 2 1 1 2 2 (4) As shown in Fig. 6, the main shaft is bent under the action of the offset mechanism, the main shaft is subjected to the force P of the eccentric mechanism. At the same time, the main shaft gives a counterforce P\u2032 to the shell through the eccentric ring, assuming the deflection angle of the shell is \u03b82. Therefore, the deflection curve of the main shaft, rotation angle of the bit and offset forceareall related to the stiffness of the main shaft, the stiffness of the shell and others. According to the geometric relations, we can know that" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003946_0954410019896430-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003946_0954410019896430-Figure1-1.png", + "caption": "Figure 1. The schematic diagram of DTSS.", + "texts": [ + " The remainder of this paper is organized as follows: in the following section, the mathematical model is developed by using Lagrange\u2019s equations; then, RBF neural network-based adaptive TSMC is proposed, and the finite-time convergence of DTSS is proven by Lyapunov theory; later, the effectiveness of the proposed controller is demonstrated by numerical simulations. Finally, conclusions are presented in the last section. In this section, the Lagrangian equation of the second kind is used to establish the deployment mathematical model of the DTSS, also the control problem is described. A typical schematic diagram of DTSS is shown in Figure 1, and the particles A, B, and C represent the main satellite, sub-satellite, and the mass center of DTSS, respectively. l is the tether length that has been deployed. Rc represents the vector from the center of the Earth to the mass center of the system. The coordinate system OXYZ is the geocentric orbit coordinate system with origin O located at the center of the Earth. The axis OX is along the spacecraft\u2019s radial direction pointing to the center of DTSS, the axis OZ is normal to the orbital plane, and the axis OY completes a right-handed coordinate system. In order to describe the attitude of system, the orbital coordinate system Cxyz and tether coordinate system Cxtytzt are defined with origin C coincident with the mass center of the system. The axes of coordinate systems Cxyz and OXYZ are parallel correspondingly. The axis Cxt points in the opposite direction of the tether tension, and the angles , between Cxyz and Cxtytzt shown in Figure 1, which are called in-plane and out-of-plane angle, indicate the location relation between the projection of the tether in the orbital plane with the direction of the vertical line, and the relationship between the tether and the orbit plane, respectively.29 The mathematical model in this paper is established in the orbital coordinate system. Before deriving the mathematical model, several reasonable assumptions can be made as shown below to simplify the modeling process. Assumption 1 1. The main satellite and sub-satellite can be considered as mass points during the tether deployment process, regardless of their shapes and attitude motions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002450_ascc.2017.8287450-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002450_ascc.2017.8287450-Figure1-1.png", + "caption": "Fig. 1. Schematics of RDIP", + "texts": [ + " \u03c3\u0307 is a continuous function in this case. This helps reducing chattering while keeping the main advantages of SMC. The main feature of this controller is that, it is applicable to the relative degree 1 systems. Outlined of the paper is as follows: Section II describes system description. Sliding surface is proposed in Section III and control development using STA. Subsequently simulation and experimental results are presented in Section IV and Section V. Finally Section VI concludes the paper. Schematic of RDIP is shown in Fig.1. RDIP comprises of three links of the pendulum. The Dc servo system provides necessary torque to Link 1 to control the whole system. Link 1 which is base arm is free to rotate in horizontal plane. A bottom link 2 connected to the rim of rotating arm. A top link 3 is connected to link 2 through hinge. Two links of the pendulum moves as an inverted pendulum in a plane perpendicular to the rotating arm. Each link has an optical encoder installed that measures the angular position of the links of the pendulum" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001425_1.4006251-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001425_1.4006251-Figure6-1.png", + "caption": "Fig. 6 Initial position", + "texts": [ + " In addition, a DA board (Interface Corp., PCI-3521), counter board (Interface Corp., PCI-6204), servo amplifier (Mitsubishi Corp., MR-J2S-40A), and servo motor (Mitsubishi Corp., HC-KFS43 (maximum torque: 3.8 Nm)) were used. The second and third links are not directly controllable because they do not have sensors or actuators at their joints. We consider the initial condition in which the configuration of the underactuated manipulator is in the rest position, i.e., the free links hang downward in the direction of gravitation in Fig. 6. We sinusoidally and horizontally excited the tip of the first link with frequencies in the range of from /2p\u00bc 3 (Hz) to /2p\u00bc 42 (Hz). At each excitation frequency, the relative angles of the second and third joints were measured using a high-speed camera (Fotron Corp., FASTCAM-APX RS 250K/250KC) under the condition in which the frame rate was 250 fps and the resolution for one frame was 1024 pixels 1024 pixels. The 1 pixel on the frame was corresponded to 0.2794 mm. 3.2 Experimental Results. Figures 7(a) and 7(b) show the change in the relative angles of free links at equilibrium points depending on the excitation frequency of the first link. The production of two types of pitchfork bifurcations in the third link, as shown in Figs. 7(a) and 7(b), is substantially different from the two-link system [14]. At the low frequency ( /2p\u00bc 6 Hz), indicated by the black circles in Figs. 7(a) and 7(b), free links remain in the rest position, as shown in Fig. 6. When the excitation frequency /2p is from 6 Hz to 8 Hz, the manipulator has two types of stable equilibrium points, as indicated by the blue and red squares in Figs. 7(a) and 7(b), where the combination of h2 and h3 at the equilibrium configuration of the manipulator at each excitation frequency is expressed by the combination of blue square plots in Figs. 7(a) and 7(b) or red square plots in them. Figure 8 shows the configuration of the three-link underactuated manipulator. This configuration corresponds to the equilibrium points indicated by the arrows in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003882_s11071-019-05343-5-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003882_s11071-019-05343-5-Figure3-1.png", + "caption": "Fig. 3 Schematic diagram of forces acting on cage", + "texts": [ + " The differential equation [24] for the vibration of the j th ceramic ball can be described as Fbx j + F\u03b7o j cos\u03b1o j \u2212 F\u03b7i j cos\u03b1i j + T\u03b7i jcos\u03b1i j \u2212 T\u03b7o j cos\u03b1o j + Qi j sin \u03b1i j \u2212 Qo j sin \u03b1o j + Qcx j \u2212 P\u03b7 j = mb x\u0308b j Fby j + F\u03be i j \u2212 F\u03beo j + T\u03beo j \u2212 T\u03be i j + Gby j + Qcy j = mb y\u0308b j Fbz j \u2212 F\u03b7o j sin \u03b1o j + FR\u03b7i j sin \u03b1i j \u2212 T\u03b7i j sin \u03b1i j + T\u03b7o j sin \u03b1o j + Qi j cos\u03b1i j \u2212 Qo j cos\u03b1o j \u2212 Gbz j + Qcz j \u2212 P\u03be j = mb z\u0308b j [(T\u03be i j \u2212 F\u03be i j ) cos\u03b1i j + (T\u03beo j \u2212 F\u03beo j ) cos\u03b1o j + Qczj \u2212 P\u03be j ]DW 2 \u2212 Jx \u03c9\u0307x j = Ib\u03c9\u0307bx j (T\u03b7i j \u2212 F\u03b7i j + T\u03b7o j \u2212 F\u03b7o j ) DW 2 \u2212 Jy\u03c9\u0307y j = Ib\u03c9\u0307by j + Ib\u03c9bz j \u03b8\u0307b j [(T\u03be i j \u2212 F\u03be i j ) sin \u03b1i j + (T\u03beo j \u2212 F\u03beo j ) sin \u03b1o j + Qcx j \u2212 P\u03b7 j ]DW 2 \u2212 Jz\u03c9\u0307z j = Ib\u03c9\u0307bz j + Ib\u03c9by j \u03b8\u0307b j (1) where DW is the diameter of the ceramic ball; mb is its mass; x\u0308b j , y\u0308b j , and z\u0308b j are the displacement accelerations of the barycentre of the j th ball along the direction of each coordinate axis in the coordinate system {O; X,Y, Z};\u03c9bx j , \u03c9by j , and \u03c9bz j are the angular velocities of the j th ball in the coordinate system {O; X,Y, Z}; \u03c9\u0307bx j , \u03c9\u0307by j , and \u03c9\u0307bz j are the angular accelerations of the j th ball in the coordinate system {O; X,Y, Z}; \u03b8\u0307b j is the orbit speed of the j th ball in the coordinate system {O; X,Y, Z}; and Ib are the moments of inertia of the ball in the coordinate system {O; X,Y, Z}. 2.2 Differential equations of vibrations on cage During the operation of the bearing, the cage contacts only the ball, and friction and impact are generated. The forces of the j th ceramic ball on the cage of the bearing are shown in Fig. 3. In Fig. 3, ec is the eccentricity between origin Oc in the cage coordinate system {Oc; Xc,Yc, Zc} and origin O in the inertial coordinate system {O; X,Y, Z};\u03c6c is the deflection angle of the coordinate system {Oc; Yc, Zc} relative to {O; Y, Z};\u03c6p j is the position angle of the j th pocket; \u03c6b j is the position angle of the j th ceramic ball relative to the cage; \u03c6\u2032 b j is the cosine angle of\u03c6b j ; Fcy , and Fcz are the components of the hydrodynamic force acting on the cage; and Mcx is the friction moment acting on the cage" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000908_0020-7403(65)90021-4-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000908_0020-7403(65)90021-4-Figure9-1.png", + "caption": "FIG. 9.", + "texts": [ + " However , collapse would occur because of mater ia l failure r a the r t han the loss of init ial elastic s tabi l i ty . 676 S . J . BR~TV~C In general, eigenvalue buckl ing of axial ly hyperstat ic cont inuous frames is possible and i t takes the same general form as tha t of an isostatic frame. However, condit ions of 0 Fro. 8. Stable axial ly hyperstat ic rigidly jointed frame. geometrical compatibility o fflexural contractions must be used in addi t ion to those already employed. This is demons t ra ted in the following example. Example (3) Consider the simple frame in Fig. 9. In the analysis, the members are assumed axial ly rigid. They are rigidly jo inted a t B and p inned at A, D, C. The frame is symmetr ica l abou t the loading axis. The sloping members of length L are inclined at 60 \u00b0 to the horizontal . Example 3 ; buckl ing of an axially hyperst.atie rigidly jointed frame. The parameters A,, 2 z, A 8 and the body rotat ions \u00a21, ~b2, ~b3 nms t satisfy the following linearized compat ib i l i ty conditions, given by equat ions (8) in section 2, ~/(3) ~b x = A, -- 222 (253) ~/(3) ~bz = 221 --2z (25b) 4/(3) \u00a23 = 421 - 328 (25c) and the additional compatibility condit ion, A , + 2 ~ - - ~ 2 3 = 0 ( 2 6 ) (There are four independent compat ibi l i ty condit ions in this case" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003852_j.mechmachtheory.2019.103628-Figure12-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003852_j.mechmachtheory.2019.103628-Figure12-1.png", + "caption": "Fig. 12. The interference status at different grinding positions.", + "texts": [ + "5 mm allowance, the initial parameters of grinding wheel are obtained as shown in Table 3 . Based on the initial parameters, the tool path of gear grinding can be planned, and the interference test of non-working side is carried out. It can be seen that the non-working side of the grinding wheel has obvious interference with the other end of the opposite tooth surface when grinding both ends. The interference status between the non-working side of the grinding wheel and the opposite tooth surface at different grinding positions are shown in Fig. 12 , and the detail interference values are shown in Fig. 13 . It can be seen that the interference is more serious when grinding the heel of concave side and the toe of convex side. According to the method described in Section 6 , such a large amount of interference in the non-working side can only be avoided by tilting the axis of the grinding wheel. The tilt angle of grinding wheel axis increases gradually. Then the pressure angle and the top radius of the grinding wheel are modified according to Eqs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003826_mdp2.112-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003826_mdp2.112-Figure2-1.png", + "caption": "FIGURE 2 Parts orientation in XY plane", + "texts": [ + "com/journal/mdp2 1 biocompatible and can be used in custom\u2010made disposable guiding elements used surgery. Alumide is an EOS material consisting of polyamide PA 12 and aluminum particles, in a mechanical mixture. The main purpose of introducing aluminum in the mixture is to enhance the stiffness of the parts and to control the shape stability. The sintering process was conducted on EOS Formiga P100 (EOS GmbH Electro Optical Systems) for constructing the parts designed according to ASTM D 5045\u201099.13 In the Figure 1, the process scheme of Formiga P100 machine is presented, while in the Figure 2, the orientation of the parts (0\u00b0, 45\u00b0, and 90\u00b0) and the fabrication bedding example can be observed. The process parameters were set identical for both materials, and the main values used are laser power 25 W, laser speed 2000 mm/min, scan spacing 0.25 mm, and building chamber temperature 170\u00b0C. Six parts for each orientation angle were manufactured using polyamide and Alumide, a total number of 36 sample being obtained. After cooling and post\u2010processing all parts, the mechanical test was conducted on 5\u2010kN Zwick machine (Figure 3)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002076_optim.2017.7975006-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002076_optim.2017.7975006-Figure1-1.png", + "caption": "Fig. 1. Computation domain of the 3D electromagnetic field associated with the motor operation", + "texts": [ + " Based on the experience in the finite element analysis of stator short-circuit faults with 2D models [5, 6, 8, 9, 11, 14, 19-21, 24, 26, 28, 29] and replying the investigations with the 3D finite element model of a squirrel-cage induction motor described in [34], this paper presents results related an onephase short-circuit and a two-phase short-circuit faults. The first short-circuit is generated by the contact of the two terminals of an elementary coil, the second correspond to the contact of the output terminal of an elementary coil belonging to a phase and the input terminal of an elementary coil that belongs to another phase. The possibility of early detection of these short-circuits is studied. A two poles squirrel-cage induction motor of 7.5 kW, 3 x 380 V, 50 Hz supplied, represents the support of different numerical applications. Fig. 1 reflects the 3D computation domain of the electromagnetic field developed with the Flux software [38], including a part of the infinite box region. This domain contains half of the motor geometry, delimitated by the transversal symmetry plane xOy, where the center point of the motor O[0, 0, 0] is placed. There are considered the main assemblies of the motor - the rotor squirrel-cage, which is a region of solid conductor type, the stator and rotor magnetic cores, which are magnetic nonlinear and nonconductive regions and the stranded coils of the stator winding, which are regions with constant current density in the cross-section", + "05 ms of the time step considered by the step by step in time domain computation ensures a good enough precision of the numerical results for the evaluations of the amplitude of harmonics of the currents and of the magnetic field until 2000 Hz. An electric circuit is associated to the field model of the motor in order to take into account the three phase 380 V rms voltage supply of the motor and to model the short-circuits through low resistance resistors. The upper image in Fig. 2 highlights in yellow the first elementary coil of the phase U, red colored in Fig. 1, which is the object of the one-phase short-circuit. This short-circuit is modeled through the resistor Rshc in the lower image of the associated circuit, resistor which connects the input and output terminals of the short-circuited coil. A very high resistance 107 \u2126 of the Rshc resistor is considered for the healthy (HE) state of the motor operation. The values corresponding to faulty states with increasing fault severity FA1, FA2, FA3 and FA4 are 90 \u2126, 9 \u2126, 5 \u2126, and 0.9 \u2126 respectively. For the first three faulty states, the post processing of the results of the correspondent Flu3D applications show the values of the current in the short- circuited elementary coil of phase U are lower than the motor rated current" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002349_s12289-017-1388-x-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002349_s12289-017-1388-x-Figure6-1.png", + "caption": "Fig. 6 Rhinoceros 5 output geometry", + "texts": [ + " Boolean geometry intersection and contact length calculation At this point of the calculation, all 2D geometries have been created, extruded along the z-direction and located in the right position considering a specific slice of the ring in a specific instant of the process. The geometries data and the relevant position are now fed to BModel implementation^ section, Fig. 2, where the Boolean intersection between the outer area of the ring and the main roll and between the inner area of the ring and the mandrel are carried out, as shown in Fig. 6. The Boolean intersection allows determining the contact area as well as the extremum points of contact (BModel validation^ section), on both main roll side and mandrel side, Fig. 7. As previously anticipated, the chosen criteria to calculate a unique solution is to consider the projection of the contact arc between tools and ring to have the same value on both main roll and mandrel side. By applying this assumption, the algorithm varies \u0394ss until the coordinate of the extremum points of contact have the same value on both sides", + " However, the computational time of the combined CADanalytical model proposed in this paper is directly linked to the number of slices, as it is hereafter explained. After the calculation of the geometry of the ring, the resulting data are exported and submitted to the CAD algorithm implemented in the Grasshopper plug-in of Rhinoceros 5, as also shown in BContact geometry estimation utilizing Grasshopper^ section of Fig. 2, and the calculation is iterated until the projection of the contact arc onto the y-axis, Fig. 6, is Case FY [9] FQ [10] FP [6] Fauthors FFEM eY eQ eP eauthors same on both mandrel and main roll side. This value is considered the result for the projection of the contact arc between the tools of the considered slice in the considered round of the process. The time required for one single computation, thus for the calculation of the projection of the contact arc for one single slice, is averagely 2.5 s on an i7\u20136700 3.4GHz processor with 16GB of RAM installed. The small deviations between the calculation times are caused by the proximity of the initial guess solution to the final solution, which is not preconditioned" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002165_j.jmapro.2017.08.008-Figure15-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002165_j.jmapro.2017.08.008-Figure15-1.png", + "caption": "Fig. 15. (a) Example 4, (b and c) support point clustering, (d) Support model interface surface, (e) Support generated by proposed method, and (f) Support with model.", + "texts": [ + " Comparing with the commercial machine, he proposed methodology shows 14% improvement of total build ime (TBT) where 53% savings of support material as compared in able 4. Fig. 14(c\u2013e) shows the comparison of the support structure etween commercial support generators and proposed methodlogy. The comparison with respect to support volume, contour umber, and build time between commercial support generators nd proposed methodology is shown in Table 5. The implementation sequence of the proposed methodology on xample 4 is shown in Fig. 15(a\u2013e). The normal vector parameter ange is calculated as 0 \u2212 0.85, which is divided into four groups o generate coherent region as shown in Fig. 15(b\u2013c). This object as a volume of 4030 cubic mm with a build height of 11.9 mm. 30 164 13 144 Following the proposed technique, it generates 1051 cubic mm of support volume. Comparing with the commercial machine, the proposed methodology shows 26% improvement of total build time where 64% savings of support material as compared in Table 6. Even though, both build time and support volume has reduced in the proposed method, but number of contour is higher than the commercial software for this object" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003787_j.ymssp.2019.106415-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003787_j.ymssp.2019.106415-Figure1-1.png", + "caption": "Fig. 1. Example of a real vane segment of a low pressure turbine module for aeronautical applications. Courtesy of GE Avio.", + "texts": [ + " In particular, it will be shown how the GSI method is exploited to handle with the huge number of contact DoFs in order to evaluate the forced responses of the system that otherwise would be expensive to predict from a computational point of view. The goodness of the proposed methodology is quantified both in terms of accuracy of the solution and time costs savings on the calculation of non-linear forced responses by comparison with benchmark results obtained by the Policontact software [24]. Deeper considerations will be finally carried out on the convenience of properly design the hook joints in order to extend the engine components\u2019 life improving the safety margins. A typical example of a real stator vane segment is shown in Fig. 1. This represents the fundamental \u2018sector\u2019 of a stator wheel located in between two subsequent rotor stages. Each vane segment usually exhibits three lap joints through which the interaction with the neighbors components occurs: - interlocking joint: located at the inner radius of the sector it allows the coupling of neighboring segments. When all the segments are assembled resulting in a full stator wheel, the design interference at the interlocking guarantees a certain level of normal preload. During vibration friction onset causes energy dissipation, which is responsible of a response amplitude decrease in resonance conditions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001502_wcica.2012.6358164-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001502_wcica.2012.6358164-Figure1-1.png", + "caption": "Fig. 1. sketch of NSVOW", + "texts": [ + "0468 \u221a T (3){ \u03c1 = pbi exp(\u2212g0/ (RTbi) (H \u2212 Hbi))/RTbi \u03c1 = pbi(1 + Lbi/Tbi(H \u2212 Hbi))\u2212g0/(RLbi\u22121)/RTbi (4) where H is geopotential height, h is geometric height, r0 is radius of the Earth, R is air constant, g0 is sea level acceleration of gravity, \u03c1 is atmospheric density, T is temperature, Vs is speed of sound. pbi, Tbi, Hbi and Lbi are respectively initial atmospheric pressure, initial temperature, initial geopotential height and vertical temperature gradient of a level. Remark 1: The definitions of all symbols in the (1)-(4) are given in [9]. In the following development of modeling for the NSVOW, the atmospheric density \u03c1 and speed of sound Vs will be used to obtain dynamic pressure and flight Mach number. A ketch of the NSVOW is shown in Fig. 1. This vehicle has a wing skew angle that is variable from 0 to and 60 deg, and it is equipped with right aileron, left aileron, right elevator, left elevator and rudder. The deflection of aileron, which is denoted by \u03b4a, generates the rolling moment. The deflection of elevator, which is denoted by \u03b4e, generates the pitching moment. And the deflection of rudder, which is denoted by \u03b4r, generates the yawing moment. \u03b4a, \u03b4e > 0 means the trailing edge is displaced downwards, and \u03b4r > 0 means the trailing edge is displaced to the left" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001381_kem.490.237-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001381_kem.490.237-Figure6-1.png", + "caption": "Fig. 6. Generating system of the pinion \u2013 initial position", + "texts": [], + "surrounding_texts": [ + "Due to simultaneous cutting of the both sides of the tooth space (pinion and gear), the bearing contact bias occurs after assembling of a transmission made in that way. Such an effect is undesirable since it causes loud work of the transmission and non-uniform transfer of the motion which leads to the faster fatigue wear. In order to remove the bias the helical motion is applied during pinion finishing. It is realized by the axial offset of the fixed headstock of the workpiece connected with the generating gear rotation. The generating system of the pinion is the technological hypoid gear obtained thanks to the hypoid offset of the workpiece\u2019s axis by the E value with respect to the cradle\u2019s axis. The cradle with inclined toolhead (tilt) creates the bevel generating gear. For the pinion cutting the head\u2019s axis inclination is applied in order to compensate the difference between the pressure angle and tool profile so in the initial position the c S system is rotated with respect to the d S system by the angle of j around d X axis. The head\u2019s axis inclination angle (in degrees and minutes) and the value of the hypoid technological offset E is determined on the basis of the gear geometry analysis and calculation cards. Positive values mean down-shift by the pinion with left inclination line direction of a tooth or up-shift by the pinion with right inclination line direction of a tooth. The negative value means upshift by the pinion with left inclination line direction of a tooth or down-shift by the pinion with right inclination line direction of a tooth. Mathematical model of tooth flank surface Mathematical notation of the tooth flank is apparent from the equation of the surface of action of the tool, kinematics, and accepted treatment technological system. The following discussion presents the side of the tooth surface obtained with the use of a technological system with a bevel generating gear - this is the case more generally in comparison with a ring generating gear. While processing the envelope, the equation of the tooth flank, which is the bounding surface of the utility, is determined from the system of equations [1, 2, 6]. This system includes the equation of the family of tool surfaces and the equation of meshing, resulting from the method for determining the kinematic envelope: ( ) ( ) , , , , 0 t t t t t t s s \u03b8 \u03c8 \u03b8 \u03c8 \u22c5 = 1 t1 1 1 r n v (1) where: ( ), ,t t ts \u03b8 \u03c81r - determining the vector function of the family tool surfaces system bounded with treating pinion ( 1S ), 1 n - the unit normal vector defined in 1S , ( ), ,t t ts \u03b8 \u03c8t1 1v - the relative velocity vector defined in 1S . Based on the defined technological model, the family of tool surfaces is determined as follows: ( ) ( ) ( ), , ,t t t t t ts s\u03b8 \u03c8 \u03c8 \u03b8= \u22c51 1t tr M r (2) where: ( ),t ts \u03b8tr - vector equation of the tool surface referred to the system associated with the tool, tS , ,t ts \u03b8 - curvilinear coordinates surface form, ( )t\u03c81tM - the transformation matrix being the product of a transformation matrix representing the rotations and translations of homogeneous coordinate systems included in the technological gear model t\u03c8 - parameter of motion (in this case - angle of the cradle rotation). Vector equation of the tool surface as a function of curvilinear coordinates ,t ts \u03b8 shows the relationship (3), which involves the processing of the active side of the tooth. ( ) ( ) ( ) cos sin , sin sin cos t wk t wk t t t wk t wk t wk r s s r s s \u03b8 \u03b1 \u03b8 \u03b8 \u03b1 \u03b1 + = + \u2212 tr (3) where: wk\u03b1 - the angle of the external blades, wkr - the radius of the cutterhead. Based on the model of technological gear it is designated a family of the tool surfaces according to equation (2), for which it determines conversion matrix equation (4). ( ) ( )( ) ( )1t t t\u03c8 \u03c8 \u03c8 \u03c8= \u22c5 \u22c5 \u22c5 \u22c5 \u22c5 \u22c5 \u22c51t 1w wr rh hm mk kc cd dgM M M M M M M M M (4) where: ( )t\u03c81tM - conversion matrix, ij M - the elementary transformation matrices representing the rotations and translations of homogeneous coordinate systems, ij - subscript indicating the direction of the transformation from system jS to system iS . Technological gear model is also used to determine (in the chosen system, for example 1(S or )mS a unit normal vector and relative velocity. The normal vector to the surface of the tool sets at any of the predefined layouts. In this way we can get the components of meshing equation. Solving equations (1) by eliminating one of the variables such as: ts from the meshing equation and then substituting into the equation of the family of tool surfaces ( )1 , ,t t ts \u03b8 \u03c8r we can obtain the tooth flank surface equation in two-parametric form (5). ( ) ( )( ), , , ,t t t t t t ts\u03b8 \u03c8 \u03b8 \u03c8 \u03b8 \u03c8=1 1r r (5) where: ( ),t t\u03b8 \u03c81r - the equation of the pinion tooth surfaces in the two-parametric form, ( ),t t ts \u03b8 \u03c8 - the variable ts in the function of other parameters. Model of technological gear is designed to create the flank surface of the gear and pinion teeth, which will be used for the analysis of meshing for constructional spiral bevel gear. In order to obtain a tooth surface as the family of tool surfaces it should be designated in the system rigidly bounded with cutting pinion S1 , which represents the envelope to family of surfaces of the cut gear tooth ( )\u03a3 1 . Family of tool surfaces is shown in the following way: ( ( )) ( ( ( ) ( )) )\u03c8 \u03c8 \u03c8\u2190 \u2190 \u2190 \u2190 \u2190 \u2190 \u2190 \u2190 \u2190= \u22c5 \u22c5 \u22c5 \u2212 + \u22c5 \u22c5 \u22c5 \u2212 \u2212(t) (t) 1 1 w 1 t w r r h h m m k t c d d t t k c r hr L L L T L L L r T T (6) In this equation the vector record (t) 1 r , (t) m r , (t) gr , concerns a family tool surfaces ( )\u03a3 t (as evidenced by the superscript ( t )), set respectively in the cutting gear system S1 , the basic system (stationary) and the tool system. Other signs are bounded with the transformations of coordinate systems used in the model of technological gear. The applied rotation and translation matrixes of the coordinate system are as follows: cos i 0 sin i 0 1 0 sin i 0 cos i , \u2190 = \u2212 d g L (7) cos j sin j 0 sin j cos j 0 0 0 1 , \u2190 = \u2212 c d L (8) U cos q U sin q 0 , \u2190 \u2212 \u22c5 = \u2212 \u22c5 k c T (9) cos sin 0 ( ) sin cos 0 0 0 1 , \u03c8 \u03c8 \u03c8 \u03c8 \u03c8\u2190 \u2212 = t t m k t t t L (10) 1 B1 1 0 A X p ( ) ,\u03c8 \u2190 = \u2212 + h m t" + ] + }, + { + "image_filename": "designv11_33_0000546_j.1460-2695.2010.01540.x-Figure13-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000546_j.1460-2695.2010.01540.x-Figure13-1.png", + "caption": "Fig. 13 Crack tip radial and angular coordinate from the first propagation point (X1).", + "texts": [ + " A positive kurtosis distribution (leptokurtic) exhibits a sharper peak, thus indicating that the majority of the evaluations remain close to the mean. On the other hand, a platykurtic or negative kutosis distribution has a more rounded peak. The first error analysis compares the crack tip locations. In order to have an unbiased evaluation, the crack tip coordinates are now redefined from the initial crack tip position at a0 (arel = 0.0): r becomes the radial position and \u03b8 turns into the angle between the initial crack segment and the straight line defined by propagation points arel = 0.0 and arel = i (Fig. 13). Figures 14 and 15 present the descriptor evolution for the radial position and angular position, respectively. Two series of data are plot in the figures. The first series (\u25e6) considers the full crack propagation path of both the prediction model and BE model. The results show a maximum mean error around \u22120.2 mm on the radial position and \u22121.1\u25e6 on the angular position. The standard deviation becomes more important for the second half of the propagation path with highest values of 0.75 mm and 7" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002968_15440478.2018.1541774-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002968_15440478.2018.1541774-Figure1-1.png", + "caption": "Figure 1. Fiber bridging mechanism in FMLs (Khan, Alderliesten, and Benedictus 2009).", + "texts": [ + " In FMLs, the fatigue crack initiates at the exterior metallic layers since the constituent with high stiffness attracts more load during cyclic loading. Due to the presence of fibers in the laminates, part of the load is transferred from the metallic layers to the fibers. The fibers remain intact during the crack growth in the aluminum layers, thereby restraining the crack propagation. The presence of fibers in the laminates hinders the crack tip opening, consequently resulting in a reduction of stress intensity in the aluminum layers. Because of the fiber bridging mechanism as shown in Figure 1, the crack growth rate in FMLs is relatively lower than monolithic aluminum (Khan, Alderliesten, and Benedictus 2009). Composite as the core in FMLs has been a well-known material since the past few decades (Sivakumar, Ng, and Salmi 2016). Indeed, composites alone have the excellent potential to substitute the metallic materials in a wide variety of applications such as the fuselage of airplanes and interior door panels of vehicles. Nevertheless, the poor impact resistance and damage tolerance of carbon fiber reinforced thermoset composite materials have impeded their usage in engineering applications (Sinmaz\u00e7elik et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002540_jifs-169406-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002540_jifs-169406-Figure1-1.png", + "caption": "Fig. 1. Dynamic finite element model.", + "texts": [ + " In the following geometric model, some factors which have relatively little influence on the internal stress distribution and deformation of the bearing are simplified, such as chamfer, oil groove, the axial and radial play of the bearing, etc. Combining 2D CAD drawing software with the finite element analysis software ANSYS-LS/DYNA, the tapered roller bearing model is established. The crown design of roller generatrix is completed in the 2D CAD, which is then imported into the finite element analysis software. Finally the whole tapered roller bearing design is finished in the LS-DYNA. Dynamic finite element model of tapered roller bearing is shown in Fig. 1. Selecting the outer ring and cage for rigid body material model, rollers and inner ring for linear elastic material model, this is conducive to simulate some practical problems and improve the speed of calculation. Material of inner and outer rings and rollers of the tapered rolling bearing is GCr15, and the cage is made of brass. The parameters of material model are shown in Table 1. Solid164 unit and shell163 unit are selected as the finite unit types. Finite element mesh method generally includes free meshing, mapped meshing and sweep meshing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003882_s11071-019-05343-5-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003882_s11071-019-05343-5-Figure6-1.png", + "caption": "Fig. 6 Schematic diagram of coordinate systems for piston sound source", + "texts": [ + " It is assumed that the end surface of the ring is divided into many infinitely small surfaces, dS, and each is treated as a point sound source. The intensity of the point sound source is denoted as dQ0 = udS. Since the vibration of only half a sphere contributes to the half-space sound field, the sound pressure emitted from the end surface of the ring can be expressed as [46] p = \u222b \u222b S j \u03c10c0k 2\u03c0h ue j (\u03c9t\u2212kh)dS (21) where S is the end surface area of the ring and h is the distance from observation point to point sound source. As shown in Fig. 6, the sound field is rotational symmetric relative to the X -axis of the piston sound centre. In the sound field, the point P is placed in the XOZ plane, and analysed by geometry, the relational expressions can be obtained as follows: h = \u221a r2 + l2 \u2212 2rl cos( r , l) (22) cos( r , l) = sin \u03b8 cos\u03d5 (23) where l is the distance between the surface, dS, and coordinate origin, O , and \u03d5 is the azimuth angle of the surface, dS. When the distance from the observation point to the centre of the piston sound source is far greater than the radius of the piston, h \u2248 r \u2212 l sin \u03b8 cos\u03d5 (24) By replacing Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002062_978-3-319-61431-1_2-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002062_978-3-319-61431-1_2-Figure1-1.png", + "caption": "Fig. 1. General configuration of a single cable.", + "texts": [ + " The final geometry of the cable will be given as the composition of three steps: unwinding, shaping and rotation. Let us consider a single cable #k operated by a winder #k. The cable is tangent to the winder at point Wk and has an end-point denoted Pk. The unwinded portion of the cable is the planar curve between Wk and Pk of length lk. Let Fb = (Ob,xb,yb) and Fk = (Wk,xk,yk) denote respectively the fixed global reference frame and the local reference frame attached to the winder #k. The position of Wk and the orientation of the cable at Wk are defined by (xWk , yWk ) and \u03d5k respectively as indicated in Fig. 1. 1 http://icube-avr.unistra.fr/fr/index.php/Planar cable robot with non straight cables. From an initial configuration where the cable is straight along the xk direction, let us now consider a small displacement of the cable that alters the cable shape but preserves its point of tangency Wk on the winder. In this elementary displacement, a point of coordinates (x, 0) with x \u2208 [0, lk] is moved to the point Mk of coordinates (xMk = x+\u03b4xMk , yMk = \u03b4yMk ) in the local frame Fk. Finally, the coordinates (xk, yk) of the point Mk expressed in the global frame Fb can be obtained using an homogeneous transformation as \u23a1 \u23a3 xk yk 1 \u23a4 \u23a6 = \u23a1 \u23a3 cos \u03d5k \u2212 sin \u03d5k xWk sin \u03d5k cos \u03d5k yWk 0 0 1 \u23a4 \u23a6 \u23a1 \u23a3 xMk yMk 1 \u23a4 \u23a6 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002588_mwent.2018.8337172-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002588_mwent.2018.8337172-Figure3-1.png", + "caption": "Fig. 3. The accelerometer and gyroscope accuracy checking experiment scheme.", + "texts": [ + " Gyroscope sensor output data is almost always shifted relative to the zero level due to it structural features and offset value depends on various factors such as temperature, rotation speed etc. As a result error of tilt angle estimation is accumulated during integration process what causes slow \u201crunaway\u201d of tilt angle value. Thus the main disadvantage of tilt angle estimating by using accelerometer is the distortions caused by acting external forces and the main disadvantage of gyro sensor is a zero drift. To show these properties of sensors the following experiment was carried out. MPU6050 sensor was placed on the rod and the rod was rigidly fixed to the shaft of motor (Fig. 3). The shaft was rotated by applying the voltage to the motor coils. In first stage 10% of maximum voltage level was applied and 50% in the second. The output data of accelerometer, gyro sensor and motor encoder were captured. Comparing measuring result obtained by MPU6050 sensor to the data obtained by encoder (in this experiment encoder is a reference sensor) it is possible to estimate the accuracy of the accelerometer and gyro angle measurement. From the experiment result (Fig. 4 \u2013 7) one can see that angle estimation obtained by the accelerometer has significant distortion in the second case (50% of voltage) due to more fast start of motor shaft rotation and hence greater angular acceleration" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000854_robot.2010.5509330-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000854_robot.2010.5509330-Figure7-1.png", + "caption": "Fig. 7. Experiment set-up a) Semi-cylindrical soft fingertip and the distributed ridges on the surface. b) Entire view of the experimental apparatuses.", + "texts": [ + " In the next section, experiment results will be shown to validate these simulation results. To validate the proposed model, we conducted experiment on an objective soft fingertip. This is a semi-cylindrical soft fingertip which has the radius of 20 mm, and the thickness equals to 4 mm. This soft fingertip was made from polyurethane rubber (KE-12, Exseal, Japan) after an 8-hour curing phase at room temperature. On the outer surface of the fingertip, there are many ridges distributed uniformly (Fig. 7(a)). Each ridge has the shape of 90o arc of a circle which has diameter of 0.5 mm. Distance of two continuous arcs (or ridges)\u2019 centers is 1 mm. These distribution of ridges represent epidermis ridges on the human\u2019s fingertips, and free ends of cantilevers in the model reported in section II. Fig. 7(b) illustrates set-up of the experiment. The soft fingertip is attached on a 2-DOF (degree of freedom) xz-motorized linear stage (XMSG615 and ZMSG413, Misumi, Japan) to give vertical and horizontal translation of the fingertip on the flat rigid plane. There are two kinds of force/moment sensors employed to measure force/moment acting on the fingertip during the experiment. The 3-DOF load cell PD3-32-10105 (Nitta, Japan) is fixed upon the fingertip to measure the total forces (normal force and friction force) acting on the fingertip during sliding motion" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002109_1350650117723484-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002109_1350650117723484-Figure3-1.png", + "caption": "Figure 3. Relationship between the center of the ball and groove curvatures.", + "texts": [], + "surrounding_texts": [ + "Bearing characteristics, constant preload, bearing stiffness, thermal effects, lubricant film thickness Date received: 5 April 2017; accepted: 26 June 2017" + ] + }, + { + "image_filename": "designv11_33_0000926_amr.443-444.160-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000926_amr.443-444.160-Figure1-1.png", + "caption": "Figure 1. The schematic diagram of milling", + "texts": [ + " The workpiece dimension was 80mm\u00d750mm\u00d730 mm. All the machining experiments were carried out on a three axis Mikron HSM 800 high speed milling center with an ITNC 530 controller, whose rotational speed range from 0 to 3600 r/min. During the experiments, 10 mm uncoated cemented carbide milling cutters with four teeth K44 were used. Cutting tool relief angle was 10\u00b0 and helix angle was 30\u00b0. The milling mode was down milling. The schematic diagram of milling applied by the experiments is shown in Figure 1. The milling experiments focus on the surface residual stress of workpiece under different cutting conditions. Three cutting tool rake angles were select as 4\u00b0, 8\u00b0 and 14\u00b0. Meanwhile, three different cooling methods including emulsion liquid, oil mist and dry were selected. The influence of cutting conditions to surface residual stress in high-speed milling of titanium alloy TC11 was investigated by four-factor and three-level orthogonal method. Three cutting speeds were selected as 251 m/min, 314 m/min, and 377m/min" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003262_tmag.2019.2900447-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003262_tmag.2019.2900447-Figure1-1.png", + "caption": "Fig. 1. Rack\u2013pinion gear structure. (a) Mechanical gear. (b) Magnetic gear.", + "texts": [ + "ieee.org. Digital Object Identifier 10.1109/TMAG.2019.2900447 within a few seconds with high accuracy. In addition, it is possible to understand the underlying physical phenomenon by approximating the governing equation and reduce the number of modifications to the design variables [4], [5]. This paper investigates the torque characteristics of a magnetic rack-and-pinion gear using an analytical method. Furthermore, we fabricated a rack\u2013pinion gear and verified its feasibility through experiments. Fig. 1(a) and (b) shows the mechanical and magnetic structures of the mechanical and magnetic rack\u2013pinion gears, respectively. The mechanical rack\u2013pinion gear has a structure in which the gear teeth are physically engaged with each other, and the linear motion is converted into rotational motion. As shown in Fig. 1(b), the magnetic rack\u2013pinion gear based on PMs can act as a mechanical gear without any physical contact. Fig. 2 shows the model used for the analytical method. As magnetic rack\u2013pinion gears have both linear and circular structures, the air-gap length is not constant, unlike in a typical machine. Hence, it is necessary to consider the variation in the air-gap length in the magnetic field analysis process. Using the analysis model, shown in Fig. 2, we can consider the varying air-gap length from the center. In addition, we can analyze the magnetic field and force using the Cartesian coordinate system without requiring complex methods such as coordinate transformation [6]. 0018-9464 \u00a9 2019 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. In this paper, we analyze the magnetic rack\u2013pinion gear as shown in Fig. 1(b). The pinion gear with a circular structure is made up of 16 pole PMs, and the width of the linear gear of the rack is the same as that of the pinion magnets, making the magnetic gear ratio 1:1. Thus, magnetic rack\u2013pinion gears based on PMs operate in a magnetically coupled manner and help convert linear motion into rotational motion. In addition, as the gear ratio is 1:1, the relationship between the linear and rotational speeds can be expressed as follows: V = r \u00d7 RPM \u00d7 0.10472. (1) Fig. 2 shows the analytical model for the magnetic field distribution analysis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000759_detc2011-48502-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000759_detc2011-48502-Figure1-1.png", + "caption": "FIGURE 1. MACHINE TOOL MODEL COUPLED TO AN UNGROUNDED NES.", + "texts": [ + " To properly understand the suppression mechanisms that appear similar to those in the previous aeroelastic applications, numerical bifurcation analysis are performed by utilizing DDEBIFTOOL [18]. The CxA technique is also utilized for analytical study of TET mechanisms (i.e., resonance captures). Finally, applications of an NES to a practical machine tool model are introduced. In this section we provide a quick review of SDOF machine tool dynamics model \u2013 i.e., permanent contact model [3] and contact loss model [5]; then introduce an ungrounded NES to the machine tool model. Figure 1 provides a conceptual depiction of turning process and its mathematical modeling coupled to an ungrounded NES. First, neglecting the NES in the modeling, we can write the equation of motion for the SDOF machine tool as x\u0308+2\u03b6 \u03c9nx\u0307+\u03c9 2 n x =\u2212 1 m \u2206Fx (1) where \u03c9n is the linearized natural frequency of the undamped free oscillating system; w, the cutting width; and \u03b6 = c/(2m\u03c9n), the damping factor. \u2206Fx is the cutting force variation which can be calculated by curve fitting of data generated by cutting tests; and it is a function of chip thickness f such that Fx( f ) = Kw f \u03b1 where K is related with test technical parameters, and \u03b1 is the so-called cutting force exponent (In this work we set \u03b1 = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001627_1.4001668-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001627_1.4001668-Figure2-1.png", + "caption": "Fig. 2 An example of a wedge artifact \u202023\u2021", + "texts": [ + " This is a generating type of tooth lead easurement method and, in this study, this measurement method s investigated. In this method, the relation between the amount dz f probe movement in the axial direction and the rotational angle of the measured gear is expressed in the following equation: rmd = dz tan m 1 he difference between the theoretical helix form and the actual orm is expressed in the direction of the line tangential to the base ircle of the gear dotted line in Fig. 1 . 2.2 Evaluation Using Wedge Artifact. Figure 2 shows an xample of the wedge artifact. Plane S1 is measured instead of the ooth flank by the lead measurement method. In the evaluation of he accuracy of the gear-measuring instrument using the wedge rtifact, first, the virtual measurement result without error theoetical measurement curve of the wedge artifact is calculated, here an ideal gear-measuring instrument without any errors is ssumed. This theoretical lead measurement curve is compared ith the actual lead measurement result, which is output from the 71006-2 / Vol", + " However, if the lead measurement circle radius rm becomes larger, the inclination of the theoretical measurement curve around the origin does not change. Figure 9 indicates that some tens of millimeters of change in the lead measurement circle radius rm is acceptable. Therefore, a specific wedge artifact can be applied to a wide variety of lead measurement circle radii. This is an important advantage of the wedge artifact. In the case of the conventional helicoid artifact, it is impossible to evaluate the lead measurement accuracy under such a wide variety of lead measurement circle radii. 4.2 Experiment. The manufactured wedge artifact cf. Fig. 2 is measured using a gear-measuring instrument while changing the helix angle m and the lead measurement circle radius rm. The inclination angle is 60.16582 deg, which is measured using a coordinate measuring machine. The uncertainty for this measurement is 2.1 10\u22125 deg 23 . The flatness of the lead measurement plane S1 is about 0.5 m. This flatness value applies to the whole plane. Actually, in the tooth lead measurement, only a limited area of plane S1 is used, and therefore it is thought that the flatness is actually much smaller than 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003987_sa47457.2019.8938077-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003987_sa47457.2019.8938077-Figure3-1.png", + "caption": "Figure 3. Device design with double-sensor tip [19]", + "texts": [ + " where the pipe section is too small, the medium or flow is too aggressive or the accuracy requirements can be relaxed. The use of thermowells, however, is inherently problematic. It is complicated and expensive to engineer a thermowell, a circumstance which has recently been exacerbated by increased safety standards [16], [18]. As a result, many important temperature measurements in industry are just not done. An approach to solve the problem has been presented in [19]. The device setup is given in Figure 3. Two point measurements of temperature are carried out close to the surface of a pipe, but with slightly different distance. An extrapolation concept of the temperature field can then be used to calculate the surface temperature of the pipe. The specific development here was a considerable shift of the second sensor close to the main sensor and close to the surface. This led to a dramatic reduction of response time. The large potential of dynamic response time reduction was calculated before for a simplified model system", + " The latter is assumed to be just a stainless-steel rod which is assumed to touch the surface to be measured with its tip at position x = 0. Moreover, two T-measurement elements are attached to the thermometer rod at distances x1 and x2 from the tip. The picture shows that response time can be decreased by more than one order of magnitude when moving the second sensor close to the first one. As a consequence, in the real device a second thermometer inset has been introduced, leading to the design in Figure 3. The increased performance created the very special situation that a non-invasive sensor of modern type can actually compete with the traditional high-accuracy invasive sensors, a circumstance that may fundamentally change the industrial temperature measurement practice. Probably for this reason, the device was shortlisted for the prestigious \u201cHermes Award\u201d at Hanover Fair 2019 [20]. From a user perspective, apart from the striking feature of measuring non-invasively, the device is very close to previous standard temperature devices (handling in practice, certification, etc)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003552_1350650119866037-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003552_1350650119866037-Figure4-1.png", + "caption": "Figure 4. Meshing process of the gear.", + "texts": [ + " Based on the rough surface contact model, the contact pressure distribution on the tooth surface during the meshing process can be calculated according to the geometric parameters, mechanical properties and the distribution parameters of the asperity, and the stress field in the subsurface can be further calculated. After the stress distribution in the subsurface of the gear is obtained, considering that the gear teeth are mostly in time-varying multi-axis complex stress state, the multi-axis fatigue life model proposed by Liu and Mahadevan9 can be used to calculate the fatigue life of the gear during the meshing process. Simplification of gear meshing process Figure 4 is the schematic diagram of gear meshing process. As we can see, G is a point on the meshing line, N1 is the meshing point of engaging in and N2 is the meshing point of engaging out, and B1 and B2 are the boundary points of the single and double teeth engagement area. On the meshing line, B1B2 is a single tooth meshing area and the rest is a double teeth meshing area. According to the geometric relationship of gear meshing, it is easy to get R1 \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi R2 G 0:5d1 cos \u00f0 \u00de 2 q \u00f09\u00de R2 \u00bc 0:5d1 sin \u00fe0:5d2 sin R1 \u00f010\u00de R \u00bc R1R2 R1 \u00fe R2 \u00f011\u00de In the equation, RG is the distance between G and the center of the driving wheel; R1 and R2 are the radius of curvature of the driving wheel and the driven wheel, respectively; d1 and d2 are the pitch circle diameter of the driving wheel and the driven wheel, respectively; is the pressure angle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000256_robot.2009.5152179-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000256_robot.2009.5152179-Figure5-1.png", + "caption": "Fig. 5. Parameters Definition, Left: Swing, Right: Heelstrike", + "texts": [ + " In this paper, we will particularly study the effect of the first energy feedback on the Virtual Slope Walking stability. B. Model and Assumptions To simplify the problem in the largest extend, this paper adopts the model that composed of a mass point trunk and two massless telescopic legs. The extending and shortening of the legs are completely equivalent to the bending and unbending of the knee joints since the legs are massless. Some parameters used in this paper are defined as the following (Fig. 5). \u03b8 angle between the two legs during swing, when the swing leg is before the stance leg, 0\u03b8 > , at heelstrike, 0\u03b8 \u03b8= , where 0 0\u03b8 > ; r length of the stance leg during swing, before heelstrike it is r \u2212 , after heelstrike the stance leg and swing leg exchange, and it become r + ; \u03c6 inclination angle of the stance leg during swing, when the stance is after the zenith, 0\u03c6 > , before heelstrike, \u03c6 \u03c6 \u2212 = , where 0\u03c6 \u2212 > , after heelstrike it is \u03c6 \u03c6 + = , where 0\u03c6 + < ; \u03c9 angular velocity of the stance leg during swing, \u03c9 \u03c6\u2032= , before heelstrike it is \u03c9 \u2212 , where \u03c9 \u03c6 \u2212 \u2212 \u2032= , after heelstrike it is\u03c9 + , where \u03c9 \u03c6 + + \u2032= , the mean of \u03c9 in a walking step is \u03c9 , where 0 T\u03c9 \u03b8= ; F force on the stance leg during swing; M mass of the trunk; g acceleration of gravity; T step period; t time variable in a walking step", + " Let r r r \u2212 + \u0394 = \u2212 , (6) becomes 0 0 sin( ) sin( ) (2 ) ct c Q Mg r Q Mg T r M r r r\u03c9 \u03c9 \u03c6 \u03bb\u03b8 \u03c6 \u03bb\u03b8 \u03bb \u03c9 + + + = \u2212 + \u0394 = \u2212 + \u0394 \u2212 + \u0394 \u0394 \u23a7 \u23a8 \u23a9 (20) Since r\u0394 is far smaller then r + , its second-order items are neglectable, thus we have 0 0 sin( ) sin( ) 2 ct c Q Mg r Q Mg T r M r r\u03c9 \u03c9 \u03c6 \u03bb\u03b8 \u03c6 \u03bb\u03b8 \u03bb \u03c9 + + + = \u2212 + \u0394 = \u2212 + \u0394 \u2212 \u0394 \u23a7 \u23a8 \u23a9 (21) Similarly, we neglect the second-order items of r\u0394 , (15) becomes 1n n nA B t\u03c9 \u03c9 + \u0394 = \u0394 + \u0394 (22) where 2 20 02 2 0 02 cos ( sin( ) 2 ) cos sin( ) A g T r r r r r B g r r \u03b8 \u03c6 \u03bb\u03b8 \u03bb \u03c9 \u03c9 \u03c9 \u03b8 \u03c6 \u03bb\u03b8 + + + + + + = \u2212 + \u0394 + \u0394 \u2212 = \u2212 + \u0394 \u23a7 \u23aa\u23aa \u23a8 \u23aa \u23aa\u23a9 Substitute (2), (9) into c rE E= and neglect the second-order items of r\u0394 , we have 2 2 2 2 0 0 1 cos( ) tan 2 Mg r M r r Mr\u03c6 \u03bb\u03b8 \u03c9 \u03c9 \u03b8 + + + + \u0394 \u2212 \u0394 = (23) From (23), we have 2 0 02 tan 2( cos( ) 1) r r g r \u03b8 \u03c6 \u03bb\u03b8 \u03c9 + + + \u0394 = + \u2212 (24) Substitute (24) into (21) and use 0 T\u03c9 \u03b8= , we have 3 2 20 03 2 20 0 0 tan 2( 1) ( 2) tan 2( 1) ct c Mr F TQ G F Mr TQ G\u03c9 \u03b8 \u03b8 \u03b8 \u03bb\u03b8 \u03b8 + + = \u2212 \u2212 = \u2212 \u23a7 \u23aa \u23aa \u23aa \u23a8 \u23aa \u23aa \u23aa\u23a9 (25) where 2 02 0 2 02 0 0 2 2 0 0 2 0 sin( ) cos( ) 1 cos ( cos ) sin 2( 1) tan 2( 1) gT F r gT G r K H K G K G \u03c6 \u03bb\u03b8 \u03b8 \u03c6 \u03bb\u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 + + + + = \u2212 + = + \u2212 = \u2212 + \u2212 = + \u2212 \u23a7 \u23aa \u23aa \u23aa \u23aa \u23aa \u23a8 \u23aa \u23aa \u23aa \u23aa \u23aa\u23a9 Substitute (24) into (22), and combine it with (18) and 0 T\u03c9 \u03b8= , the eigenvalues of the Jacobi matrix in (19) are 2 1,2 ( ) 4 2 A D A D BC \u03bb + \u00b1 \u2212 + = (26) where 2 20 0 2 2 0 0 2 2 0 ( 2) sin cos 2( 1) sin 2 ( 1) 1 F A G F B T G T C D \u03bb\u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u2212 = + \u2212 = \u2212 = = \u23a7 \u23aa \u23aa \u23aa \u23aa \u23a8 \u23aa \u23aa \u23aa \u23aa \u23a9 In (26), F, G, H, K are the same as those in (25). In (25) and (26), all ctQ , cQ \u03c9 and 1,2\u03bb are determined by six parameters: M , r + , \u03bb , 0\u03b8 , \u03c6 + , and T , while five of them are independent. Here, we choose \u03c6 + being determined by the others. From geometry relation in Fig. 5 right, we have 0 0 1 ( ) sin tan cos r r r r \u03b8 \u03c6 \u03b8\u2212 + + + = + \u0394 = + (27) Substitute (24) into (27), we have 2 0 2 0 0 0 2 0 tan 1 1 sin tan cos 2( cos( ) 1) gT T r \u03b8 \u03b8 \u03b1 \u03b8 \u03c6 \u03bb \u03b8 + + = \u2212 + + \u2212 (28) By satisfying (28), we can calculate the numerical value of \u03c6 + according to \u03bb , 0\u03b8 , and T . Then by setting M and r + with the body parameters of our robot, there are three parameters to be studied, which are \u03bb , 0\u03b8 , and T . A. Robot Our planar robot Stepper-2D is shown in Fig. 15. It is 36.8cm by height and 780g by weight" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003893_0954406219888240-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003893_0954406219888240-Figure4-1.png", + "caption": "Figure 4. Notation diagram of the universal joint.", + "texts": [ + " The mapping relation of two coordinate systems denoted by R1 can be presented by two Y-X Euler angles ( around V-axis and around U-axis), which is expressed as R1 \u00bc cos sin sin sin cos 0 cos sin sin cos sin cos cos 2 64 3 75 \u00f01\u00de Since the two hinges of the universal joint are symmetrical about the orthogonal plane where two cross shafts are located, only the quarter section is studied under the assumption that the bottom hinge is fixed. Therefore, two sides (Side A and Side B) and two faces (Face 1 and Face 2) are marked to facilitate the study, as shown in Figure 4. Sides A and B belong to the bottom and upper hinges, respectively. Faces 1 and 2 denote the side and inner planes of the bottom hinge, respectively. Two end points of Sides A and B are A1, A2 and B1, B2, respectively. Two projections of the bottom hinge in the o-UV and o-UW planes are shown in Figures 5 and 6, respectively, where the projections of Faces 1 and 2 correspond to Sides 1 and 2. Accordingly, the values of two end points of Side A in the o-UVW coordinate system are A1 \u00bc d 2 , c 2 , 0 T , A2 \u00bc d 2 , c 2 , L4 T \u00f02\u00de In the o-uvw coordinate system, the values of two end points of Side B are B01 \u00bc c 2 , d 2 , 0 T , B02 \u00bc c 2 , d 2 ,L4 T \u00f03\u00de According to the mapping relation of two coordinate systems, the values of two end points of Side B in the o-UVW coordinate system can be solved as B1 \u00bc R1B 0 1 \u00bc \u00bdB1U,B1V,B1W T, B2 \u00bc R1B 0 2 \u00bc \u00bdB2U,B2V,B2W T \u00f04\u00de where B1U \u00bc c 2 cos \u00fe d 2 sin sin , B1V \u00bc d 2 cos , B1W \u00bc c 2 sin \u00fe d 2 cos sin , B2U \u00bc c 2 cos \u00fe d 2 sin sin \u00fe L4 sin cos , B2V \u00bc d 2 cos L4 sin , B2W \u00bc c 2 sin \u00fe d 2 cos sin \u00fe L4 cos cos " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000698_s207510871201004x-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000698_s207510871201004x-Figure1-1.png", + "caption": "Fig. 1. Geometric representation of the motion control problem for a ship on a preset path.", + "texts": [ + ", this group of equations is enveloped by a certain feed back coupling. If the equations that are not intercon nected by feedback coupling have a triangular struc ture [7, 8], the backstepping method allows us to mod ify the already known intermediate control law and DOI: 10.1134/S207510871201004X 42 GYROSCOPY AND NAVIGATION Vol. 3 No. 1 2012 DOVGOBROD include one more equation into a group of equations enveloped by feedback coupling, etc. DESIGNING ADAPTIVE CONTROL ALGORITHM Assume that it is necessary that the center of gravity of the ship p (see Fig. 1), move as close as possible to the preset path represented as a smooth curve. The predetermined path is described by the vector function pd(u), depending on parameter u. The vector function pd(u) must have coordinate wise second order deriva tives. The absolute value of the ship current speed rel ative to the inertial coordinate system {a} is denoted by U, and the velocity vector direction by \u03c6. Let B be a representing point moving along the preset path at a certain variable velocity UB at \u03c6B direction in the coor dinate system {a}" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000908_0020-7403(65)90021-4-Figure13-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000908_0020-7403(65)90021-4-Figure13-1.png", + "caption": "FIG. 13. Effect of geometrical imperfections on the post-critical equilibrium paths of pin-jointed frameworks.", + "texts": [], + "surrounding_texts": [ + "Wmax = 1 - ( tan 2 7+- ~)~ 0 \u00b0t + . . . (45) where it is a s sumed t h a t 7+ < 30 \u00b0. I t is s ignificant t h a t in the ease of p in- jo in ted f rames the reduct ion in the peak load p a r a m e t e r from the crit ical e igen-value is p ropor t iona l to 0 \u00b0t, which is in con t ras t to the behaviour of con t inuous frames, when th is reduc t ion depends on the order of 0\u00b0i. Also, in p in- jo in ted f rames the peak occurs a t larger d is tor t ion , i.e. for 0 oc 00+, t h a n in cont inuous f rameworks for which 0 ~ 00t. W h e n 7+> 30 \u00b0 the p a t h has a m i n i m u m for nega t ive values of 0. Then the increase in the m i n i m u m value of the load The theory of elastica in the non-linear behaviour of plane frameworks 681 parameter from the lowest eigenvalue is proportional to ~l, i.e. wml n = 1 + ( t a n 2 y - ~)~ 8 a! + . . . (46) CONCLUSIONS The closed expressions for flexural equilibrium of prismatic members make it possible either in their exact or approximate form to refer any buckled member within a rigidly or pin-jointed framework to one of the plane elastic curves. The character of the curve (inflexional, non-inflexional, compressive or tensile) is determined by the appropriate choice of the \"shape\" function and a single distortion parameter. In general, the non-linear relations are found to yield eigenvalue solutions which are identical with those obtained earlier from the linear theory by several authors. In its first approximation the more accurate theory is in agreement with the approximate analysis developed in Ref. 13, but ultimately, the non-linear expressions of the elastica make the theory more accurate. The theory is found equally useful in analysing axially hyperstatic continuous and pin-jointed frameworks which exhibit eigenvalue buckling. However, by suitable bracing, loss of initial elastic stability in such frameworks may be prevented altogether. The equilibrium paths of geometrically imperfect continuous frames, studied by means of a model show a reduction in the peak load compared to the 682 S . J . BRITVEC critical eigenvalue which is proportional to the initial imperfection (nondimensional) raised to the power of \u00bd. In contrast to this, such a reduction in the case of pin-jointed frames, is proportional to the initial imperfection raised to the power of ~. The fl)rms of the equilibrium paths studied by means of the model are found to be in agreement with the experimental observations in Ref. 13. Finally, it can bc concluded on the basis of the generalized statical analysis tha t a l l continuous frames, whether of the portal or braced type, h)ade(l centrally at the joints by static conservative loads, represent unstable structures which may in general develop dynamic buckling in the critical region of loading. Acknou'ledgeme~lts--A major par t of this work was carried ()tit between the years 1960 and 1964 while the au thor was Assistant Professor in the Depa r tmen t of St ructura l Engineer ing at Cornell Univers i ty . The prepara t ion and wri t ing of' the paper were done at, H a r v a r d Univers i ty and the support for this was received from the Office of Nava l Research, Contract NONR-1866(02). The au thor wishes to acknowledge the visit ing lectureship ex tended to him by the l )epar tn len t of\" Civil Engineer ing tit Univers i ty College, London, in the smnmer of 1962, when sonic of the theoreti(:al work on the behaviour of portal-I ype frames was earried out. He wouhl like to thank Profbssor A. H. Chilver for a discussion re la t ive to tha t paiq of t, he work. Final ly , thanks arc due to Professors B. Budiansky and J . L. Sanders. ,lr., of Harvar( t Univers i ty , for reading t.he manuscr ipt . A 1' P E N D I N Compressive members distorted by small emt-couples and a large axial force (P shape function dis tor t ion pa ramete r % = (cp, _ q)~)2 1 1 sin2(1)~-sin2d)~ l ( c o s ( P l - c o s % = ((P'-q) ' )2 [8 - -16 c l , ~ - d ) s - 2 , , ....... 7~(~(Ps (Ps) ~ ] #ql = ((P, - ~,) cos q), [ 1 21 sin 2(l)i--sinq),_.(1)~ 2_~.j) b3,. = ~(~P~-q) i ) \\ 3 - (I L COS cos (I)~ -- COS ~)i ~'~ = sin (D,+ .................... q)i - (P~ 7 cos (P, - cos q)j 3 cos (P i - cos (PJ ( s in2 (P i_s in2 (P j ) e3~ = :ig ..... i~-~:--% 32 ( ~ , - %)2-- _ _l [c._o:,~ (P,-_eo_s ._~s _ ~ sin q), cos -~ O, 3 \\ q)i - (P~ ] 1 1 s in2(Pz-s in2(P j 1 [ e o s d ) , - e o s ( D ~ 2 d., = ~i-~ ........ ~ , \u00b1 i ~ . . . . . 2~- ~, % l 5 1 sin 2(D i - sin 2d)~ 1 sin 4(I)~--sin 40~ d4 = - ~ + 6 - 1 6 . . . . . d ) , - ~ + 2-1-6 ~ (I)~- (1)~ pin (eo \u00a2, - ~ ~ q) , - \u00a2 , _ - ,~. ~ o ~ --- -~ ,=-~] --! +,~ q , % q , , - ~ ~ .... + ~ -~ ,_% / e~ = (q~,- q~,)' [1 1 sin 2 ~ - s ! n . 2 ~ ] 4 + ~ ~p~- (p~ J [ 1 1 sin 2 ~ - s i n 2q)~ ,,, = ((t,, - % ) ~ t 3 s 4 - ~:-,'i2 . . . . 4~i - (E- l sin 4(D,-s in 4(I)~ 1 / s in2q) , - - a in2~\u00a2 , ~] -2 .162 ( p , Z ~ , \" - i 6 ' \u00a2 , - % t ) J The coef~cie.nts d.o and e~ are always positive. Only (I)~ need be negative. The theo ry of e las t ica in the n o n - l i n e a r b e h a v i o u r of p l ane f r ameworks 683 Inflexional tensile members distorted by small end-couples and a large axial.force q\" shape f u n c t i o n a ' d i s to r t i on p a r a m e t e r a o = (TI - ~t',) ~ [ 1 s inh 2 t F i - s i n h 2tI's l ~sinhXt',zs!nh.._q'~i2 ] a2 - ( t l '~- t t ' i )2 ~ + 1 6 ~t\"~-~'~ - 2 \\ ~t',.-~t\"~ ] J b,i = ( q ' , - t F ~ ) s inh ~ ' i b a t = ]~(q ' i -~t . ' j ) ( ~ + ~ s inh2q\"~- - s inh2~F~)s inh~ t \", tt\" i _ q ' j s inh t F , - s inh t|.j exi = cosh ~ ' i . . . . 7 s inh ~'~ -- s inh ~'s 3 s inh 2~ ' i - s inh 2tF~ (sinh tt', - s inh tU~) \u00a2'~' - 4 s - - % - % 32 i v - , - % ) ' 1 (s inh ~ ' i - s inh + ~ \u2022 ~l'i) a + -2~ sinh2 tt'i cosh ~t.' i tF~ - tl ' j 1 1 s i n h 2 ~ F , - s i n h 2 ~ F j 1 ( s inh tF , . . -_s inh~ ' j l2 d.,_ = 4 + ~ \u2022 tFz_W, ~ - ,~ ~ ' , - W j ,t 5 1 s inh 2~\" i - s inh 2Wi 1 s inh 4~\" i - s inh 4tI 'j d4 384 6\" 16 . . . . ~t\"(- ~ ' j 2-162 -~'~ ~ ' ~ 1 (s inh 2~ '~ - - s inh 2~'~\\ ~ 1 / s inh t F ~ - s i n h \" l s inh 2~ ' , --_ s..'.m_h 2~ ' i {sinh t F , - s inh \" 1 (s inh tF , - sinh ~',~* - - 3 2 - t I\" , - tF~ ~ ~ '~-~F~ t t ~ ) ~ - 8 t t ' ~ - -T , . . . . . / . . , / l l s i n h 2 t F ~ - s i n h 2 t F ~ ) e~ = ( ~ t ' ~ - ~ ) [ - ~ + ~ ,t. _~t . ~ 1 1 s inh 2~'~ - s inh 2tF\u00a2 + 1 s inh 4q ' i - s inh 4~['_~ e~ = ( ~ ' _ ~ - ) z 384 96 ~'~--q~'~ 2.16 z q '~--q '~ 1 (sin_h 2~ ' , - s inh 2~\u00a2~ *] + 16 ~ \\ ~'~ -- ~'~ ] J d, a n d e, are a lways pos i t ive ; on ly ~'~ need be nega t ive . (Note : a 0 and a , are pos i t ive w h e n the axia l force is tensi le .) Pin-jointed compressive members X = ~ , a 0 =~r~ ,a2 = r r2 /8 ,b~ = b~ . . . . . 0, c~ = l , c~ = c~ . . . . . 0, d 2 = ~[,d~ = - ~ , e 2 = ~ / 4 , e~ = ~ r ~ / 3 8 4 . Non-inflexional tensile members I n th i s case no inf lexions occur in the elast ic curve a n d P is tens i le (negat ive) . E q u a t i o n ( la) is modif ied accord ing ly a n d it is easi ly shown t h a t c~ is i m a g i n a r y (c~ is the slope a t a n inf lexion). Since i q) s in [ ( ~ ) ] = i s i n h (.~.) E q u a t i o n ( ld ) gives, ~ _0-_~ s i n h ( ~ - ) s i n h ( ~ ' ) = s in 2 684 S . J . BRITVEC Because the R.H.S. is real, it follows that a = - i e and ~P = i~F \", where e and H\" are also real. The same transformations can be deduced rigorously on integrating the modified equation (la). R E F E R E N C E S 1. A. E. H. LOVE, A Treatise on the Mathematical Theory of Elasticity. Cambridge University Press (1920). 2. i . V. SOUTHWELL, An Introductio~ to the Theory of Elasticity. Clarendon Press, Oxford (1936). 3. S. TIMOSHENKO, Theory of Elastic Stability. McGraw-Hill, New York (1936). 4. E. Cr~WALLA, Die Stabilitiit lotrecht belasteter Rechteckrahmen, Dcr Bauingenieur, January (1938). 5. H. ZIE('LER, Advanc. Appl. Mech. 4 (1956). 6. C. E. PEARSON, Quart. J. Appl. ~lath. 14 (1956). 7. S. J. BRITVEC, \"The Post-Buckling Behaviour of Frames\", Ph.D. Dissertation, Cambridge University (1960). 8. l~. FRmCH-FAY, _F/exib/e Bars. Butterworths, London (1962). 9. W. T. KOITER, Elastic Stability a~td Post-Bucklil~g Behaviour in Non-Li~ear Problems. University of Wisconsin Press (1963) (\"On tho Stability of Elastic Equil ibrium\", Thesis (in Dutch), Delft, 1945). 10. ,I. M. T. Tr~o.~IsoN, J . Mech. Phys. Solids, 11,000 (1962). 11. S. ,I. BRITVEC, Int. J. ,ilech. Sci. 5, 447 (1963). 12. S. J. BaXTVEC, \"Overall Stability of Pin-jointed Frameworks after the onset of Elastic Buckling\", Ingenieur Arch iv, Band 32, Springer-Verlag, Berlin (1963). 13. S. ,l. BRrTVEC and A. H. CHILVER, \"Elastic Buckling of Rigidly-jointed Frames\", J. Engng Mech. 89, 217 ASCE (1963). 14. J . G. EISLEY, Appl. illech. Re~'. (Sept. 1963)." + ] + }, + { + "image_filename": "designv11_33_0003233_s12206-019-0209-1-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003233_s12206-019-0209-1-Figure3-1.png", + "caption": "Fig. 3. Gearbox compound fault test-rig.", + "texts": [ + " And a fitness function is proposed to evaluate the optimization effect of the algorithm which is shown as follows: 1( , ) 1 1 ( , )fit i i vc i i F C a C = - + g g , (40) where ( , ) [0,1]vc i ia C \u00ceg is the cross validation accuracy of SVM using the parameters Ci and \u03b3i. The flow chart for optimization of SVM parameters by WOA is shown as Fig. 2. An experiment of compound fault diagnosis in a parallel shaft gearbox was performed to verify the effectiveness of the proposed method. The test rig of compound fault was composed of a 4 kW stepless speed motor, a parallel shaft gearbox, a magnetic powder brake, a speed torque sensor and signal acquisition system. Fig. 3 presents the test rig. The magnetic particle brake can provide the load for the experiment. The test gearbox is equipped with a set of four gears and three shafts. In this paper, the output shaft gear is called gear 1, the big gear of intermediate shaft is called gear 2, the pinion gear of intermediate shaft is called gear 3, the gear of input shaft is called gear 4, the structure of the test gearbox is shown as Fig. 4. Fig. 4 also displays the tooth number of each gear for the test gearbox. The data is collected by an acceleration sensor which is vertically mounted on the outside of the gearbox, and the specific layout of the sensors is the location of the digital symbol used in Fig", + " Therefore, the recognition results demonstrate that the proposed method is more competitive compared to other methods. Furthermore, it is also proved that the method proposed in this paper still performs well for variable condition signals. In addition, to illustrate the difference between compound fault and single fault in classification, the traditional EMDANN method was applied to deal with the single fault data of gearbox. The single fault test was also completed in the test rig as shown in Fig. 3. Two single faults of 10 mm gear broken tooth fault and 8 mm gear crack fault are preset in experiment. Similarly, the loads of three working conditions are 10 Nm, 15 Nm and 20 Nm, respectively. For other details of the experiment, please refer to Ref. [47]. There are 20 samples in each working condition and fault state. The data contains the signals of gearbox in three states: Normal state, broken tooth fault and crack fault. So there are 180 samples, half of them are training samples and the other half are testing samples" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002870_joe.2018.8345-Figure10-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002870_joe.2018.8345-Figure10-1.png", + "caption": "Fig. 10 Modal simulation of stator core (a) 2nd f = 7113 Hz, (b) 2nd f = 7312 Hz, (c) 3rd f = 7747 Hz, (d) 3rd f = 7747 Hz, (e) 4th f = 12,221 Hz, (f) 4th f = 15,623 Hz", + "texts": [ + " The frequency of the changing electromagnetic force is directly related to the rotational speed of the rotor and the number of slots. As shown in Fig. 9, the peak value of 750 Hz in flux density spectrum was generated by the rotor of 30 slots at 1500 rpm speed. The properties of the materials of each part of the stator are shown in Table 3. The difference between the stators of the four motors is the pole arc coefficient of the stator permanent magnets. The variation range is 0.722\u20130.8. The range of variation is not large, so the modalities of the four motors are similar. The main mode of motor C is shown in Fig. 10. Previous research has shown that the electromagnetic vibration of the motor is mainly caused by the vibration of the stator; while the influence of the rotor is small. The Ansys Workbench was used to Table 1 Basic configuration of model motor Motor Poles Slot number Pole arc Slot per pole \u03b1 A 4 36 0.778 9 7 B 4 36 0.722 9 6.5 C 4 30 0.8 7.5 6 D 4 30 0.733 7.5 5.5 Table 2 Structural parameters of motor C Parameter Value Parameter Value, mm pole pairs 2 rotor out diameter 94 slot number 30 slot inner diameter 59 stator outer diameter 130 mm shaft diameter 35 stator inner diameter 104 mm stator length 220 Table 3 Material properties Material Density, kg/m3 Young's modulus, GPa Poisson's ratio structure steel 7850 200 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001045_jsea.2010.312134-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001045_jsea.2010.312134-Figure3-1.png", + "caption": "Figure 3. The motor assembly [1].", + "texts": [ + " The element is defined by ten nodes having three degrees of freedom at each node. The element has plasticity, creep, swelling, stress stiffening, large deflection and large strain capabilities. The FEA of the assembly is carried out to determine deformation due to inertia effects like gravity, velocity, acceleration, etc., resulting in increase or decrease in the critical assembly feature. A motor assembly consisting of an x-base, a motor, a shaft, a motor base and a crank are investigated using the proposed tolerance synthesis approach discussed previously (Figure 3). The four features of x-base flatness, motor base flatness, motor shaft size, and the motor shaft perpendicularity affect the clearance measurement and they are treated as controllable factors. The dimensioning and tolerancing schemes and tolerance levels are summarized in Table 1 and Table 2; shows the costs for each component tolerance at various levels [1]. The details of full factorial experiment design and response data are obtained from reference paper [1].The output response in this example is the total cost, consisting of manufacturing cost and quality loss as expressed in Equation(1)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002897_irsec.2017.8477423-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002897_irsec.2017.8477423-Figure3-1.png", + "caption": "Figure 3. Representation of SRG", + "texts": [ + " The magnetic flux created by the ampere-turns Ni oscillates between two extreme values corresponds [4]: - An opposing position in which the magnetic circuit has a maximum reluctance as illustrated in Fig.2a. - A conjunction position in which the magnetic circuit has a minimum reluctance as explained in Fig. 2b. The SRM can operate either as: Motor: By proper positioning of current pulses during increasing inductance profile or as, Generator: By proper positioning of current pulses during decreasing inductance profile. The Switched Reluctance Generator (SRG) has three phases: excitation phase, intermediate freewheeling phase, and generation phase as shown in Fig. 3. Excitation phase: when [K 1, K 2] are off, the winding of the stator is excited by an external circuit, the electromechanical energy supplied by the external circuit is converted into magnetic energy. Generation phase: when [K 1, K 2] are on and [D 1, D 2] are off, the magnetic and mechanical energies are converted into electrical energy. The SRG drive consists of the machine itself, the converter and also the controller as depicted in Fig. 4. machine The equations for the SRM phase are written as: , dt id iRV ,, iiLi dt di i LL i dt di LiRV ,, 0 i diiiLiW T The sign before V in (1) is determined by the operating mode of the system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000908_0020-7403(65)90021-4-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000908_0020-7403(65)90021-4-Figure1-1.png", + "caption": "FIG. 1. Examples of inflexional and non-inflexional buckling of prismatic members.", + "texts": [ + " ) The basic equat ions for n o n - i n f l e x i o n a l memb er s can be ob ta ined f rom the relat ions (i) to (v) on se t t ing (for exp lana t ion see. Appendix) q~ = i q \" and a = - ie, where i = ~:- 1. q \" is defined as the shape funct ion of non-inflcxional tensile members in the interval . 0 ~< ' t\" ~ Arsinh [s~-nt~-~/2) and e as the cor responding dis tor t ion pa ramete r . ~\" may be posit ive or nega t ive and when the m e m b e r is s t ra ight , in the limit, e -* 0. In th is case the m e m b e r is par t of the imn-inflexional e las t ics as indicated in Fig. 1 (e). The origin of Lt\" is at the point of lea,~t curvatm'e . In the case of symmet r i ca l ly d i s to r ted tensile m ember s the origin of tF' is ini(tway between the ends, so t ha t ~'~ = - ~',. In all cases the origin of the arc coincides with t h a t of the re levan t shape funct ion, which is assumed to increase wi th the posi t ive value of s towards the end (i). In compress ive members (I) has the value of Tr/2 at. the first posi t ive inflexion I t, the value 3rr/2 a t the next , etc. ( -~r /2 , -31r /2 at the opposi te inflexions)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000674_j.proeng.2010.03.006-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000674_j.proeng.2010.03.006-Figure7-1.png", + "caption": "Fig. 7. Debris from failed roller cage. Fig.8. Debris from the failed cage. The flat piece in the centre is a sidewall of the roller bearing cage.", + "texts": [ + " The failure was detected because of a high level of vibration. Upon dismantling, it was observed that there was a segment missing from the cage in which the rollers were located (Fig. 5). At the end of the missing section of the ring, the tabs or fingers retaining the rollers were observed to be bent, their normal orientation being shown in Fig. 6. Fig.5. Broken-out segment of the roller bearing cage. Fig.6. Roller held in place by tabs. One end of the section of the cage where the break occurred showed a smooth fracture face (Fig. 7, right side). The debris collected from the failed bearing consisted of a number of parts of the cage, most being heavily damaged. (Fig. 8). However one of the side pieces of the cage shown in Fig. 8 (centre) was examined further. One end showed clear signs of fatigue failure (Fig. 9). 1 mm Fig.9. Fracture on the end of the side section of the cage that retained the rollers. The fatigue cracking had initiated at a sharp corner of the pocket that was punched out of the brass cage retaining the rollers: cracking had initiated at two points" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001473_aero.2010.5446818-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001473_aero.2010.5446818-Figure1-1.png", + "caption": "Figure 1 \u2013 Disk geometry.", + "texts": [ + " This paper also presents some results obtained from analyzing a set of tip timing data acquired in a rotor spin test program. The results show that a disk crack as opposed to a blade crack can cause changes to the IBS of several blades near the crack. These changes are dependent on the crack length and the rotating speed. The study concludes that, in spin testing conditions, multiple cracks are detectable and crack propagations can be tracked using the blade tip timing data. The geometry of a rotating disk is described in Figure 1. The crack opening displacement (COD) \u03b4 is determined by the hoop stress and the crack length together with material properties and disk geometry. The hoop stress level of a rotating disk is determined by the rotating speed. The COD at the edge of a cracked body can be approximated by an empirical solution using the following formula from page 53 of [4]. \u239f \u23a0 \u239e \u239c \u239d \u239b= b aV E a ' 4\u03c3\u03b4 (1) where \u03c3 is the uniformly distributed stress, a is the crack length and b the width of the cracked body. Constant E\u2032 is the modulus of elasticity E under plane stress condition, and it equals to E/(1-\u03bd2) under plane strain condition, where \u03bd is the Poisson\u2019s ratio", + " Basically, the postulation is that as the cracks in dovetail slots open up the two closest blades are splitting, which causes the associated IBS\u2019s to increase. The longer crack gives the larger amount of splitting and in turn the larger increases in IBS. The possibility of determining the phase of crack growth for the two cracks is examined here. Referring back to the Beta curves in Figure 3, if we could estimate the Beta values from the differential IBS we might be able to get a feeling on how far the cracks had grown by comparing the estimated Beta values with that of an equivalent ring disk in Figure 1. Figure 17 illustrates the Beta plot from the blade tip timing data using the Beta value corresponding to the initial notch of 1\u2033 plus the estimated Beta value associated with the crack growth. The initial notch related Beta value is in the order of 1.1\u00d710-10 based on the calculation by Eq. (5). The Beta value related to crack growth is estimated using the differential IBS and the blade geometric information, and it is in the order of 0.6~0.7\u00d710-10 towards the end of the spin testing. By locating the total Beta value of 1.7 ~ 1.8\u00d710-10 in Figure 3, we can find that the crack growth had past the inflection point but was still in the steady growth phase with an equivalent a/b ratio of around 0.4, which could mean that the extent of crack induced damage to this test disk is equivalent to the damage caused by a radial edge crack of a/b=0.4 to the ring disk shown in Figure 1. It is important to point out that this result must be used in conjunction with the finite element analysis (FEA). This kind of crack growth (or damage tolerance) testing is normally accompanied by a comprehensive FEA of the test article. However, the FEA for this spin test is not made available for this paper. One must be aware that there are some differences in spin testing and real engine conditions for disk crack detection. One main difference would be the wobble motion of the test article in spin testing which is not exhibited in the real engine condition" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000215_cefc.2010.5481322-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000215_cefc.2010.5481322-Figure1-1.png", + "caption": "Fig. 1. Axial-type magnetic gear in this study", + "texts": [ + " Especially, a magnetic gear operating with axial magnetic flux is thought to be practical due to its simple structure. Then we proposed a new axial-type magnetic gear and confirmed its static transmission torque using the 3-D FEM [1]. In this paper, the effect of eddy current on transmission torque and eddy current loss in an axial-type magnetic gear is discussed employing T- method in 3-D finite element analysis. The validity of this calculation is verified by carrying out the measurement on a prototype. Fig. 1 shows an axial-type magnetic gear used in this study, which mainly consists of a high-speed rotor, a low-speed rotor, and stationary pole pieces. A high-speed rotor generates magnetic harmonics in the air gap between stationary pole pieces and the low-speed rotor. While a high-speed rotor rotates, a low-speed rotor rotates in accordance with the gear ratio. In this analysis, the T- method is employed: 0graddiv em TT (1) gradrot1rot eme TTT t (2) where Tm and Te are current vector potentials of equivalent magnetizing current density and eddy current density, respectively, is the permeability, and is electrical conductivity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002033_elma.2017.7955470-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002033_elma.2017.7955470-Figure4-1.png", + "caption": "Fig. 4. Field distribution of the flux density in the two investigated constructions of the coaxial magnetic gear", + "texts": [ + " The inner rotor and the outer rotor of the first construction of the gear rotate with 150 rpm and 27.3 rpm, respectively. The inner rotor of the second construction rotates with 150 rpm, and the outer rotor rotates with 50 rpm. The steel segments stay stationary for both magnetic gear\u2019s constructions. The results of the analysis of the harmonic components of the magnetic field of the two constructions of the magnetic gear are obtained using Fast Fourier Transform (FFT) method via MATLAB software. Results for the flux density of both constructions are given in the Fig. 4. The upper limit of the flux density is confined to 2 T. It can be seen from Fig. 4 that in some regions the flux density has higher values at the first construction of the magnetic gear than the second one. The dynamic torque transmission characteristics for both constructions are depicted in Fig. 5. As shown above the dynamic torque transmission characteristic of the first construction has lesser ripples of the torque than the dynamic torque transmission characteristic of the second construction of the coaxial magnetic gear. The FFT results for the dynamic torque of the inner rotors for both constructions are depicted in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000239_sii.2010.5708334-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000239_sii.2010.5708334-Figure2-1.png", + "caption": "Fig. 2. The concept of a quad rotor tail-sitter", + "texts": [ + " It can hover like a quad rotor helicopter, and can fly long distance like a fixed-wing airplane. In order to verify this concept, a simulator of a quad rotor tail-sitter UAV is developed. Simulated results show the advantage in energy efficiency of the proposed UAV over a conventional quad rotor helicopter type UAV. This paper 978-1-4244-9315-9/10/$26.00 \u00a92010 IEEE - 254 - SI International 2010 also describe the design and development of a quad rotor tail-sitter UAV. The concept of a quad rotor tail-sitter UAV is shown Fig. 2. A quad rotor tail-sitter has a four flight forms, hovering flight, low speed flight, level flight and transitional flight. In case of approaching a target or ground, the UAV hover and make a low speed flight like a quad rotor helicopter. In case of flying long distance, it tilt the fuselage with the propulsion unit forward and cruises like a fixed-wing aircraft. Hence, the fixed-wing generates the lift and improves energy consumption in level flight. Additionally, transitional flight is the transition from the hovering flight to level flight or from the level flight to hovering flight" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002586_apec.2018.8341172-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002586_apec.2018.8341172-Figure9-1.png", + "caption": "Fig. 9. Photographs of 12/8 SSHE-SRM prototype. (b) Stator. (b) Rotor.", + "texts": [ + " Finally, Table I summarizes the predicted performances of the SSHE-SRM based on the NBC and AHBC at different conditions and control modes. As shown in Table I, the characteristic of the SSHE-SRM based on NBC is better than that based on AHBC, such as higher average torque, lower torque ripple and higher torque density under the same conditions. In order to validate above theoretical analysis, predictions and simulations of dynamic performance, a 12/8 SSHE-SRM prototype is manufactured and the SSHE-SRM drive with NBC and AHBC is developed for experimentally test. Fig. 9 shows detailed photographs of the 12/8 SSHE-SRM prototype. The static flux linkage and average torque characteristics of the SSHE-SRM are tested and validated. In order to measure the static magnetic characteristics, a simple indirect method, namely locked-rotor test, is performed. The measured results of the SSHE-SRM with FEA results are compared and shown in Fig. 10. It can be seen that the FEA result agrees well with the measured values, which validates aforementioned analysis. The errors may be attributable to the neglect of end effect in 2D-FEA, nonuniformity of the air gap and manufacturing tolerances" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002850_icra.2018.8461156-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002850_icra.2018.8461156-Figure3-1.png", + "caption": "Fig. 3. Spatial decomposition of instantaneous optic flow for straightline flight along a corridor laden with two small-field obstacles (poles).", + "texts": [ + " \u00b0 For motion restricted to a plane, it is well-known that wide-field optic flow patterns are spatially periodic that reside in L2[\u2212\u03c0, \u03c0], the space of square-integrable and piecewise-continuous functions, and can be modeled using the first N harmonics of the Fourier series as [18], Q\u0307WF (\u03b3,x)\u2248a0 2 + N\u2211 n=1 (an cosn\u03b3+bn sinn\u03b3) , an= 1 \u03c0 \u222b \u03c0 \u2212\u03c0 Q\u0307(\u03b3,x) cosn\u03b3 d\u03b3, bn= 1 \u03c0 \u222b \u03c0 \u2212\u03c0 Q\u0307(\u03b3,x) sinn\u03b3 d\u03b3. (2) For vehicle flight close to the centerline of a straight corridor, the tangential optic flow profile resembles a sinewave like pattern, with the peak amplitude proportional to the forward speed and inversely proportional to the perpendicular distance of the vehicle to the wall. Fig. 3 depicts the instantaneous optic flow field induced by simulated vehicle flight in a textured corridor with two small-field obstacles (poles), while Fig. 4 depicts the instantaneous small-field residual signal reconstructed with sensor noise for the case shown in Fig. 3. From these figures, it is clear that the introduction of a small-field obstacle induces a high-frequency spatial perturbation in the nominal optic flow pattern. The objective then is to eliminate the low-frequency (wide-field) content present in optic flow so that only the small-field (SF) signal, which encodes motion cues necessary for small-field obstacle avoidance, remains. Additionally, a thresholding mechanism will need to be implemented to mitigate the influence of high-frequency noise. \u00b0 \u00b0 (deg) O p ti c f lo w r e s id u a l (p ix e l/ s ) -200 -100 0 100 200 0 -0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001045_jsea.2010.312134-Figure11-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001045_jsea.2010.312134-Figure11-1.png", + "caption": "Figure 11. Deformation due to velocity effect.", + "texts": [], + "surrounding_texts": [ + "In this research, the proposed approach provides better formulation of cost-tolerance relationships for empirical data. BP network architecture of configuration 4-6-1 generates a suitable model for cost-tolerance relationship of R2 value 0.9997, there by eliminating errors due to curve fitting in case of regression fitting. And it also generates more robust outcomes of tolerance synthesis. The proposed non conventional optimization technique obtains an optimal solution better than that of simulated annealing [6] and Response surface methodology (RSM) [1].This study proposes a tolerance synthesis based on BP learning, a NSGA II based optimization algorithm and CAD interface, in order to ensure that the proposed values of controllable factors (tolerances) satisfies the assembly constraint, even before the start of manufacturing process. There by reducing scrap and rework cost." + ] + }, + { + "image_filename": "designv11_33_0002636_j.mechmachtheory.2018.05.003-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002636_j.mechmachtheory.2018.05.003-Figure3-1.png", + "caption": "Fig. 3. Apex seal geometry and dimensions.", + "texts": [ + " 2 , as follows R lc = \u221a ( L \u2212 C ) 2 + ( k ) 2 (5) In the following sections, the forces and pressures, as well as the motion of the seal will be investigated. 3. Apex seal kinematics model formulation The apex seal is modelled as a simple spring-mass-damper system in which the seal is positioned inside a groove measured W g in width. A spring is attached to the inner end of the groove and one end of the seal; the other end of the seal pushes against the housing. The cross-sectional area of the seal is almost rectangular with one rounded end, dimensions and geometry of which is depicted in Fig. 3 . The seal width and height are denoted W s and d s , respectively. With the case of lima\u00e7on-to-lima\u00e7on machine, the truncation distance, z , can be calculated as: z = L \u2212 ( r sin 2 \u03b80 + L cos \u03b80 ) (6) where \u03b80 is the constant angle between the rotor axis and a line connecting the rotor centre point to edge point 3, which is shown in Fig. 3 . The angle \u03b80 can be derived from the lima\u00e7on Eq. (3) or (4) as follows: \u03b80 = sin \u22121 ( \u2212 L 4 r + \u221a L 2 ( 4 r ) 2 + W g 4 r ) = sin \u22121 ( \u2212 1 4 b + \u221a 1 ( 4 b ) 2 + W g 4 Lb ) (7) With the case of the circolima\u00e7on and the lima\u00e7on-to-circular machines, where the rotor lobes are manufactured of circular arcs, the truncation, z , can be calculated as: z = L \u2212 C r \u2212 \u221a R lc 2 \u2212 ( k + W g 2 )2 (8) Let k s be the seal spring stiffness, and k w be the equivalent stiffness of the wall of the machine housing ( k w k s ); and let I be the instantaneous centre as shown in Fig", + " (15) and \u03d5\u0308 is the angular acceleration of the seal, m s is the seal\u2019s mass and I CG is its mass moment of inertia about the centre of gravity, F x is the sum of forces acting on the seal in the x -direction F y is the sum of forces acting on the seal in the y -direction M CG is the sum of moments acting on the seal about its centre of gravity. At every rotor angle, \u03b8 , the three non-linear simultaneous differential equations given in (80) are solved numerically to find the seal displacements, being x, y and \u03d5. 4.4. Possible seal-groove interactions As pointed out by Phung and Sultan [8] , the interactions between the seal and seal groove can be categorised into nine different cases. To differentiate between the cases, seal-groove contact points have been numbered as shown in Fig. 3 . Points n, p and q respectively represent the negative, neutral, and positive sides of the seal groove inner end; while points 1, 2, and 3 represent the negative, neutral, and positive sides of the seal groove outer end, respectively. The negative side of the groove is on the negative side of the negative side of Y 1 axis, while positive side of the groove is on the positive side of Y 1 axis. The two imaginary points 2 and p , point 2 sits between points 1 and 3 while point p sits between points n and q , are also shown in Fig. 3 above. The seal cases can now be categorised by a letter, which represents the position of back-contact point, followed by a number, which represents the position of front contact point. i.e.: Case q1 is when the top left-hand corner of the seal contacts with the positive side of the groove, denoted q , and the negative side of the seal contacts with the negative inner side of the groove modified apex, denoted 1, as shown in Fig. 7 a. Of note is that back-contact happens when the back corner of the seal contacts with the groove surface; front-contact is when either the negative or positive side of the seal comes into contact with the inner side of the modified apex" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002076_optim.2017.7975006-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002076_optim.2017.7975006-Figure6-1.png", + "caption": "Fig. 6. Position of points for magnetic field evaluation", + "texts": [ + " 5 show the increase of the deviation of the current Iu(t) from the harmonic variation in time. The increase of the amplitude of particular harmonics in comparison with healthy state is evident. The rms values of the three phase currents are Iu = 7.215 A, Iv = 8.905 A and Iw = 6.357 A. The value 40.47 A of the current in the shortcircuited elementary coil of the phase U in this faulty state, much higher than the motor rated current, shows that FA4 is a severe short-circuit fault. Points in the air region outside the motor, M[116, 0, 0] and N[-116, 0, 0], Fig. 6, which are placed in the transversal symmetry plane z = 0 of the motor and P[116, 0, 90] and Q[-116, 0, 90], Fig.6, in the plane z = 90 mm crossing the stator winding volume outside the magnetic core are considered for the evaluation of the radial component Bx and of the axial component Bz of the magnetic flux density. The comparison of the time dependences and spectra of the radial component Bx[M](t) of the magnetic flux density in the point M in Fig. 7 and Fig. 8, highlights changes from the healthy state HE of motor operation to the faulty state. A relatively reduced decrease of the amplitudes of the 50 Hz harmonic of this magnetic field component, from 6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000091_3.4135-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000091_3.4135-Figure1-1.png", + "caption": "Fig. 1. Notation.", + "texts": [ + " For end conditions other than the classical simple support, procedure similar to that in Ref. 9 can be used. Basic Equations In Ref. 9, three simultaneous linear differential equations were derived for the displacements-u*, v*} and w* of a typical point of the middle surface of a thin-walled circular cylindrical shell. In this derivation, the method devised by Sanders8 was used, and the effect of initial membrane stresses was included. The three equations are given below (for sign conventions see Fig. 1): i <-\u00bb \\-*- */* 2 : + L 2 \" ~8~ (1 ~ r fl + v _ 3Ko _ , I 2 ~ T\" ( ' '+ aox*o

is mathematically defined by,( xc yc ) = ( (xt + xR)\u2212R sin(\u03b8t + \u03c9t\u2206t) (yt + yR) +R cos(\u03b8t + \u03c9t\u2206t) ) (11) where \u03c9t\u2206t is the rotation increment in the last period of time, and (xt, yt) is the actual robot position that arise from the inertial measurement system", + " the ICR model [14] we establish, ( xt+1 yt+1 ) = ( xt yt ) + ( xc \u2212 xR \u2212R sin(\u03b8t + \u03c9t\u2206t) yc + yR +R cos(\u03b8t + \u03c9t\u2206t) ) (12) The property of skid-steering gives the robot the ability to change its turning axis, in accordance to the wheels lateral slippages. The authors took advantage of this effects, by calculating a point coordinate called (xR, yR). During preliminary motion tests, some interesting variations in the motion patterns were observed in the mobile robotic platform that were experimentally disclosed when applying different wheels speed. A kinematic restriction of this effect establishes that the rover turning Z-axis only moves within a squared area bounded by the wheels\u2019 contact point (as depicted in fig.2). By modelling the observed effects when parametrizing different angular speeds on the front and the rear wheels, we changed its equilibrium point, so the equations to calculate xr and yR were formulated to model such inertial effects. Where yR is the turning Z-axis displacement along its longitudinal axis; L is the distance between the rear and front side wheels. vmax is the maximal allowed robot velocity reached up to a contact point. Similarly, xR is the displacement or shift of the robot reference through its transverse axis", + " The inertial coordinates of the front side accelerometer is given by xAI = \u222b t \u222b t (axA cos\u03b8 \u2212 ayA sin\u03b8)dt2 (32) yAI = \u222b t \u222b t (axA sin\u03b8 + ayA cos\u03b8)dt2 (33) These are the inertial coordinates of the rear side accelerometer, xBI = \u222b t \u222b t (axB cos\u03b8 \u2212 ayB sin\u03b8)dt2 (34) yBI = \u222b t \u222b t (axB sin\u03b8 + ayB cos\u03b8)dt2 (35) Where xI and yI are the location of each accelerometer in an inertial frame. Two kinds of experiments were carried out, simulation testing, and laboratory experiments deploying \u201cRoverto\u201da rover-like mobile robot shown in figure 2. This robotic platform is instrumented with independent control speed in each wheel, wireless communication, an in-house made inertial unit compounded by a couple of accelerometers, a vision system, and computer on-board. The mobile robot used in this research work possess a differential drive based mechanical configuration, using four wheels for stability and very good controllability capabilities. A solution to inverse and direct kinematics for trajectory control of a four-wheeled mobile robot was proposed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003921_b978-0-12-803581-8.11722-9-Figure16-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003921_b978-0-12-803581-8.11722-9-Figure16-1.png", + "caption": "Fig. 16 Schematic of a mechanical part of a research station used to examine hydraulic clutches and brakes with ER fluid: 1 \u2013 electrical engine, 2 \u2013 examined clutch, 3 \u2013 connecting clutch, 4 \u2013 torque meter, 5 \u2013 controlled brake, 6 \u2013 frame.", + "texts": [ + " Omitting the movement resistance of the research station allows to use one torque meter which, as far as possible in the construction conditions of the research stations, should be placed behind the examined clutch prototype. Main physical values measured during experimental research of clutches and brakes with ER fluid are: angular velocity of input shaft o1 and output shaft o2, transmitted torque M, working fluid temperature T and voltage U supplied to the clutch\u2019s electrodes. According to needs, other values are also measured, e.g., pressure decrease Dp in channels of the hydrodynamic clutch, intensity of the load current I or relative air humidity w. Fig. 16 depicts a schematic of a mechanical part of a typical research station used to examine hydraulic clutches and brakes with ER fluid. In research stations used to examine hydraulic clutches and brakes with ER fluid, instead of a torque meter, a force sensor can be used. In such case the propeller part of the brake is connected to rotatably mounted lever whose tip touches the sensor. The torque M is obtained by multiplying the sensor\u2019s indications by the length of the lever. Such a way of measuring the torque M is also used during research conducted on clutches with ER fluid for o2 \u00bc 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002795_gt2018-77151-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002795_gt2018-77151-Figure5-1.png", + "caption": "Fig. 5 Cross-section of bearing test rig", + "texts": [ + " [9] note the importance of a proper discretization model to produce accurate temperatures in a turbulence flow analysis, in particular the number of nodes across the film, at least 101. In this study, the film flow domain in a pad has 38 nodes in the circumferential direction, 101 nodes across the film, and 22 nodes in the axial direction. In addition, 9 nodes are used for a pad thickness. An iterative method searches for the journal position due to an applied load and convergence to steady film and pad temperature fields. The convergence criteria must satisfy a 0.1% difference in applied load and a difference in temperature within 0.1K. Fig. 5 shows a cross-section of the test rig for measurement of the static load performance of lubricated bearings, and Fig. 6 shows a photograph of the test bearing, rotor and support bearings. An electrical motor and gear box drive a rotor to a speed as high as 24k rpm. The rotor, 101.6 mm in diameter, is supported on a pair of tilting-pad journal bearing; and a four pad test bearing, flexure pivot type, is located at the rotor midspan. Table 2 lists the test bearing geometry and the lubricant material properties" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003886_s40684-019-00170-w-Figure10-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003886_s40684-019-00170-w-Figure10-1.png", + "caption": "Fig. 10 Simulation scheme using 20-Sim for model validation", + "texts": [ + " Figure\u00a07 illustrates dynamic model of parallel gear involving damping and meshing coefficients. Gearbox subsystem is composed of: \u2022 Planetary gear connected to the output rotor shaft; \u2022 Parallel gear connected to the output planetary gear shaft; Figures\u00a08 and 9 show BG detailed model of the studied transmission. This model is characterized by the addition of passive elements R where Sanchez and Al. [18] considered only the energy storage elements C. In order to validate the above model, the scheme of Fig.\u00a010 is carried out under 20-Sim. The verification scenario consists of subjecting the multiplier to an upstream torque that models the torque from the rotor, and a downstream flow source modeling the rotational speed imposed by the generator. The results of the simulation shown in Fig.\u00a011 illustrate input and output torque and flow. It reveals a ratio of 60.5 with a yield of 98% calculated using Eq.\u00a0(1). With: \u2022 C1: input torque; \u2022 \u03c91: speed of rotation at the entrance of the train; \u2022 C2: output torque; \u2022 \u03c92: angular output speed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000225_004051756703700306-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000225_004051756703700306-Figure1-1.png", + "caption": "Fig. 1. Apparatus for measuring bending and extension moduli : G―galvanometer; A-galvanometer mount ; Ri, IZ4 R,-sliding rods ; Rf-threaded rod ; B--galvanometer positicming screw ; and E--clamp securing screw.", + "texts": [ + " This method was developed by Pierce [9] and used hy Carlene [11 ] for viscose rayon filaments. This method is unsuitable for fihers which are not uniform enough in cross section and are difficult to handle. , In an attempt to clarify differences between static bending and stretching, an instrument was built for the determination of static bending, while an Instron tensile tester was used in determining extension moduli. Measurements were conducted on mohair and kemp fibers. Experimental Apparatus The appar atus (Fig. 1 ) consisted of a galvanometer G, with a segment of a razor hlade mounted on the end of its pointer to provide a knife edge. Current passing through the coil deflected- the needlethe torque produced heing proportional to the current [7] ] in the coil. This torque was opposed hy the restoring torque in the galvanometer restoring springs. By hanging different weights over the knife edge, the current required in each instance to restore the needle to its zero position (i.e., restoring torque of the springs equal to zero) was determined, and from that a calihration curve was drawn" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001426_icca.2010.5524116-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001426_icca.2010.5524116-Figure2-1.png", + "caption": "Fig. 2. The schematics of the TORA system", + "texts": [ + "1, then the input/output constraints in (2) can be satisfied for the whole LPV systems, with the ranges of magnitude and rate of varying parameters given in (3a) and (3b), if the following matrix inequalities are also imposed on the matrix variables { }j j,Y Q and 0,\u03b1 > j 0 j,x0 j 0 ( ) : 0 \u03b1\u03a4 \u00a7 \u00b7 = \u2265\u00a8 \u00b8 \u00a9 \u00b9 Q x Q x $ (19) ( )u j iu j 2j,u j j iu,maxu iu j ( , ) : 0 \u03b1 \u03a4\u00a7 \u00b7 \u00a8 \u00b8 = \u2265\u00a8 \u00b8 \u00a8 \u00b8 \u00a9 \u00b9 Q Y Q Y u Y A A $ (20) ( ) ( ) j 2 j,z j j jz,maxz j j * ( , ) : 0, jz \u03b1 \u00a7 \u00b7 \u00a8 \u00b8 = \u2265\u00a8 \u00b8 +\u00a8 \u00b8 \u00a9 \u00b9 Q Q Y z CQ DYA $ (21) for 1, , ,j k= \" 1, , , u i m= \" 1,2, , z j p= \" , where u 1 m iu \u00d7\u2208A * with its iu-th element equaling to one and others zeros, 1 pz jz \u00d7\u2208A * with its jz-th element equaling to one and others zeros. Sketch of Proof. As an extension from Lemma 2.2, the proof can be obtained based on the derivation in Theorem 3.1. IV. ILLUSTRATIVE EXAMPLE The proposed controller design algorithm is demonstrated by a benchmark continuous-time nonlinear system, namely the translational oscillator with rotational actuator (TORA) system which has been studied previously, e.g., in [18]\u2013[20]. The TORA system is depicted in Fig. 2. The oscillator consists of a cart of mass M connected to a fixed wall by a linear spring of stiffness k . The cart is constrained to have one-dimensional travel. The rotating proof mass actuator attached to the cart has mass m and moment of inertia I about its center of mass with distance e from the point about which the proof mass rotates. The control torque N is applied to the proof mass. The object of this problem is to use the control torque N to attenuation disturbances acted to the base translational mass" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001591_acc.2011.5990742-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001591_acc.2011.5990742-Figure2-1.png", + "caption": "Fig. 2. Level sets of the function V\u2217(x,x). The level set V\u2217(x, x) = 1 is identical to the the boundary \u2202G = {x : V (x) = 1} = {x : v(x) = 1}", + "texts": [ + " The following deformed version of (10) is used to overcome the tracking problem V\u2217(x,x) = inf { \u03b3 > 0 : x + 1 \u03b3 (x \u2212 x) \u2208 G } (11) with x \u2208 intG where intG denotes the interior of G. Note that (11) is convex and for any x \u2208 G and any fixed x \u2208 intG V\u2217(x,x) = 0, (12) V\u2217(x,x) < 1 if x \u2208 G, (13) V\u2217(x,x) = 1 if x \u2208 \u2202G (14) hold. Therefore, the function V\u2217(x,x) seems to be a suitable Lyapunov function for the tracking problem. Next, we can state the tracking controller u = kT(x, r) = k2(x\u0302)V\u2217(x,x) + (1\u2212V\u2217(x,x))u. (15) Therein, the vector x\u0302 = x + 1 V\u2217(x,x) (x \u2212 x) (16) lies on the boundary of G, i.e., x\u0302 \u2208 \u2202G holds. Fig. 2 illustrates the level sets of the function V\u2217(x,x) together with x and x\u0302. Note that the level set {x : V\u2217(x,x) = 1} is identical to the level set {x : V (x) = 1} = {x : v(x) = 1} and therefore identical with \u2202G. The main result ensuring the stability of the tracking control is stated in Theorem 1: Given a domain of attraction G of the system x\u0307 = Ax + B\u03c3 (K1x + T\u03c3(u)) under the controller u = k2(x) that satisfies k2(cx) = ck2(x) as well as the conditions \u2202v(x) \u2202x (Ax + BM(v,h1(x),K1x + Th2(x), K1x + Tk2(x))) < 0 \u2200x \u2208 G,v \u2208 V and G \u2286 L(h1(x)) \u2229 L(h2(x)) for all v \u2208 V " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003914_icems.2019.8922162-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003914_icems.2019.8922162-Figure2-1.png", + "caption": "Fig. 2. The finite element models of 9s6p and 12s8p PMSMs In this paper, 9s6p and 12s8p PMSMs are taken as examples, and the distribution characteristics of the radial electromagnetic force of non-coaxial eccentricity and elliptical eccentricity with eccentricity ratio of 20% are compared by finite element software. Fig. 2 shows the finite element models of 9s6p and 12s8p motors.", + "texts": [], + "surrounding_texts": [ + "FREQUENCY ORDER (k=0,1,2\u2026) \u03c9p2 kZp \u00b1\u03c52 16 \u00b1= k\u03bd \u03c9p2 kZp \u00b1+ )( 21 \u03c5\u03c5 1616,21 \u2212=+=\u2260 kk \u03bd\u03bd\u03bd\u03bd \u03c9p2 kZp \u00b1\u2212 )( 21 \u03c5\u03c5 16,16 21 kk =\u00b1= \u03bd\u03bd ( ) \u03c9\u03bc p\u00b11 kZp \u00b1\u00b1 )( \u03c5\u03bc 16,12 +=+= kk \u03bd\u03bc ( ) \u03c9\u03bc p\u00b11 kZp \u00b1)( \u03c5\u03bc 16,12 \u2212=+= kk \u03bd\u03bc ( ) \u03c9\u03bc\u03bc p21 \u00b1 kZp \u00b1\u00b1 )( 21 \u03bc\u03bc 12 += k\u03bc" + ] + }, + { + "image_filename": "designv11_33_0000523_978-94-007-0020-8_28-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000523_978-94-007-0020-8_28-Figure1-1.png", + "caption": "Fig. 1 A beam segment and the analogous motion planes", + "texts": [ + " A locking free Timoshenko beam finite element formulation [12] is carried out, in which the exact solution of differential equations of motion is used to derive the shape functions. The rotor is treated as a free-free body in space upon which combination of static and dynamic forces (e.g., gravity, bearing, imbalances and gyroscopic moments) can act. The rotor is modeled by circular Timoshenko beam finite elements to account for shear deformation and rotary inertia [15]. The equations of motion of a rotating Timoshenko beam element, in two perpendicular planes of motion XZ and YZ (Fig. 1) can be written as follows: @ @Z AGS @x @Z \u02db C px.Z; t/ D mz @2x @t2 @ @Z AGS @y @Z \u02c7 C py.Z; t/ D mz @2y @t2 @ @Z EI @\u02db @Z C AGS @x @Z \u02db D JT @2\u02db @t2 C !rJz @\u030c @t @ @Z EI @\u030c @Z C AGS @y @Z \u02c7 D JT @2\u02c7 @t2 !rJz @\u02db @t (1) Here the following relations hold in the XZ plane. \u02db D @x @Z C sxI sx D Vx GsA I MY D EI @\u02db @Z (2) The corresponding relations in the YZ plane are employed. Further px and py are the distributed external forces with respect to XZ and YZ planes respectively, Gs and are the shear modulus and the shear form factor " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000623_tac.1962.1105481-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000623_tac.1962.1105481-Figure2-1.png", + "caption": "Fig. 2-Contours of equal 1,fSE p ~ ( e , b) . The point (al, b ~ ) gives minimum SlSE with input r l ( t ) and (a2, bn) gives minimum MSE with input r2(fj.", + "texts": [ + " The error cl(tj formed b,- taking the difference between Model I1 and the desired (Model I) outputs is squared to obtain an instantaneous error measure. This measure contains a component at the perturbation frequency X. The amplitude and phase of this component give the amplitude and sign of a signal which can be recovered b!, multiplication and integration (Fig. 1). This is fed back negativel!. to a. to reduce the short time average of Two parameters ]nay be controlled by perturbing them at different frequencies (or a t least in theory by the sine and cosine of the same frequenq-) and using two adaptive loops. Fig. 2 shows contours of equal mean square error (IISE), il?, as a function of the two parameters a and b of Model 11. Suppose for a particular input spectrum the minimum MSE occurs a t (al, b1). If originally the parameter values \\\\:ere (ao, bo), then when the adaptive loops are operating the parameters a and b change and ultimately equal (al , blj . I t will be shown in Section 111 that if the gains of the tn-o adaptive loops are suitably chosen, the adjustment of a and b will be a t all times perpendicular to the tangent of the kiSE contours, that is, the parameters are adjusted along the path of steepest descent. If the input spectrum changes, the 1ISE contours shift and may have a new minimum a t (a?, b?) (Fig. 2). The adaptive loops now force the parameters ( u , b j to the new optimum values (a?, b?). To sum up, the upper part oi the adaptive scheme chooses the \u201cbest\u201d values of the parameters a and b of AIodel I1 corresponding to various input signals. Referring now to Fig. 1, the purpose of Adaptive loop 2 is two-fold: 1) It adjusts the controller parameter to compensate for the unpredictable variations of the process parameter; and 2) i t forces the controller parameter to track the 3lodel I1 parameter as it varies under the influence of Aklaptive loop 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003617_01691864.2019.1657947-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003617_01691864.2019.1657947-Figure3-1.png", + "caption": "Figure 3. Coordinate systems configuration.", + "texts": [ + " In order to simplify the model, we only focus on the bending of a single direction, and certain assumptions are made, as follows: A1: The friction force between the cable and backbone equals the maximum static friction force when the cable slides along the guiding rings. A2: Gravitational force is ignored because all the backbones are very small (the entire weight is less the 10 g). A3: The extra length of the cable added by stretching is ignored because the tension force is small. A coordinate system is defined at each joint, as shown in Figure 3. The origin is set at the joint center the Z axis directed along the axis of the backbone. The endoscope ismainly used for viewing in surgical procedures, there is no significant load requirement and no external force applied. So its mechanical balance can be built starting from its proximal end. Our model makes use of numerous quantities, which are defined in Table 1. Figure 4 shows the force and moment applied on the Link L1n. The proximal link (which is also the base) is L0, and the base coordinate O0 is located at L0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001301_1.52410-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001301_1.52410-Figure4-1.png", + "caption": "Fig. 4 Record disk orbit.", + "texts": [ + "x 0 _y 0 0 (22) Under this condition, the trajectory of the deputy satellite in the x-y plane is an ellipse that has its center at xc; yc 0; y 0 2 _x 0 ! (23) and minor and major radius of a\u0302 x2 0 _x2 0 =!2 p (24) b\u0302 2 x2 0 _x2 0 =!2 p (25) in the x-axis, and y-axis, respectively. In addition, if the initial position and velocity in the z-axis are selected as z 0 x 0 tan (26) _z 0 _x 0 tan (27) and is selected as =3, then the motion in the z-axis has the same phase as that in the x-axis, and the trajectory in the Hill\u2019s coordinate becomes a circle of radius 2 x2 0 _x2 0 =!2 p (28) This trajectory, shown in Fig. 4, is well known as a cart orbit, or a record disk orbit. Figure 5 shows the geometrical relation between the inclination i and eccentricity e in the record disk orbit, which is drawn from a side view. If the relative inclination 2i and eccentricity e are small enough to treat the geometrical relation in term of linearity, the parameter can be represented as arctan i e (29) whichmust be =3 to obtain a record disk orbit, therefore, the relative inclination i must be related to the eccentricity e, by i 3 p e (30) D ow nl oa de d by U N IV E R SI T Y O F O K L A H O M A o n Ja nu ar y 24 , 2 01 5 | h ttp :// ar c" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001487_j.sna.2010.07.024-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001487_j.sna.2010.07.024-Figure2-1.png", + "caption": "Fig. 2. Experimental set-up for measuring the deflection of PVCCF in thermal cycling.", + "texts": [ + " DSC measurements Differential scanning calorimetry (DSC) measurements were carried out for PVC and CFRP in order to establish their thermal characteristics. PVC and CFRP were put through the preheat treatment before their DSC measurement. 3.6. Theoretical analysis of deflection and force generated Employing the experimental data obtained by the measurements described so far, a theoretical analysis based on the beam theory for PVCCF deformation under a load against an environmental temperature change was carried out. 4. Results and discussion 4.1. Deflection of PVCCF under thermal cycling Fig. 2 shows the experimental set-up for measuring the temperature dependence of PVCCF deflection. The PVCCF was suspended in a water bath, the temperature of which was repeatedly raised and cooled between 300 and 340 K in order to control the PVCCF temperature. The horizontal displacement of the PVCCF 6 cm away from the clamping point was measured by a laser displacement sensor, and the displacement was converted into a curvature. Fig. 3 shows the temperature dependence of the PVCCF curvature by the two thermal cycles" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002159_gt2017-63208-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002159_gt2017-63208-Figure3-1.png", + "caption": "Fig. 3 Schematics of fluid dynamic loss", + "texts": [ + " The gearbox loss-reduction methods can be evaluated effectively by using the CFD of the oil and air inside a gearbox [7\u201310]. Therefore, we perform CFD modeling based on the loss mechanisms. Then, the CFD simulation results and the experimental results are compared for each loss component. 1 Copyright \u00a9 2017 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 08/19/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Figure 2 shows the classification of loss in a gearbox. Figure 3 shows schematics of the loss. Airflow is governed by the momentum conservation law of continuum because the air molecules are always dispersed. On the other hand, oil follows two governing-laws depending on the dispersing state in the air. One is the momentum conservation law of point mass treated as \u201coil particle\u201d which does not conserve the momentum as continuum. The other is the momentum conservation law of continuum treated as \u201coil flow\u201d (e.g., [11]). Accordingly, the present classification has the following characteristics that are different from the characteristics of other classification methods [3\u20136]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001434_046009-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001434_046009-Figure4-1.png", + "caption": "Figure 4. (a) A microscopic view of a short piece of the filament shows the arrangement of the model flagellins in the protein network. (b) One model flagellin viewed from the side (top) and from the front (bottom). Rigid black arms emanate from a central black spot to neutral binding sites 1, 2, 7, and 8, which exist in both states of the flagellin. The green sites 3, 4 and the red sites 5, 6 are only activated in the L- and R-state, respectively. The blue transparent surface only symbolizes the borders of the rigid body. The central black spot is also the origin of a Cartesian coordinate system attached to the rigid body. The origin lies in the plane of binding sites 1\u20136 and its horizontal position is in the middle between sites 1, 2 and 3\u20136. The y coordinates A and B define the respective axial positions of the binding sites 3, 5 and thereby the internal twist of the pure L- or R-form of the filament. The distance between the two binding sites 3, 4 and 5, 6 are given by QL and QR , respectively. The distance between sites 1, 2 is E and their y coordinates are E/2 and \u2212E/2. The inner binding sites 7, 8 are characterized by the axial displacement or y coordinate D of their center and their distance F. For further use, we define the difference C = QL \u2212 QR , which is the difference in length of the flagellin in the L- and R-state. Finally, b and t are the width and depth of the model molecule.", + "texts": [ + " In our rigid-body model for the flagellin, we follow work by Namba and Vonderviszt [2] who set up a characteristic binding scheme. Via distinct binding sites each molecule forms elastic bonds with its nearest neighbors. The L- and R-states of flagellin differ by the location of two of these binding sites. With their model, Namba and Vondervistz could reproduce curvature and torsion of the 12 possible polymorphic configurations of the flagellum [17]. In our model we have extended the binding scheme. As we will demonstrate in section 4.1, this proved to be crucial for stabilizing the normal configuration of the filament. Figure 4(a) shows a short piece of the filament and in figure 4(b) one of its flagellin molecules is viewed from the side (top) and from the front (bottom). Rigid black arms emanate from a central black spot to the binding sites. Whereas the black backbone symbolizes the rigidity of the molecule, the transparent blue form should only give an impression of its silhouette but has no physical meaning for the model. The gray binding sites with labels 1 and 2 are present in both states of flagellin, while the red binding sites (labels 5, 6) are only activated in the R-state and the green sites (labels 3, 4) in the L-state", + " Our early investigations showed that the elastic network following from this binding scheme is unstable and collapses to a non-tubular form, like a balloon which loses its form after deflation [42]. The model is unstable against disturbances along the radial direction. For this reason we also include the interactions between the D0 domains of the flagellin proteins in the inner core of the filament. We already mentioned that they are also crucial for realizing the helical configurations of the filament [35]. In figure 4(b), the corresponding inner binding sites are labeled by 7 and 8. They form bonds with neighboring molecules along the 1-start helix as summarized in table 1. With this extension the elastic network relaxes into its stable ground state at a given spin configuration {\u03c3i}. The figure caption of figure 4 contains a detailed description of the model parameters. We stress here the distance E between the neutral binding sites 1, 2 and the distances QL, and QR of the state-specific binding sites 3, 4 and 5, 6. Furthermore, A and B are the axial positions or y coordinates of binding sites 3 and 5. Figure 5 illustrates how pure L-type (blue) and R-type (red) rigid bodies pack together and form bonds with their neighbors along the 5-start and 6-start directions. The bonds, in the D1 domain, always form between a neutral and an active binding site", + " The distance F of these binding sites does not depend on the flagellin state, so we choose an average value F = (2E + QR + QL)/(2 \u00b7 11) = 5.227/11 nm. Furthermore, we have the spring constants kO and kI of the outer and inner bonds. Since the absolute value of the elastic energy is not relevant in this section, we give the energy in terms of kO times the square of a characteristic length and choose kI relative to kO . Finally, we need the positions A and B of the state-dependent binding sites (see figure 4(b)). We calibrate their values such that the twist of the relaxed rigid-body network in the pure L- and R-states corresponds, respectively, to the measured values of pure Land R-state flagellar filaments, namely \u03c4L = \u22125.62 \u03bcm\u22121 and \u03c4R = 13.45 \u03bcm\u22121 [37]. This calibration is repeated whenever a parameter is changed. In figure 6, we plot curvature versus torsion for the 12 polymorphic states. The different colors show results when the strength of the inner bonds varies from strong (gray: kI = 2kO) to weak (red: kI = kO/20) relative to the outer spring constant", + " This confirms the common view [2, 19, 35, 39, 40] that the L- and R-states of the outer (D1) domain of flagellin enable the formation of the different polymorphic forms of the flagellum, whereas the inner (D0) domain plays a crucial part in selecting the energetically preferred helical configuration. We thank the Deutsche Forschungsgemeinschaft for financial support through the research training group GRK1558. The rigid-body networks of the model proteins exhibit helical forms. We describe here how we determine their curvatures and torsions from the position vectors ri of the centers of the rigid bodies indicated by the central black spheres in figure 4(b). First, we introduce a unit tangent vector to the filament at subunit or model protein i. We define the difference vectors to neighbor proteins in the same protofilament, ri+j+11 \u2212 ri+j , average them along a 1-start helix over all 11 protofilaments and normalize the resulting vector: ti = \u22115 j=\u22125(ri+j+11 \u2212 ri+j ) | \u22115 j=\u22125(ri+j+11 \u2212 ri+j )| . (A.1) The curvature of a space curve is calculated from the derivative of the tangent vector with respect to its arc length s, \u03ba = |dt/ds|. Discretizing the derivative gives the local curvature \u03bai at location ri : \u03bai = |ti+11 \u2212 ti | |ri+11 \u2212 ri | . (A.2) The torsion of a rod is defined by \u03c4 = d\u03b1/ds. Here, d\u03b1 is the angle by which the cross section of the rod is rotated about its tangent t when one proceeds a distance ds along the rod. Assuming small changes along one protofilament, we determine d\u03b1i by the cross product of the radial unit vectors ez i , ez i+11 of neighboring subunits in one protofilament (indicated by the z axis in figure 4(b)) and project it onto the local tangent vector ti . Thus, the local torsion \u03c4i becomes \u03c4i = ( ez i \u00d7 ez i+11 ) \u00b7 ti |ri \u2212 ri+11| . (A.3) Finally, we determine the curvature \u03ba and the torsion \u03c4 of the model filament consisting of N subunits by averaging over the local values of the 11 most centrally located subunits: \u03ba = N/2+5\u2211 N/2\u22125 \u03bai, \u03c4 = N/2+5\u2211 N/2\u22125 \u03c4i . (A.4) This algorithm has been tested at model filaments with known curvature and torsion. [1] Berg H C 2003 E. coli in Motion (Biological and Medical Physics, Biomedical Engineering) (Berlin: Springer) [2] Namba K and Vonderviszt F 1997 Molecular architecture of bacterial flagellum Q" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003466_j.mechmachtheory.2019.06.017-Figure18-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003466_j.mechmachtheory.2019.06.017-Figure18-1.png", + "caption": "Fig. 18. Coordinate systems used in the pinion generation by a worm.", + "texts": [], + "surrounding_texts": [ + "In this section, the pinion generation with double-crowned tooth surfaces is achieved. Two methods are presented for this purpose [29] . In both approaches, the profile crowning of the pinion is performed by a profile crowned worm, which is generated by a profile crowned rack ( Section 2 )." + ] + }, + { + "image_filename": "designv11_33_0002636_j.mechmachtheory.2018.05.003-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002636_j.mechmachtheory.2018.05.003-Figure4-1.png", + "caption": "Fig. 4. Kinematics of apex seal.", + "texts": [ + " (3) or (4) as follows: \u03b80 = sin \u22121 ( \u2212 L 4 r + \u221a L 2 ( 4 r ) 2 + W g 4 r ) = sin \u22121 ( \u2212 1 4 b + \u221a 1 ( 4 b ) 2 + W g 4 Lb ) (7) With the case of the circolima\u00e7on and the lima\u00e7on-to-circular machines, where the rotor lobes are manufactured of circular arcs, the truncation, z , can be calculated as: z = L \u2212 C r \u2212 \u221a R lc 2 \u2212 ( k + W g 2 )2 (8) Let k s be the seal spring stiffness, and k w be the equivalent stiffness of the wall of the machine housing ( k w k s ); and let I be the instantaneous centre as shown in Fig. 4 . Of note is that local deformation at the seal-housing contact point has been taken into account in the equivalent stiffness coefficient, k w . Given that the initial deflection of the seal spring, \u03b4s , is known at the rotor radial displacement \u03b8 = 0 ; \u03bb measures the angle between the rotor chord, p 1 p 2 , (as shown in Fig. 1 a) and a line that connects the correspondent point of p 1 on the housing to the instantaneous centre, I (which is shown in Fig. 4 ). If \u03bc was the coefficient of friction between the housing and the seal, the initial forces exerted by the machine 1 housing and the spring on to the seal can be calculated as: { F cos \u03bb( 1 \u2212 \u03bc1 tan \u03bb) = k w \u03b4w F s = k s \u03b4s (9) where \u03bb = tan \u22121 ( 2 r cos \u03b8 2 r sin \u03b8+ L ) . At the position \u03b8 = 0 , \u03bb = \u03bb0 = tan \u22121 ( 2 r L ) \u03b4w and \u03b4s are the initial deflections of the housing wall and the spring, respectively. Initially k s \u03b4s = k w \u03b4w Prior to the formulation of seal kinematics, some reasonable approximations need to be made" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001166_power2010-27012-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001166_power2010-27012-Figure1-1.png", + "caption": "Figure 1. A typical C-gear: 24-tooth pinion with a face advance of one pitch, face width of 20 x module, and tooth inclination angle of about 32\u02da at the side planes", + "texts": [ + " Gears for heavy power transmission must not be confined to the limitations imposed by spur and helical gearing. A parallel-axis gear pair with connectivity of more than three and without having any of its constraints in either of the above described ways of action would be very advantageous. Curvedtooth gears, C-gears for short, are later shown to satisfy this need. These are cylindrical gears with their teeth curved in the lengthwise direction, featuring convex and concave flanks, as the sample in Fig. 1 depicts. However, C-gears have been much more often portrayed than produced, and some of the disclosed drawings even reflect an awareness of the suitable proportions. But these gears are not of one type; a glance at Fig. 1 may indicate any one out of a dozen C-gear types. The literature abounds in suggestions of how the kinematics and/or machines for cutting and/or grinding C-gears could be configured. In the enumerative study by Arafa [13] the suggested geometries of C-gears were classified into 11 types, while Bedewy [14] made a comparative study of them, including one further possibility. Some of those types are of subordinate relevance, two of them being of \u2018approximate\u2019 geometry, and will be omitted from the following discussion" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002472_rpj-06-2017-0120-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002472_rpj-06-2017-0120-Figure5-1.png", + "caption": "Figure 5: Basic and subtraction cubes (Brajlih et al., 2010).", + "texts": [ + " During this research, manufacturing times were acquired from the software package of the Mlab CUSING\u2122 machine. It was presumed that the software-predicted manufacturing times are accurate enough for the purpose of this experiment. The basis of the test part\u2019s definition is a cube with an edge length a. The cube\u2019s edges are parallel to coordinate system\u2019s directions, and one of the corners concourses with the coordinate origin. From this cube, a smaller cube volume (with parallel edges) is subtracted, in order to obtain a part with resulting wall thickness in limits of a/10 and a/3 as shown in Figure 5. D ow nl oa de d by I N SE A D A t 0 0: 13 2 6 A pr il 20 18 ( PT ) To measure the achievable speeds of the machine, we have used an experiment based on a simple 2k factorial design principle requiring the test to be performed at combinations of high and low levels of both influential factors. The variation in volume ratio is achieved by changing the test part\u2019s wall thickness between the recommended limits of a/10 and a/3, consequently, designing two test parts with volume ratios of 0.25 (low level) and 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002839_j.mechmachtheory.2018.06.021-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002839_j.mechmachtheory.2018.06.021-Figure6-1.png", + "caption": "Fig. 6. Structural schematic for inner cone CVT.", + "texts": [ + "lee@gmail.com (Q. Li). https://doi.org/10.1016/j.mechmachtheory.2018.06.021 0094-114X/\u00a9 2018 Elsevier Ltd. All rights reserved. Nomenclatures \u03b8 half cone-angle of roller \u03b3 titling angle of roller D distance ( Fig. 3 ) r radius of circle ( Fig. 3 ) C cone apex P arbitrary point ( Fig. 3 and Fig. 5 ) v movement speed of the cone apex \u03c9 angular speed of axis t time parameter \u03bb individual parameter that represents the length of CP A and B contact points on input disc and output disc A\u2019 and B\u2019 foot points ( Fig. 6 ) A\u201d and B\u201d foot points ( Fig. 6 ) \u03c9 i and \u03c9 o rotational speeds of input disc and output disc \u03c9 r rotational speed of roller i speed ratio R length ( Fig. 7 ) \u03c9 spin spin speed \u03c9 dg rotational speeds of driving component \u03c9 dn rotational speeds of driven component \u03b8dg angle between the contact area and the rotational axis of the driving component \u03b8dn angle between the contact area and the rotational axis of the driven component \u03c3 spin spin ratio \u03c9 spinT and \u03c3 spinT spin speed and spin ratio of half-toroidal CVT \u03c9 spinL and \u03c3 spinL spin speed and spin ratio of logarithmic CVT \u03c9 spinI and \u03c3 spinI spin speed and spin ratio of inner cone CVT \u03c9 spinO and \u03c3 spinO spin speed and spin ratio of outer cone CVT e the distance of the toroidal cavity from the disc rotation axis k aspect ratio Cr in creep coefficient at input contact point Cr out creep coefficient at output contact point \u03bcin tangential force coefficient at input contact point \u03bcout tangential force coefficient at output contact point \u03c7 in spin momentum coefficient at input contact point \u03c7out spin momentum coefficient at output contact point t in input traction coefficient t out output traction coefficient \u03b7 total efficiency direction at the point of contact [17,18] ", + " In contrast, when v / \u03c9 < 0 then y < 0, the envelope is in quadrants III and IV and its second derivative is smaller than zero. Thus, the concavity or convexity of the derived envelope is unsuitable for the contact of the introduced rolling cone, as shown in Fig. 5 (a). To overcome this drawback, a particular structure is needed. As shown in Fig. 5 (b), the straight lines, which are the generatrices of roller, are always under the envelopes. Furthermore, the lines with positive and negative slopes should be used in pairs. Therefore, the contact surface of the roller should take the form of an inner cone. As shown in Fig. 6 , two inner cone rollers are introduced with rotation shafts bolted together at the intersection point of the generatrices of the inner cone surfaces. On the left of the roller, an input disc is placed with a contact surface generated by the derived envelope. Similarly, an output disc is placed on the right of the roller. To regulate speed ratio, the intersection point should be moved along the disc rotation axis, and the rollers will then tilt automatically because of the contact. At the same time, the axial positions of the discs are fixed. This scenario differs from the zero-spin logarithmic CVT, in which all the positions of the components need to be changed. Furthermore, as discussed in Section 1 , the logarithmic CVT needs a synchronous controlling motion. With regard to the inner cone CVT, only a simple controlling motion is necessary because the axial fixed input and output discs are expected to reduce the control difficulty. As shown in Fig. 6 , the normal forces at points A and B between the roller and the discs cannot point to a same point on the roller rotation axis. Thus, a net force moment on the roller appears which needs to be equilibrated, meaning that a particular ratio changer is required to be designed. The formula of speed ratio, which represents the ratio of input speed to output speed, is also derived. As shown in Fig. 6 , A is the contact point between the roller and the input disc, while B is the contact point between the roller and the output disc. Segments AA\u2019 and BB\u2019 are perpendicular to the disc rotation axis at A\u2019 and B\u2019 , respectively. Segments AA\u201d and BB\u201d are perpendicular to the roller rotation axis at A\u201d and B\u201d, respectively. Finally, C is the cone apex. The length of segment AA\u2019 can be derived from Eq. (15) \u2223\u2223AA \u2032 \u2223\u2223 = v \u03c9 cos 2 (\u03b8 + \u03c9t) , (23) We can infer from the geometry that \u2220 A \u2032 A C = \u03b8+ \u03c9t. (24) Thus, | AC | = v \u03c9 cos 2 (\u03b8 + \u03c9t) cos (\u03b8 + \u03c9t) = v \u03c9 cos (\u03b8 + \u03c9t) , (25) \u2223\u2223AA \u2032\u2032 \u2223\u2223 = | AC | sin \u03b8= v \u03c9 cos (\u03b8 + \u03c9t) sin \u03b8 ", + " 11 , the spin speed of the logarithmic CVT between the input disc and the roller can be derived as \u03c9 spinL = \u03c9 i cos ( \u03b8 \u2212 \u03b3 ) \u2212 \u03c9 r cos \u03b8 (48) where \u03c9 spinL is the spin speed of the logarithmic CVT between the input disc and the roller. We assume a no-slip condition in the traction between the input disc and the roller, that is, the linear speeds of the input disc and the roller are the same. \u03c9 i cos ( \u03b8\u2212\u03b3 ) = \u03c9 r cos \u03b8 (49) By substituting Eqs. (48) and (49) into Eq. (44) , we can obtain \u03c3spinL = cos ( \u03b8 \u2212 \u03b3 ) \u2212 cos ( \u03b8 \u2212 \u03b3 ) cos \u03b8 cos \u03b8 = 0 , (50) where \u03c3 spinL is the spin ratio of logarithmic CVT between the input disc and the roller. The logarithmic CVT can also transmit power without the spin motion. In accordance with Fig. 6 , the spin ratio of the inner cone CVT between the input disc and the roller can be written as \u03c3spinI = cos ( \u03b8+ \u03c9 \u00b7 t ) \u2212 \u03c9 r \u03c9 i sin \u03b8, (51) where \u03c3 spinI is the spin ratio of the inner cone CVT between the input disc and the roller. By substituting Eq. (28) into Eq. (51) , we obtain \u03c3spinI = cos ( \u03b8 + \u03c9 \u00b7 t ) \u2212 cos ( \u03b8 + \u03c9 \u00b7 t ) sin \u03b8 sin \u03b8 = 0 . (52) In accordance with Fig. 9 , the spin ratio of the outer cone CVT between the input disc and the roller can be written as \u03c3spinO = cos ( \u03b8+ \u03c9 \u00b7 t ) \u2212 \u03c9 r \u03c9 i sin \u03b8, (53) where \u03c3 spinO is the spin ratio of the outer cone CVT between the input disc and the roller", + " [39,40] observed that the rollers always spall from the surface crack before the other components are damaged, which they attributed to the rollers that constantly accumulate contact fatigue damage. In contrast, the discs undergo discrete damage at each speed ratio because the positions of the contact points relative to the rollers and discs ( Fig. 4 ) are fixed and unfixed, respectively. In Fig. 11 , the contact ranges of the logarithmic CVT components are same as those of the half-toroidal CVT. On this basis, the durability performance of the logarithmic CVT has not been improved relative to the half-toroidal CVT. In Fig. 6 , the contact positions relative to the inner cone roller and disc both vary with the speed ratio. Thus, all regions of the working surfaces of the inner cone roller and the disc may experience contact. A similar conclusion can be derived with mathematical equations. For instance, in Eq. (10) , \u03bb is the independent variable that represents the distance from the cone apex to the contact point, similar to the one depicted in Fig. 5 . Additionally, in Eq. (12) , \u03bb varies with parameter t , which is the independent variable of speed ratio in Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003882_s11071-019-05343-5-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003882_s11071-019-05343-5-Figure2-1.png", + "caption": "Fig. 2 Schematic diagram of forces acting on ceramic ball", + "texts": [ + " The radiation noise of the bearing is caused by the superposition of noises of all components, i.e. balls, inner ring, cage, and outer ring. The vibration characteristics of each component are analysed below. 2.1 Differential equations of vibrations for rolling element As an important component of angular contact ball bearings, the balls contact all components, i.e. the cage, inner ring, and outer ring. It is supposed that all balls are equal in mass and size, and the sizes of the cage pockets, which are distributed uniformly along the circumference, are the same. Figure 2 illustrates the forces acting on a ball when the ceramic ball bearing operates at a high speed. In Fig. 2, \u03b1i j and \u03b1o j are the contact angles between the j th ball and inner and outer raceways; Qi j and Qo j are the normal contact forces between the j th ball and inner and outer raceways, respectively; T\u03b7i j , T\u03b7o j , T\u03be i j , and T\u03beo j are the traction forces of the contact surface between the j th ball and raceways; Qcx j , Qcy j , and Qcz j are the decomposition components of the collision force between the j th ball and cage; Gby j and Gbz j are the decomposition components of the gravity of the j th ball; P\u03b7 j and P\u03be j are the friction forces acting on the surface of the j th ball, including rolling and sliding friction forces; Fbx j , Fby j , and Fbz j are the components of the hydrodynamic force acting on the centre of the j th ball; F\u03b7i j , F\u03b7o j , F\u03be i j , and F\u03beo j are the hydrodynamic friction forces at the lubricant inlet of the contact zone of ball and raceways, including rolling and sliding friction forces; Jx , Jy , and Jz are the components of the moment of inertia for a ball rotated around its own centre; \u03c9x j , \u03c9y j , and \u03c9z j are the components of the spin angular velocity of the j th ball in its own coordinate system; and \u03c9\u0307x j , \u03c9\u0307y j , and \u03c9\u0307z j are the components of the spin angular acceleration of the j th ball in its own coordinate system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002307_012055-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002307_012055-Figure1-1.png", + "caption": "Figure 1. The sample shell.", + "texts": [ + " Phenomena of this kind are often accompanied by an attached mass to the plate or a system of attached masses. Although this circumstance has not been fully studied, plate operation is present. And interest in analyzing the dynamics of thin plates is very high in broad areas of activity. [1-4] To investigate the effect of the presence of an attached mass or a system of attached masses, a special experiment was conducted in the laboratory of building structures of KnASTU. The prototype, located in a special bench (Figure 1), is equipped with an attached mass, was subjected to an additional system of attached masses, of a different nature, in the presence of natural oscillations. This work reflects the investigation of the natural oscillations of a thin rectangular plate in a plane that is hinged on both sides. Equations of natural oscillations of the plate are obtained according to the generally accepted theory of plate oscillations, as well as experimental data reflecting the dependence of the effect of the attached mass and the system of attached masses on the natural frequencies of the shell oscillations. Oscillations with moderate amplitudes of natural oscillations were decomposed, 2 according to the equations [5,6]. A discrete nonlinear model of vibrations of a thin plate supported at the edges, obtained during research, was investigated using a method of many scales. The sample is a thin plate, rectangular in plan, made of galvanized steel. The composition of the plate is given in Table 1. The geometric characteristics of the object: L = 890 mm, B = 370 mm, H = 0.4 mm. A sample is shown in Figure 1 The sample consists of grade 3 steel, semi-boiling. The shell model is fixed in a steel stand. This stand has the form of a table made of equal-angled corners L45x3 steel st3sp, especially for this experiment. Hinge support is realized with the help of glass plates 30x30 mm 4 mm thick, laid in the form of a corner. The boundary conditions are as close as possible to the real ones. The attached mass is the accelerometer \u0412\u0421110, located on the sample according to figure 1. Accelerometer \u0412\u0421110 measures the oscillation frequency with maximum accuracy. The accelerometer BC 110 transmits the readings to the signal amplifier, which, amplifying the signal, transmits further along the circuit to the analog-to-digital converter, which transmits the signal further to the personal computer. The thermometer is located closest to the shell, but does not touch it, transmits the readings on the software \"Zet-lab\", where in the compartment with the data of the oscillations the dependence of some parameters on others is reflected, in the real time" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002933_icelmach.2018.8507125-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002933_icelmach.2018.8507125-Figure1-1.png", + "caption": "Fig. 1. Structure of the electric motors with one-phase open-circuit fault (a) Traditional three-phase topology; (b) Modular topology (4 units).", + "texts": [ + " It combines the superiority of the mature Analysis of Vibration in Modular Fault-tolerant PMSM under One-phase Open-circuit Fault Zaixin Song, Yulong Pei, Yi Li, Shibo Li and Feng Chai, Member, IEEE R 978-1-5386-2477-7/18/$31.00 \u00a92018 IEEE 2565 technique for traditional three-phase motors and the faulttolerant ability. It reduces the current level that each power device unit in the system needs to support. The electromagnetic, physical and thermal isolation can be realized, which can be more effective and more reliable to deal with possible fault conditions (see Fig. 1), compared with traditional topologies. It is easy to cut off bad units and to reconstruct the normal windings. Applied in in-wheel motors, the modular structure can save more space for EVs and make the system more convenient to integrate. PMSM with modular structure may be faced with different kinds of fault conditions. Nonetheless, the one-phase OC fault, which is discussed in this paper, is the fundamental and most common fault condition. III. THEORETICAL ANALYSIS It is well-known that air-gap radial force density is the main source of electromagnetic vibration" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001216_icc.2012.6363714-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001216_icc.2012.6363714-Figure1-1.png", + "caption": "Fig. 1. Tilted antenna pattern.", + "texts": [ + " While current draw data is not available for all 16 levels, a second order polynomial interpolation is used to determine the missing values. These currents are multiplied by the supply voltage and divided by data rate to find the transmit energy consumption per bit. It is assumed that the Crossbow motes are connected to a standard dipole antenna. The gain of this antenna is omnidirectional with azimuth angle and changes with elevation angle according to [17] G(\u03b8) = 1.5 cos2(\u03b8) (1) where \u03b8 is elevation angle and \u03b8 = 0 corresponds to the horizontal plane. As a result, the gain of a dipole assumes a toroid shape as shown in Fig. 1. When the dipole antenna lands with a random orientation, the ground can be represented by a plane that intersects the antenna pattern at a random angle \u03b1 with respect to the vertical axis of the antenna gain pattern. As shown in Fig. 1, we can use this ground plane to take a cross section of the toroid dipole pattern. This allows us to use the following method to calculate the 2D node antenna gain. First, define \u03c6 as the azimuth direction of transmission in the ground plane. Utilizing the geometry of right angle triangles, the gain in this direction can be written as G(\u03c6, \u03b1) = 1.5 cos2 [ sin\u22121 (sin(\u03c6) sin(\u03b1)) ] (2) By default, the expression in (2) is defined to have maximum gain in the direction of \u03c6 = 0 and \u03c6 = \u03c0. A random angle offset \u03b6 uniformly distributed over [0, 2\u03c0) is then added to \u03c6 to reflect the different orientations of the nodes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002889_0142331218794811-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002889_0142331218794811-Figure1-1.png", + "caption": "Figure 1. Neural emulator with fully connected structure.", + "texts": [ + " Fuzzy adapting rate for NE and its application are then cited. A simulation example is also proposed, illustrating the performances of the NE based on the fuzzy adapting rate compared with the case of classical choice of adapting rate and the efficiency of the proposed method is given. An experimental validation of the proposed method to chemical reactor is then presented. We end with our conclusions. Neural emulator for nonlinear systems The NE was developed with fully connected recurrent neural networks (Figure 1). Let us define NIN and NOUT as the inputs and outputs, respectively, where IN and OUT denote the set of inputs and outputs. We consider square systems NIN =NOUT =N , where IN = 1, . . . ,N and OUT = 1, . . . ,N . The total number of neurons Ne of NE is chosen equal to 2N . To avoid perturbing any output signal with input ones, any node is either an input or an output neuron but not both at the same time. In discrete time, the dynamics of the Ne neurons of NE are given by the following equation (Atig et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000945_ssd.2010.5585533-Figure11-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000945_ssd.2010.5585533-Figure11-1.png", + "caption": "Figure 11. Mesh of the rotor claws and the associated magnetic rings.", + "texts": [ + " Moreover it is to be noted that the two claw plates are magnetically decoupled. In order to reduce the computation time, the FEA study domain is limited to a one pair of poles of the CPAES. Figures 7 and 8 show the stator and the rotor study domains, respectively. The meshing of both stator and rotor study domains has been automatically generated in the Ansys-Workbench environment. Figure 9 shows a mesh of the stator yoke and of the two associated magnetic collectors. A mesh of the stator lamination is illustrated in figure 10. Figure 11 shows a mesh of the rotor claws and the as sociated magnetic rings. 3.2.1. Main Flux Paths The flux moves through the magnetic circuit of the CPAES within three dimensional (3D) paths. These are described as follows: \u2022 axially in the stator yoke (see figure 12), \u2022 radially down through the collector and crossing the air gap down to the magnetic ring (see figure 12), \u2022 radially then axially in the magnetic ring (see figure 12), \u2022 axially-radially in the claws (see figure 13), \u2022 radially up in air gap facing the stator laminations (see figure 14), \u2022 radially up in the stator teeth and circumferentially in the stator core back (see figure 15), \u2022 radially down in stator teeth facing the two adjacent claws (see figure 16), \u2022 radially-axially in the adjacent claws (see figure 16), \u2022 axially then radially in the magnetic ring (see figure 17), \u2022 radially up through the air gap and the mag netic collector on the other side of the machine (see figure 17), \u2022 axially in the other side of the stator yoke (see figure 18)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002561_978-3-319-69748-2_2-Figure2.13-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002561_978-3-319-69748-2_2-Figure2.13-1.png", + "caption": "Fig. 2.13 a The telemetrically powered and interrogated implantable continuous blood glucose monitoring system with a three-electrode amperometric glucose sensor and ASIC and the front-end potentiostat electronics implemented as in b using a novel current mirror-based topology. Reprinted with permission from [207]. \u00a9 IEEE 2009", + "texts": [ + " In [205] perovskite-type oxide La0.88Sr0.12MnO3 (LSMO) nanofibers made by electrospinning and calcination, demonstrated wide linear range, high sensitivity, and a rapid response time. A xerogel-based amperometric sensor was developed in [206] (Fig. 2.12), which demonstrated a linear range of detection (>24\u201328 mM glucose), fast response time, and discrimination against interferences, such as acetaminophen, ascorbic acid (AA), sodium nitrite, oxalic acid, and uric acid. A remotely powered telemetric implantable sensor chip (Fig. 2.13) for subcutaneous applications with integrated electronics was presented in [207]. The potentiostat developed generated an output current which in turn was converted into an output frequency. The signal was used together with a load modulator circuit to transmit the data. The sensor was comprised of a Au counter electrode and an Au working electrode with an enzyme-immobilized layer on top and an Ag RE. The sensors had a glucose detection range of 0\u201340 mM. Monitoring glucose levels of diabetic patients is an enormous market, which is constantly increasing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003849_s00500-019-04452-y-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003849_s00500-019-04452-y-Figure1-1.png", + "caption": "Fig. 1 Cumulative probability distribution of the threshold", + "texts": [ + " More intuitively, whether each individual joins a collective action is related to whether the number of persons that joined the collective action reaches its own critical activation threshold. This phenomenon is prevalent in life, such as shopping, stock trading, strikes and incidents of mass riots. Granovetter (1978) proposed a classical threshold model to explain collective action. This model assumes that all persons can see others\u2019 behavior and assumes that the cumulative probability distribution of the threshold is function F(x), it is shown in Fig. 1. In Fig. 1, the x-axis represents the size of the threshold (the value of x-axis represents the percentage of expected number to the total number), and the y-axis represents the percentage of person\u2019s number that not large than x to the total person\u2019s number when the threshold value is x. The change trend of the whole threshold curve is conformed with the realistic law, the person\u2019s number that conforms the condition is very less when the threshold is low, but the person\u2019s number will gradually increase with the increasing of threshold, and the increase trend will gradually become larger" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001070_1.4002694-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001070_1.4002694-Figure1-1.png", + "caption": "Fig. 1 A general PM with n linear active legs", + "texts": [ + " However, since most linear legs are not perpendicular to their angular velocity, i T i=0 is not always satisfied. Hence, the solutions of dynamics analyses of PMs become approximate. This must impose difficulty on the control and analysis of kinematics/dynamics of PMs. For this reason, this paper focuses on the derivation of the velocity and acceleration of some linear legs with different structures as i T i 0 for limited-DOF PMs based on the unified kinematics/statics equations of the limited-DOF PMs 10 . 2 Kinematics/Statics of General Limited-DOF PM A general PM with n linear active legs is shown in Fig. 1. It includes a fixed base B, a moving platform m, and n linear active legs ri i=1,2 , . . . ,n 6 with the linear actuators or a passive constrained linear leg r0. Each of ri connects m at joint bi with B at joint Bi. r0 connects m at central point o with B at central point O. Let M be the DOF of PMs, m be a coordinate frame o-xyz fixed on m at central point o, and B be a coordinate frame O-XYZ fixed on B at central point O, be parallel constraint, and be a perpendicular constraint. The displacement equations of limited-DOF PMs are derived from Ref", + " 3 Angular Velocity/Acceleration of Linear Legs 3.1 Some Linear Legs With Different Serial Structure. A limited-DOF PM may include some different active/passive linear legs ri. Each of ri may be connected by various serial chains, as shown in Fig. 2 and Table 1. Each of U includes two crossed revolute joints Rij j=1,2 in leg ri. Let Rij be the unit vector of Rij, ij be the rotational angle of Rij, and ij be the angular velocity of Rij. The velocity vbi of leg ri at the connection point bi with m can be derived from Fig. 1 as follows: vbi = vri i + i ri i = vri i + i ri = v + ei 9 Here, i is the angular velocity of linear leg ri. Cross-multiplying both sides of Eq. 9 by i yields i vbi = i vri i + i ri i = i i ri i = ri i i \u00b7 i \u2212 ri i i \u00b7 i = ri i \u2212 ri i i \u00b7 i 10 In order to simplify the kinematics analysis of various legs, it is supposed that i ri in Refs. 12\u201314 . Thus, Eq. 10 leads to i \u00b7 i = 0, i = i vbi /ri 11 However, Eq. 11 is only suitable to some linear legs ri, which lower end is connected with the base B by a revolute joint R, and Transactions of the ASME 016 Terms of Use: http://www" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003466_j.mechmachtheory.2019.06.017-Figure20-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003466_j.mechmachtheory.2019.06.017-Figure20-1.png", + "caption": "Fig. 20. Installation of the face-gear drive with misalignments.", + "texts": [], + "surrounding_texts": [ + "Another method to obtain the longitudinal crowning tooth surface of the tapered involute pinion is provided in this section. The rotation modification for the pinion is performed instead of the worm plunging for longitudinal crowning of the pinion. The additional rotation of the pinion \u03d5p is changed to the following form: \u03d5 p = ( s w \u2212 s 0 ) \u00b7 tan \u03b2 r pp \u00b1 a r s 2 w (54) where a r is a parabola coefficient of the modified rotation for the pinion. The surface equation of the double-crowned tapered involute pinion can be derived by applying Eqs. (44) \u2013(54) ." + ] + }, + { + "image_filename": "designv11_33_0001845_pssb.201600335-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001845_pssb.201600335-Figure4-1.png", + "caption": "Figure 4 (a) Basic unit of 2D auxetic structure geometry of bow-tie shaped hexagon. (b) 3D unit results from rotating (a) with respect to center line.", + "texts": [ + " In the case, when b/a is \u221e (or a approaches to 0), Poisson\u2019s ratio of the given structure is always \u22121, regardless of the value \u03b8 or \u03b4/b 3 3D auxetic structure In Section 2. we introduced a 2D auxetic structure composed of sliding rigid units with bow-tie shape. In this section, we use hourglass shaped rigid units and construct two deformable structures that exhibit 3D auxetic behavior. In particular, we introduce two 3D auxetic structures, respectively, with regular triangular configuration and regular square configuration. As illustrated in Fig. 4, an hourglass shape is obtained by rotating a bow-tie shaped hexagon and hence the shape of an hourglass is determined by three parameters, for which we choose three parameters a, b, and \u03b8 illustrated in Fig. 4b. 3.1 Regular triangular configuration In this subsection, we consider a deformable structure constructed in regular triangular configuration with the hourglass shaped units as illustrated in Fig. 5. Figure 5a illustrates the structure in fully compressed position. In this figure, we note that a unit has contacts with six neighboring units. As illustrated in Fig. 5a, we denote by A, B, and C, the centers of the top three neighboring units and, by D, E, and F , the centers of the bottom three, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000139_isie.2009.5213061-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000139_isie.2009.5213061-Figure2-1.png", + "caption": "Fig. 2. Main Rotor and the Blade frame.", + "texts": [ + " The helicopter's main rotor is the basic rotary wing used to generate aerodynamic forces and moments. It is responsible for the majority of forces and moments and for the helicopter's specific characteristics. The main rotor is composed of two main blades in addition to a flybar. The ability of expressing the aerodynamic equations in many frames is very important because it will reduce the helicopter modeling complexity. Therefore, there is a need for a transformation from the Body frame to the Blade frame, Fig. 2. The air attacking the blades is due to different sources (namely, helicopter's body translation V, blade's high speed rotation V , and induced air flow Ii) and the resultant air velocity expressed in the Blade frame (10) is, of the main rotor thrust T, with N blades, near hover in terms of the collective pitch angle (Kim et al., 2004) is, and angular velocity of the main rotor (tp ,Q) . The cyclic and flybar commands are given as, Equation (12) represents the average thrust about one revolution of the rotor blades" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000895_ipin.2010.5647328-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000895_ipin.2010.5647328-Figure3-1.png", + "caption": "Figure 3. 3-D Structure of a conical linear polarizer", + "texts": [ + " Polarizing planes cannot discriminate the front surface and the rear surface. Since a linear polarizer is a flat sheet in its original shape, there are some problems when we use it for the angle sensor. To solve this problem, we invented a new form for the linear polarizer for the angular sensor. We cut out a semi-circular sheet from the flat sheet of the linear polarizer. We made a cone-shaped linear polarizer from the semi-circular flat sheet by attaching each straight edge mutually around the center of the straight line, as shown in Fig. 3. We assume that the apex of the cone-wise plane is the point O, the center axis is Z, the angle along the cone-wise plane from the line OA to an arbitrary position C is \u03b8, the rotating angle of the conewise plane around the Z-axis is \u03c6, the angle between the line of vision (ray axis) and the ground is \u03b3, and the inclined angle of the mother line is \u03b30. We can obtain the next formulas from the above geometrical relation, If we assume the relation \u03b3 = \u03b30 , we can get In the above relation, the angle \u03b8 refers to the angle of the polarizing plane based on the line OA. Hence we can extend the polarizing angle \u03b8 in the range of 180\u00b0 to the rotating angle of the cone-wise plane \u03c6 in the range of 360\u00b0. Since both angles show a one-to-one correspondence, we can determine the azimuth angle uniquely. When the sharpness of the cone-wise plane is the angle \u03b30, the extending shape of the cone-wise plane is determined uniquely, as shown in Fig. 3. We can then obtain the next relation Next we can show that the rotating angle of the conewise plane around the Z-axis can be modulated as the angle of the polarizing plane around the ray axis. That is, if we set a light source inside the cone-wise polarizer and observe from outside the cone-wise plane, the angle of the observed polarizing plane is proportional to the rotating angle of the cone-wise plane. IV. EXPERIMENTAL SETUP A. Introduction The proposed sensor consists of two parts. One is the transmitter (Tx), which emits polarized-modulating IR" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002965_itmqis.2018.8524922-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002965_itmqis.2018.8524922-Figure1-1.png", + "caption": "Fig. 1. The extrudate model developed.", + "texts": [ + " The students designed it using Solid Works software. This fiber (thermoplastic filament) is necessary for 3D printer processes via a fused deposition modeling (FDM) technology. FDM works on an \u2018additive\u2019 principle by laying down material in layers. Thus, FDM is also known as a solid-based AM technology. The machine modeling using this software enables to note all the dimensions of the components that the machine consists of and make all necessary drawings to manufacture them. The model that they have made, is shown in Figure 1. Moreover, Master Control program for this machine was made using Microsoft Visual Studio software in C #. Thus, an operator is able to manage all components of the conveyor line via the PC. The customer stated the requirements needed to that software. The developed model allows taking into account necessary specifications of the process equipment used in production. The 3D printer designed following that model, gives really good performance results, therefore, FDM technology, which guides the work, ensures efficient use of raw materials [13]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003466_j.mechmachtheory.2019.06.017-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003466_j.mechmachtheory.2019.06.017-Figure1-1.png", + "caption": "Fig. 1. Rotorcraft transmission system with concentric split-torque face-gear drives.", + "texts": [ + " Numerical examples are given to illustrate the results. \u00a9 2019 Elsevier Ltd. All rights reserved. Face-gear drives show a great significance in the mechanical transmission field, attributing to their advantages including the large contact ratio, high gear ratio, non-sensitive to misalignments and the ability for split of torque [1\u20133] . Face-gear drives have been successfully applied in the advanced rotorcraft transmission system [4,5] . In recent years, the rotorcraft transmission was updated to the concentric split-torque face-gear drives ( Fig. 1 ) in the Enhanced Rotorcraft Drive System (ERDS) Program. Compared with the conventional transmission system, the new design in ERDS increased 40% in the ratio of horsepower to weight [6] . The load sharing between the pinions and idlers should be at approximately equal to ensure the system operates properly. The load sharing performance of the new design will be influenced by the backlash of the facegear drives [1] . In addition, a proper backlash of the face-gear drive is also critical for a single mesh pair [7,8] , considering that the load capacity depends on a proper contact pattern" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000701_tmag.2012.2194791-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000701_tmag.2012.2194791-Figure1-1.png", + "caption": "Fig. 1. Schematic of the levitated object above the plate.", + "texts": [ + " In obtaining eddy forces the effect of eddy current generated in the permagnet is ignored, because the conductivity and size of the plate are much larger than those of the permagnet. Also, the permagnet is modeled as a magnetic dipole. By integration over magnetic dipole elements to find the force on a real permagnet, it is shown that the simple magnetic dipole model has sufficient precision. By comparing the results from this analytical solution with the results from the finite element analysis, the analytical solution is validated. In Fig. 1, the schematic of the magnetic levitation system is shown. An electromagnet, a levitated object which is a permagnet and a conductive plate are parts of this levitation system. The levitated object can be moved in direction by exerting the magnetic field of the electromagnet, and in and directions by moving the electromagnet. The eddy-current-based force may be generated by the variation of the position of the levitated object which creates eddy current in the conductive plate. This eddy current generates a magnetic field which opposes the motion of the permagnet (Lenz\u2019s law)", + " If the size of the permagnet is small with respect to its distance from the plate, the analytical solution obtained in this paper can be used approximately by assuming the levitated object as a magnetic dipole placed right at its center. Generally we can consider the levitated object as the sum of magnetic dipoles at different coordinates and obtain the eddy forces by integrating over these small magnetic dipole elements. But this increases the complexities. By calculating these integrals, it is seen that when , the distance between the center of the object and conductive plate (see Fig. 1), is such that and , the magnetic dipole approximation is valid with 5% relative error. In these equations is the object\u2019s diameter and is its height. In most situations is in the above ranges, so we can avoid these complexities in our calculations. A. Introduction In this section, an analytical solution is presented for obtaining eddy force due to the motion of the small permagnet. If , the relative permeability of the plate, is larger than 1, the plate will pull-in the levitated object toward itself, which is undesirable", + " From the quasi-static Maxwell\u2019s equations for linear conductors we have the following equation for magnetic vector potential in the conductive plate [15] (1) where is the magnetic vector potential, is the permeability of free space, and is the conductivity of the plate. In the nonconductive places the right side of this equation vanishes. Using this equation and boundary conditions the magnetic vector potential can be obtained in order to find the eddy forces on the object. First we analyze the motion of the permagnet in z direction. According to Fig. 1, we write the position of the permagnet as , where and are the initial position and velocity of the permagnet, respectively. It is known that a magnetic dipole with a moment is equivalent to a coil with a small radius , and a large current where . Hence, using this conversion we can make this problem similar to the electromagnet based eddy-current problem presented in the next section. The space current density, , in the cylindrical coordinate is (2) where is Dirac\u2019s delta function and is the angular unit vector in cylindrical coordinate" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002306_tmi.2017.2776404-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002306_tmi.2017.2776404-Figure3-1.png", + "caption": "Fig. 3. Exception case in which E(xi, xj) intersects with segment xixj and local coordinate system for a single edge on the boundary of the farthest point Voronoi diagram.", + "texts": [ + " 2, FVB(S) is a set of edges, and each edge lies on the midperpendicular of two vertices of the convex hull of S. To locate the optimal solution on FVB(S), we could firstly compute the optimal solution on each edge and then select the one with the maximum objective value. Denote the edge corresponding to vertices xi and xj by E(xi, xj). For most cases, segment xixj does not intersect with E(xi, xj). The only exception occurs when segment xixj is the diameter of the minimum enclosing circle (MEC) of S. One example is shown in Fig. 3(a). For this exception case, the intersection point is the MEC center and we could divide E(xi, xj) into two parts from the MEC center. Then all edges in FVB(S) lie on one side of the line passing by xi and xj . After the above process, we build a local coordinate system, which is shown in Fig. 3(b). The coordinate origin is located at the endpoint of E(xi, xj) which is closer to segment xixj . E(xi, xj) coincides with the negative x-axis. Let \u03b3 be the distance from origin to segment xixj , and r be the distance from origin to xi or xj . Denote the candidate solution by (\u2212\u03bb, 0) and the point p with polar coordinate (\u03c1, \u03b8). Then the objective function of the LBWMEC problem could be 0278-0062 (c) 2017 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002795_gt2018-77151-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002795_gt2018-77151-Figure6-1.png", + "caption": "Fig. 6 Photograph of test rig opened to showcase test bearing and support bearings.", + "texts": [ + " In this study, the film flow domain in a pad has 38 nodes in the circumferential direction, 101 nodes across the film, and 22 nodes in the axial direction. In addition, 9 nodes are used for a pad thickness. An iterative method searches for the journal position due to an applied load and convergence to steady film and pad temperature fields. The convergence criteria must satisfy a 0.1% difference in applied load and a difference in temperature within 0.1K. Fig. 5 shows a cross-section of the test rig for measurement of the static load performance of lubricated bearings, and Fig. 6 shows a photograph of the test bearing, rotor and support bearings. An electrical motor and gear box drive a rotor to a speed as high as 24k rpm. The rotor, 101.6 mm in diameter, is supported on a pair of tilting-pad journal bearing; and a four pad test bearing, flexure pivot type, is located at the rotor midspan. Table 2 lists the test bearing geometry and the lubricant material properties. End seals, labyrinth type, locate in each bearing cartridge and avoid oil contamination. The critical speed of the test rotor-bearing is well above the rotor operating speed range" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000689_12.879520-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000689_12.879520-Figure5-1.png", + "caption": "Fig. 5. Structure of the produced batteries: a) photograph of the paper battery, with seven elements integrated in series, during the measurement of open circuit voltage with a multimeter; b) Scheme of an element battery, anode/paper/cathode; c) Schematic of the integrated (in series) battery.", + "texts": [ + " The structures produced comprised Al/paper/Cu (10mm x 30mm) for the prototype and Al /paper/WO3/GZO (40mm x 30mm) for the discharge experiment. For the discharge experiment of the paper battery the anode of the device was connected to a 20M\u03a9 resistor. The positive terminal of the battery was grounded, while the load was connected to an inverting input of the electrometer. Its non-inverting input was grounded, thereby creating a virtual ground between the battery and the resistor. The information was acquired by the Global Labs DT2812a data acquisition board. For the series association (figure 5c) that powered the paper transistor, it was patterned on one side of a sheet of tracingpaper four rectangular anodes of aluminium with 12x30mm 2 , after which were intercalated with three rectangular cathodes of copper with the same dimensions. The U connections between the two electrodes were deposited in copper, Proc. of SPIE Vol. 7940 79400P-2 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 05/19/2015 Terms of Use: http://spiedl.org/terms with seven elements associated in series with an open circuit voltage varying from of 3", + " However, it should be notice that different structures can also be employed in order to improve the performance of the paper batteries. For instance the use of a WO3 interlayer between the paper electrolyte and the cathode contribute for enhancing Voc to 0.68 V. Replacing the metal of cathode by a transparent conductive oxide layer (GZO-gallium zinc oxide) we are able to produce an invisible cathode while the Voc of battery is improved to 0.75 V. The photograph and the scheme of the battery are depicted in figure 5. The paper was used without any treatment, before or after the thin films deposition. The thin films were produced by thermal evaporation of high purity (>99.99%) Cu and Al metal pellets. Therefore this all-solid-state battery is conceptually cheap and environmental friendly. As already stated, the performance of the paper batteries can be optimized by adjusting the cathode and/or anode materials, as well as the paper matrix (regarding to its porosity, density, thickness and constitution chemical additives)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002061_jmech.2017.48-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002061_jmech.2017.48-Figure1-1.png", + "caption": "Fig. 1 Physical configuration for a short journal bearing with a non-Newtonian fluid.", + "texts": [ + " Based on the Rabinowitsch fluid model, a non-Newtonian dynamic Reynolds equation for journal bearings will be derived and applied to predict the linear dynamic characteristics of short journal bearings. Analytical expressions of the dynamic coefficients and the stability threshold speed are obtained. Comparing with the Newtonian-lubricant case, the dynamic stiffness coefficients, the dynamic damping coefficients and the threshold speeds are presented and discussed for different values of the nonlinear parameter. 2. ANALYSIS Figure 1 shows the configuration for a journal bearing at the mid-plane 0z with a non-Newtonian Rabi- * Corresponding author (jrlin@nanya.edu.tw) https:/www.cambridge.org/core/terms. https://doi.org/10.1017/jmech.2017.48 Downloaded from https:/www.cambridge. rg/core. Cornell University Library, on 07 Jul 2017 at 14:52:05, subject to the Cambridge Core terms of use, available at 2 Journal of Mechanics nowitsch fluid. The bearing length is L , and the journal shaft with radius jr rotates within the bearing housing with an angular velocity " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002870_joe.2018.8345-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002870_joe.2018.8345-Figure1-1.png", + "caption": "Fig. 1 Structure model of 4-pole 30-slot motor", + "texts": [ + " This motor is transformed to a rotating armature permanent magnet synchronous motor (RAPMSM) when the permanent magnets are placed on its stator, the armature windings are placed on the rotor, and a sinusoidal voltage excitation is applied in the armature windings. In RAPMSM, the air-gap fundamental magnetic field keeps unmoved in space. This is one of the significant differences from PMSM. In this paper, for the 4-pole RAPMSM, the analytical expression of the radial excitation force wave was deduced, and the order and frequency of the excitation force wave were obtained. The analysis of different numbers of slots and different polar arc coefficients was carried out. A schematic diagram of a 4-pole, 30- slot RAPMSM is shown in Fig. 1. Open-circuit field of an RAPMSM is decided by permanent magnet excitation and rotor slots together. When ignoring magnetic saturation, air gap magnetic field is expressed as J. Eng., 2018, Vol. 2018 Iss. 17, pp. 1903-1908 This is an open access article published by the IET under the Creative Commons Attribution-NonCommercial-NoDerivs License (http://creativecommons.org/licenses/by-nc-nd/3.0/) 1903 b(\u03b8, t) = f (\u03b8, t) \u22c5 \u03bb(\u03b8, t) (1) Air-gap permeance of motors with smooth stator and slotted rotor is expressed as \u03bb(\u03b8, t) = \u039b0 + \u2211 k = 1 \u221e \u039bkcos kZ\u03b8 (2) where Z is the number of rotor slots; \u039b0 and \u039bk are the amplitude of average and kth harmonic magnetic permeability, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001553_cdc.2010.5716942-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001553_cdc.2010.5716942-Figure1-1.png", + "caption": "Fig. 1. Line parameters.", + "texts": [ + " Finally, Section VI contains some conclusions and lines of future research. In this paper, a line segment in the plane is described by its endpoints coordinates, xf = [xA yA xB yB ] T or, alternatively, by geometric parameters as xp = [\u03b1, \u03c1, dA, dB ] T , where \u03b1 \u2208 (\u2212\u03c0, \u03c0] is the angle between the x-axis and the normal to the line from the origin, \u03c1 \u2265 0 is the distance from the origin to the line, dA and dB are the distances (with proper sign) of the segment endpoints A and B to the point of incidence of the normal to the line (see Figure 1). The measurement subspace, or M-Space [5], is a feature representation that attaches a local frame to each feature element, allowing for a generic treatment of many types of features. The parameters xp represent the M-Space coordinates of a feature and are expressed in the corresponding reference frame. Notice that each xp lives in a different reference frame. 978-1-4244-7746-3/10/$26.00 \u00a92010 IEEE 3826 Let \u03b4xp denote the change in the M-Space corresponding to a small change in the feature coordinates \u03b4xf " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003893_0954406219888240-Figure22-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003893_0954406219888240-Figure22-1.png", + "caption": "Figure 22. Parameters of the revolute joint: (a) o-uw plane and (b) o-vw plane.", + "texts": [ + " Comparing to the workspace analysis, it is worth noting that the rotation angle has been newly defined for the convenience of description. When the constraint is pure force or pure moment, the joint clearance can reach the maximum value. According to the first contact model and their hypotheses of revolute joints,21,22 if the constraint is a pure force and along the u-axis, v-axis and w-axis, respectively, the corresponding clearances can be obtained easily and they are c2 and c3, respectively, where c2 and c3 are the gaps shown in Figure 21. As shown in Figure 22, if the constraint is a pure moment and around the u-axis, the clearance can be solved by dR dR 2c2 cos\u00f0 \"\u00de \u00bc LR tan\u00f0 \"\u00de \u00f021\u00de where D is the contact deformation around the u-axis, LR is the equivalent length, which depends on the actual structural parameters, and dR is the inner diameter of the housing. Thus \" LR \u00fe ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi L2 R \u00fe 4dRc2 p 2dR \u00f022\u00de Similarly, when the constraint is a pure moment and around the v-axis, it has the same clearance" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000770_lindi.2012.6319490-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000770_lindi.2012.6319490-Figure1-1.png", + "caption": "Figure 1. Theoretical scheme of bicycle/car [1]", + "texts": [ + " The goal can be to reach a certain point, trajectory tracking, tracking a moving target, tracking a line, turning into a predetermined direction (angle), etc. In addition, you can choose one from the number of controllers. II. BICYCLE MODEL OF CAR-LIKE ROBOT The theoretical description and Simulink implementation of the kinematic model has been given in [1] in more details. The example, parameters of kinematic model that were used for comparison can be found in Table I. The scheme of theoretical kinematic model of a bicycle can be seen in Fig. 1. Besides the kinematic model we developed a minimal dynamic model of the bicycle which is described in this paper. The whole control loop system can be seen in Fig. 2, which also includes the main subject of research: route controller. The minimal dynamic model means for us that it can neglect the inertial and friction effect of the wheels, furthermore the dynamic reaction of steering and turning. Thus the following main part in the dynamic model remains: 1. DC motor, which drives the rear wheel or wheels of the robot-car" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003654_s11663-019-01670-5-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003654_s11663-019-01670-5-Figure6-1.png", + "caption": "Fig. 6\u2014Transient temperature distribution of the molten pool on the top and cross-sectional view as laser beam exposed direct exposure point F under various ets: (a) 45 ls, top view; (b) 45 ls, cross-sectional view; (c) 85 ls, top view; (d) 85 ls, cross-sectional view; (e) 125 ls, top view; and (f) 125 ls, cross-sectional view.", + "texts": [ + " It was clear that the temperature change of the direct exposure point F was more significant than that of the unexposed points (E and C) under different ets, owing to point F being directly affected by the laser beam. Therefore, it could be considered that the changed et had the greatest influence on the direct exposure position F compared with other unexposed positions. Since the influence of the laser beam on different positions was different, to accurately analyze METALLURGICAL AND MATERIALS TRANSACTIONS B the effect of process parameters in the point exposure scan pattern, the positions where the thermal behaviors were different should be studied if possible. 2. Temperature distribution Figure 6 depicts the transient temperature distribution of the molten pool on the top and cross-sectional views as the laser beam scanned direct exposure point F under various ets. It shows that the temperature distribution on the top surface seemed more intense at the front of the molten pool than that at the back at different ets (Figures 6(a), (c), and (e)). This indicated that the temperature gradient in front of the molten pool was larger. A main cause of the phenomenon was that the heat conductivity of the material was small in front of the molten pool", + " Generally, the thermal conductivity of aluminum alloy was obviously larger than that of titanium alloy, so the thermal effect induced by et was not significant. It was also seen that the simulated central peak temperature in the molten pool changed greatly with increasing et. The maximum temperature of the molten pool increased from 3641 K to 4200 K as the et was enhanced. This was because increasing et brought about a long interaction time between the powder particles and laser beam, resulting in greater input from the VED. 3. Molten pool dimension As shown in Figure 6, from both the top and cross-sectional views, with the increase of et, the higher temperature within the molten pool caused sufficient melting of the powder particles and a resultant larger molten pool. The specific change of molten pool dimensions with different ets is depicted in Figure 7(a). At an et of 45 ls, the calculated length, width, and depth of the molten pool were 150, 112, and 61 lm, respectively. As the et was increased by 40 ls, the enhanced length, width, and depth of the molten pool were 175, 148, and 68 lm, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001487_j.sna.2010.07.024-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001487_j.sna.2010.07.024-Figure9-1.png", + "caption": "Fig. 9. Coordinate system set to the PVCCF, internal forces and moments.", + "texts": [ + " Hence, the release of motion restriction in the PVC molecular chain and the induction of compression force on the PVC plate caused a rearrangement of the PVC molecular chain orientation, resulting in the anisotropic molecular structure of the PVC plate. Consequently, the deflection property change of the PVCCF was induced after the initial thermal cycling treatment. Bearing in mind the deflection property change of the PVCCF by the initial thermal cycling, we carried out a theoretical analysis of the PVCCF deflection [15]. Simultaneous equations represented by Eqs. (1)\u2013(5) are derived for the arbitrary cross section of the PVCCF, where the coordinate system is set to the PVCCF, as illustrated in Fig. 9. The balance of axial forces, where Ni (i = U, I, L) represents internal forces exerted on the neutral axis of the PVC plate, the glue layer, and the CFRP plate, respectively, is shown by the following equation: NU + NI + NL = 0 (1) The balance of moments Mi (i = U, I, L) represents the internal moments exerted on the PVC plate, the glue layer, and the CFRP plate, respectively; ti represents the thickness of those layers; and b and m represent the width of the PVCCF and the internal moment per unit width induced by external 360 H" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002839_j.mechmachtheory.2018.06.021-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002839_j.mechmachtheory.2018.06.021-Figure2-1.png", + "caption": "Fig. 2. Speed decomposition of rolling traction in different conditions: (a) the parallel zero-spin condition; (b) the intersection zero-spin condition; (c) the other conditions.", + "texts": [ + " The differences are the result of the inevitable shearing deformation of the oil film during torque transmission which means that the two types of losses are inevitable. However, spin loss generally dominates the power losses [19,20] but can be eliminated. In tribology, spin loss is caused by the different speed distributions in the contact areas. In geometry, spin loss can only be eliminated in two traction conditions [17] , namely, the parallel zero-spin condition and the intersection zero-spin condition. In Fig. 2 (a), the rotation axes of the driving and driven components are parallel with the contact area. Furthermore, the speed distributions of the two contact areas are even, which indicates the absence of spin motion. Fig. 2 (b) shows that the contact area and the two axes intersect at one point. The regularities of the speed distributions are also the same. Thus, the spin motion can also be eliminated. Fig. 2 (c) shows the typical operating state in toroidal CVTs (i.e. the speed distributions differ). An angular speed always remains after the pure rolling traction is extracted (i.e. the spin motion remains). This study investigates a new application of the envelope theorem to optimise the shapes of contact components in halftoroidal CVT. In this subsection of the paper, we present the pair of contact generatrices to be determined by calculating an envelope of a family of circular curves which are regarded the contact generatrices of original half-toroidal CVT" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000860_j.sysconle.2010.03.011-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000860_j.sysconle.2010.03.011-Figure5-1.png", + "caption": "Fig. 5. The prototype of miniature ducted-fan aerial vehicle described in [28].", + "texts": [ + " In fact, by denoting with u the angle of attack of the control vanes, it turns out that the torque generated by the flaps can be written as Nc = kcw2Pu while the resistant aerodynamic torque of the propeller can be approximated by Np = \u2212kDw2P , with kc and kD constant positive coefficients. Denoting with x1 the heading angle and with x2 the heading angular velocity, the heading dynamics can then be written as x\u03071 = x2 Iz x\u03072 = kcw2Pu\u2212 kDw 2 P (25) where Iz denotes the inertia around the vehicle vertical axis. The design of the heading control law for the ducted-fan consists of the choice of the input u in (25) in order to stabilize a given (constant) heading reference x\u03041. Observe that in order for (25) to be controllable a necessary condition is thatwP > 0 (see Fig. 5). 7.1. Design of the feedback control law For the purpose of designing a feedback control law we rewrite the system (25) as x\u03071 = x2 x\u03072 = a(t)(kcu\u2212 kD) (26) with a(t) := w2P(t)/Iz . Let us assume that a(t) > 0 for all t \u2265 0, which in practice requires the propeller angular velocity to be greater than zero, as happens during a standard flight. Moreover we assume that the outputs available for feedback are given by y1 = x1 + e1 y2 = x2 + e2 (27) where \u2016e1(t)\u2016\u221e \u2264 e1max and \u2016e2(t)\u2016\u221e \u2264 e2max are bounded measurement errors" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000238_icca.2010.5524364-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000238_icca.2010.5524364-Figure1-1.png", + "caption": "Fig. 1. LOS Fixed Coordinate", + "texts": [ + " Assume that both missile and target are particles in a 3- dimension space and define the LOS vector as r = rT \u2212 rM = \u03c1er, (1) where rT , rM are the position vectors of the target and missile in an inertial frame, respectively, \u03c1 is the range between target and missile and er the unit vector in the direction of LOS. The relative velocity and acceleration are r\u0307 = \u03c1\u0307er + \u03c1e\u0307r = vT \u2212 vM , (2) r\u0308 = \u03c1\u0308er + 2\u03c1\u0307e\u0307r + \u03c1e\u0308r = aT \u2212 aM , (3) where vT , aT , vM and aM are the velocity and acceleration vectors of target and missile, respectively. Assume that the angular velocity vector of LOS, \u2126, is orthogonal to LOS and and define a LOS fixed coordinate system (er, et, e\u2126) where et , e\u0307r \u2016\u2126\u2016 , e\u2126 , \u2126 \u2016\u2126\u2016 . (4) The geometry of the coordinate is shown in Fig. 1. If the accelerations of both target and missile are expressed in this (er, et, e\u2126) coordinate, aT , aTr er + aTt et + aT\u2126 e\u2126, aM , aMr er + aMt et + aM\u2126 e\u2126, 978-1-4244-5196-8/10/$26.00 \u00a92010 IEEE 866 it can be shown that the equations of motion are the followings [10], d dt \u03c1\u0307 = \u2016\u2126\u20162 + (aTr \u2212 aMr ), (5a) d dt \u03c1\u2016\u2126\u2016 = \u2212\u03c1\u0307\u2016\u2126\u2016+ (aTt \u2212 aMt ), (5b) along with one kinematic equation d dt \u03c1 = \u03c1\u0307. (6) The three scalar differential equations fully describe the relative dynamics between target and missile" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000810_j.partic.2010.10.002-Figure11-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000810_j.partic.2010.10.002-Figure11-1.png", + "caption": "Fig. 11. Trapping of a 650 m glass particle on the air\u2013water and decane\u2013water interfaces. (a) The dimensionless vertical positions (Z/R) as a function of time. Here Z is measured from the undeformed interface. Notice that the oscillatory behavior is similar to that in Fig. 5. (b) The contact lines on the air\u2013water and decane\u2013water interfaces for the particle.", + "texts": [ + " Notice that the frequency f oscillation for these cases was 20 Hz or larger, and therefore a igh speed camera was needed to see and analyze the motion. The photographs shown in Figs. 8\u201310 were taken from highpeed movies of particles undergoing adsorption at the fluid\u2013liquid nterfaces. These movies were also analyzed frame-by-frame to btain the dimensionless distance of the center of particles (Z/R) rom the undeformed interface as a function of time. The later results for a 650 m glass bead are shown in Fig. 11 for the ir\u2013water and decane\u2013water interfaces. Fig. 11(a) shows that the quilibrium height of the center of the particle relative to the ndeformed interface is lower on the air\u2013water interface. This ig. 13. The frequency of oscillation of spherical glass particles on the decane\u2013water nd air\u2013water interfaces versus the particle diameter. Both experimentally meaured frequencies and that given by Eq. (7) (indicated by \u201cth\u201d) are shown. he parameter values in Eq. (7) are assumed to be: p = 2600.0 kg/m3 and P \u2212 c = 16000 kg/m3; for the air\u2013water interface = 0.001 Pa s, 12 = 72.4 mN/m; nd for the decane\u2013water interface = 0.001 Pa s, 12 = 51.2 mN/m. w f a c a k a F t o nterface. The data were taken for 3 different test particles of the same approximate iameter of 550 m. an be also seen in Fig. 11(b) which shows that the particle floats n the decane\u2013water interface such that a smaller fraction of it s immersed in the water, whereas on the air\u2013water interface larger fraction of its lower surface is immersed in the water. his is expected since for the same floating height the buoyant eight of the particle on the air\u2013water interface is larger, and hus to balance its weight it is more immersed in the lower liquid. he angle of the three-phase contact line on the particle\u2019s surace is another important parameter, but its value is not known o us. It is noteworthy that even after the vertical oscillations of he particle subsided, its floating height on the decane\u2013water nterface slowly decreased before reaching a constant value. This s due to the partial pinning of the contact line on the partile\u2019s surface and the contact angle hysteresis (see Fig. 11(a)). his issue is discussed below in more detail. Also notice that he amplitude of oscillation of the particle was larger on the ecane\u2013water interface. This is because the densities of decane and ater are closer than the densities of air and water, and thereore the restoring buoyant force resulting from a displacement way from the equilibrium position is smaller for the decane\u2013water ase. When the diameter of glass particles in our experiments was pproximately 650 m or larger a significant fraction of the sprinled particles were not captured on the corn oil\u2013water interface nd those that were captured did not disperse", + " We believe that this is due to the fact hat when a particle smaller than 1 mm moved downward in the ecane\u2013water interface, the three-phase contact line became parially pinned on the particle\u2019s surface increasing the contact angle bove the equilibrium value. This in turn increased the vertical capllary force making the net vertical force on the particle zero even hough the particle was above its equilibrium height. As the conact line slowly moved down on the particle\u2019s surface, the contact ngle was reduced and the particle moved downward. A comparison of the two cases described in Fig. 11(a) also shows hat the time taken by the particle to oscillate once about the quilibrium height, i.e., the inverse of which is the frequency of scillation, is larger on the decane\u2013water interface. Since the effecive interfacial viscosity is larger when the upper fluid is decane and he interfacial tension and the buoyant weight of the particle for he decane\u2013water interface are both smaller than the correspondng values for the air\u2013water interface, the experimental result for he frequency of oscillation is consistent with Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002284_ssrr.2017.8088167-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002284_ssrr.2017.8088167-Figure9-1.png", + "caption": "Fig. 9 Landing position adjustment control", + "texts": [ + " To determine the position required, the height of the surrounding area can be detected using perception sensor and appropriate control of endeffectors. However, since there is error in the information acquired from perception sensor, it would be difficult to determine the position required accurately just by using the sensor alone. Therefore, based on feedback from the force sensor mounted on the robot's end-effectors, contact with an uneven surface can be detected and the appropriate landing position of the end-effectors can be determined. Fig. 9 shows the trajectory of the end-effector when landing position adjustment control is applied to the crawling motion. This control method is only applied during the phases where the robot's end-effectors are being lifted in the z-direction and during the phases where its torso is lifted in the z-direction. During regular operation, the phases where the end-effectors are traveling downwards in the z-direction halts when the reaction force in that direction detected by the force sensors exceeds a certain threshold value" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002282_978-3-319-66417-0_1-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002282_978-3-319-66417-0_1-Figure9-1.png", + "caption": "Fig. 9 Schematic diagram for non-covalent approach for synthesis of MIP", + "texts": [ + " Non-covalent Approach The frequently used method to prepare MIP is the non-covalent approach due to its simplicity. In this approach, the template and functional monomer form a complex through non-covalent intermolecular interactions. The special binding sites are formed by the self-assembly between the template and monomer, followed by a cross-linked copolymerization (Matsui et al. 2009; Lakshmi et al. 2009). The imprint molecules interact, during both the imprinting procedure and the rebinding, with the polymer via non-covalent interactions, e.g., ionic, hydrophobic and hydrogen bonding (Fig. 9). The drawback of the non-covalent approach is set by the peculiar molecular recognition conditions. Due to the random nature of complex formation which may lead to different orientations of the two species and therefore different types of imprinted cavities. The formation of interactions between monomers and the template is more stabilized in hydrophobic environment over polar environment. Some restrictions in the covalent binding between the monomer and the polymer, such as some molecules characterized by a single interacting group such as an isolated carboxyl, generally give imprinted polymers with very limited molecular recognition properties" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000151_iemdc.2009.5075258-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000151_iemdc.2009.5075258-Figure8-1.png", + "caption": "Fig. 8. Schematic view of the reduced model.", + "texts": [ + " As the working point of the PM shall not change the height of the magnet must not be altered als long as the width of air gap is the same. Therefore we have to shrink the PM\u2019s face as the length of the stacks is reduced. kSL \u2248 A\u03b4 APM \u00b7 hPM \u03b4\u2032\u2032 (15) rPM,a,red = \u221a lstack,red \u00b7 r2 PM,a \u2212 r2 PM,i lstack + r2 PM,i (16) The new model produces qualitatively the same torque as the full model but it is lstack,red/lstack times smaller because the flux is reduced in the same ratio. A scheme of the model is given in fig. 8. It is existing two, in axial direction extruded parts which are very short. The length does not matter as long results are within the calculation accuracy and are not affected by numerical errors. Left you can see the reduced permanent magnet. On the right side we have two possibilities to close the magnetic circuit. Either we don\u2019t place nothing within the area marked with a bright edge marked and use an ideal magnetic material for the stator yoke, or we fill this area with the real physical entities and leave an empty space between the inner radius of this area and the outer radius of the reduced magnet to avoid stray losses" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003240_s10921-019-0571-z-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003240_s10921-019-0571-z-Figure4-1.png", + "caption": "Fig. 4 Rolling element bearing geometry", + "texts": [], + "surrounding_texts": [ + "The fault detection experiments were carried out using a test rig which consists of a centrifugal pump as shown in Fig.\u00a03. The test bearing is one of the two support bearings which support rotor of a centrifugal pump. Four sets of bearings (SKF R7 NB 62) of motor pump were considered for the experimental instigations. The motor was operated under a speed of 1200\u00a0rpm. The sound and vibration signals were acquired simultaneously from a test bearing mounted in the experimental setup under the following conditions: (a) healthy bearings (b) inner race fault (c) outer race and (d) rolling element fault. In general, about 90% of total faults are related to inner race flaw, outer race flaw or rolling element flaw, rarely a cage defect also occurs in rolling element bearings [28]. Initially, the healthy bearing (new bearing) was fixed in the test rig. The vibration and sound signals were acquired simultaneously using B&K 4322 accelerometer and B&K 4117 microphone respectively, the accelerometer was mounted in radial direction to the central axis of bearing housing. In the present work, the pits were artificially introduced on inner race, outer race and rolling elements using electrical discharge machining method. The diameter and depth of the cylindrical pit is approximately 0.7\u00a0mm. The specifications and operating conditions of these bearings are given in Table\u00a01. The position of the microphone near to the bearing housing is important in sound measurements. Initially, the sound measurements were taken at various distances and directions; finally, the position was set at a distance of 5.5\u00a0cm near the bearing housing under near field condition, this method was successful and provided promising fault diagnostic information in our previous works [7, 9, 29, 31]. DACTRON FOCUS\u2014 F100 data acquisition system was used to acquire the sound and vibration signals at a sampling frequency of 16.4\u00a0k\u00a0Hz; these signals were amplified using a B&K 2626 signal conditioning amplifier. The time domain averaging method is most commonly widely used in fault detection of rotating machinery which increases the strength of vibration/sound data relative to the noise obscured in the signals. In the present study the vibration and acoustic signals are synchronously averaged to minimize the random noise by considering 16 sets of raw (9)h( ) = T \u222b 0 H( , t)dt 1 3 time data. The sound and vibration signals were acquired from the bearing setup under various operating conditions and stored in a personal computer for further processing. The local faults simulated on the components of rolling element bearing generate a certain fault characteristic frequencies based on their location and size i.e. dimensions of inner and outer races, number of rollers, shaft speed etc. The bearing geometry and fault locations on bearing components are depicted in Figs.\u00a04 and 5 respectively. These defects mimic contact fatigue faults which are quite common in medium/high speed rotating machines used in industrial applications. When the bearing rotates, each type of bearing defect will generate a particular frequency of impact vibrations. The healthy bearing was replaced by defective bearings which consist of inner race, outer race and ball faults. The signals were acquired for all the cases separately, under the same operating conditions. The characteristic bearing defect frequencies of inner race, outer race and balls are calculated by the Eqs.\u00a0(10), (11) and (12) respectively [19]. (10)fo = Nb 2 fr [ 1 \u2212 Bd Pd cos ] Hz (11)fi = Nb 2 fr [ 1 + Bd Pd cos ] Hz 1 3 (12)fb = Pd Bd fr [ 1 \u2212 ( Bd Pd )2 cos2 ] Hz where, Nb is the number of rollers in bearing, fr is rotating frequency of shaft, \u03b1 is contact angle and Bd and Pd are the rolling element diameter and pitch diameter respectively. For this test rig, the shaft frequency is 20\u00a0Hz. Journal of Nondestructive Evaluation (2019) 38:34 34 Page 8 of 23 The characteristic bearing defect frequency values of inner race, outer race and roller elements are 149\u00a0Hz, 84\u00a0Hz and 52\u00a0Hz respectively. The conventional signal processing techniques viz. statistical parameter and fast Fourier transform analyses methods are used to extract fault related features from time wave forms of sound and vibration signals. Further, in order to enhance diagnostic information, advanced signal processing techniques such as envelope and EMD based envelope and statistical parameter analysis methods developed in MATLAB 6.5 are used to assess the incipient bearing faults. Figure\u00a06 shows the signal analysis procedure used to detect incipient faults in rolling element bearing." + ] + }, + { + "image_filename": "designv11_33_0000227_sav-2010-0518-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000227_sav-2010-0518-Figure4-1.png", + "caption": "Fig. 4. Undamaged test pinion.", + "texts": [ + " In addition, a water-cooled magnetic brake, capable of producing an anti torque of up to 150 Nm, was connected to the output shaft of the gearbox to consume power and, consequently, to load gears within the gearbox. Moreover, an elastic coupling and V-belt units were employed in order to provide a much smoother power transmission through the gearbox and the other elements of the test rig. All these explanations are depicted in both Figs 2 and 3. In order to achieve a tooth fatigue crack (and, consequently, a tooth breakage) within a shorter time, the face width of the pinion gear at the first stage was reduced to 4 mm from its original value of 10 mm as shown in Fig. 4. This face width removal also permits the pinion to be tested at nearly twice of its nominal torque of 17 Nm. During the fatigue test, the speed of the input pinion was set to nearly 1566 rpm which yields a fundamental tooth-meshing frequency of 522 Hz for the first stage. The resulting vibrations were picked up by two accelerometers located around the bearing housing nearest to the input pinion as seen in Fig. 3. In addition, an inductive sensor producing one pulse per revolution was used to reference the angular position of the test gear" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001347_s10846-012-9725-2-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001347_s10846-012-9725-2-Figure1-1.png", + "caption": "Fig. 1 A spincopter", + "texts": [ + " The benefit of such a concept is twofold: firstly, two wings produce more lift (thorough analysis of influence that wing shape has on the lift is given in Section 4), and secondly the stability of the vehicle is further enhanced (some experiments with UAVs based on one rotating wing demonstrate degradation of the stability at low rotational speeds). Two actuators, placed symmetrically with respect to the center of gravity of the vehicle, are used to generate thrust which spins the wings that in effect produce lift. As it is demonstrated in Section 3, difference in thrust, produced by pulsations in rotational speeds of propulsion motors, is used to control the vehicle velocity in the horizontal plane. Concluding remarks and directions for future work are given in the last section of the paper. A spincopter, depicted in Fig. 1, is a flying vehicle comprised of two wings, a fuselage and two propulsors. The fuselage contains an energy source and the required electronics. The propulsors, driven by dc motors, are positioned on the opposite sides of the fuselage, equally distanced from the machine\u2019s rotation axis and attached to a tiny tube. Construction symmetry ensures that flyer\u2019s rotation axis passes through its COG. Each propulsor, composed of two blades, produces thrust that causes torque, which in term spins the vehicle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000535_s11465-011-0128-z-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000535_s11465-011-0128-z-Figure4-1.png", + "caption": "Fig. 4 Location of representative points", + "texts": [ + " To simplify boundary conditions, the sides\u2019 displacements of the arched shell are fixed. Some work must be prepared for Abaqus secondary development. First, the arched shell is created using the composite material shell model. Material properties of the basic layer, damping layer, and constraint layer of the arched shell are steel, damping, and steel, individually. The thickness of each layer in the arched shell is 12, 8, and 1 mm, individually. Then, five representative points are selected for reflecting the dynamic property of the whole gearbox, as shown in Fig. 4. Finally, some necessary settings must be done to output acceleration values of the representative points in the history output manager. This operation can be done in an input file (*.inp). The dynamic response analysis of the arched shell divided into two parts including the nature frequency analysis and random response analysis. To facilitate the secondary development, the restart analysis is directly created in the file random_restart.inp. This file is saved in the folder that includes the finite element model gongban_damp" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000041_cae.20391-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000041_cae.20391-Figure8-1.png", + "caption": "Figure 8 The resultant MTMPS unit with main elements.", + "texts": [ + " Since the plastics part is required drilling operation, the average total time for plastic part is higher than metal average total time. After completion of the processes, parts were stored according to the material type as metal or plastic. Apart from the above-mentioned four main stations, this system has an additional feature. If the process is reset by the user during material handling, transfer cylinder locates the parts in a magazine. Here, this magazine is called as thrash can. The resultant MTMPS unit is shown in Figure 8. Since the MTMPS unit contains electro-pneumatics and electronics equipment, at the design stage the following documents were taken into account: DIN VDE directives and VDMA standard sheets [14]. As described in detail above the MTMPS unit has many moving parts. A whole series of safety measures is necessary to rule out the possibility of dancer to personnel and equipment during the operation of Mechatronics systems. Accordingly safety concerns deserve to be addressed. Regarding the safety concerns every precaution is taken to prevent any accident", + " Cleaning Operation: Third Phase of the Turn Table is cleaning and when the air expelling from nozzle, ON label is displayed and red light defining the processing part position is displayed on the SCADA System. Otherwise OFF label is displayed. Exit: Exit to the main page of the SCADA System. This section deals with the course evaluation which took place in Spring Semester 2008. The evaluation was performed at the end of the Introduction to Mechatronics Course MAK368. This process took 3 weeks time which corresponds to 6 h instruction. The aim of the evaluation was to compare the MTMPS and Conventional MPS and to emphasis the motivation effect on training. The MTMPS (Fig. 8) and a 5 stationed Conventional MPS (shown in Fig. 1, the left picture) were tested by 20 volunteer and enthusiastic students. All the students are supposed to be having almost the same initial background knowledge since they completed the same prerequisite courses successfully. Handouts related to both the MTMPS and Conventional MPS units were provided for students covering all the different scenarios described (emergency stops, fault findings, etc.). The students have been able to implement the programs and make the MTMPS and Conventional MPS work" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003001_access.2018.2883332-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003001_access.2018.2883332-Figure6-1.png", + "caption": "FIGURE 6. The control volume of the relief chamber-trapped volume system (left) and a lumped-parameter representation (right): \u00ac relief chamber, trapped volume 1, \u00ae trapped volume 2, and \u00af kidney ports.", + "texts": [ + " To ensure constant tracking of a single feature within a set when the gear pair is in motion, the following algorithm is used to ensure a minimum vector projects from the previous gravity center to the next of a same feature: \u21c0pci+1 = min k=1,2,...,n \u221a( X kci+1 \u2212 Xci )2 + ( Y kci+1 \u2212 Yci )2 (12) where k denotes the number of features within set B, (Xc,Yc) is the gravity center coordinates of the desired feature, and i denotes the angular position of the gear. An initial condition of (Xc0,Yc0) is provided by user at code initialization. 77512 VOLUME 6, 2018 III. PRESSURE DYNAMICS A diagram of the trapped volume of the gear motor with the proposed relief chamber is shown in Fig. 6. The control volume can be represented by three lumped volume: a relief chamber, two kidney-shaped connection ports, and two trapped volume. The key design parameter is the relief chamber to trapped volume ratio (V1/V2), which dictates how much fluid is available for the purpose of pressure pulsation attenuation. The trapped volume is connected to the relief chamber via a kidney-shaped connection port, the lumped pressure inside the trapped volume is a function of cyclic volumetric compression and expansion of the trapped volume and the flow through the connection port, described as: p\u03071 = \u2212 \u03b2e1 V1 ( V\u03071 + q13 ) (13) where \u03b2e1 is the effective bulk modulus of oil in trapped volume 1, p\u03071 is the time derivative of the pressure of trapped volume 1, V\u03071 is the volumetric time derivative of the trapped volume 1 with respect to time, obtained from previous section, q13 is the volumetric flow rate through the kidney port from trapped volume 1 to the relief chamber, where the outgoing flow is defined positive" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000921_1.3682769-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000921_1.3682769-Figure1-1.png", + "caption": "FIG. 1. (Color online) Schematic diagram of wrinkled pattern characterized by spatial periodicity k and amplitude A. h represents the lateral etching depth.", + "texts": [ + "12 The bilayer thin film was processed by reactive ion etching to produce shallow trenches about 100 nm in depth. Then, the sample was placed in selective etchant to etch away the underlying Si buffer layer and left the SiGe film debonded, forming a free standing film. The debonded film elastically deformed in order to relax the strain energy that was initially stored in the film, and a wrinkled pattern was formed at equilibrium.8 The morphology of the wrinkled pattern was characterized by a spatial period (k\u00bc 3.78 lm) with an amplitude (A\u00bc 206 nm) and a lateral depth (h \u00bc 2 lm), as depicted in Fig. 1. Because the debonded film is spatially deformed, the strain of the wrinkled pattern cannot be treated as the strain of the flat film. Using the nonlinear plate model13 together with the bending effect,14 the strain of a deformed film, in its gen- eral form, is eij \u00bc e0 ij \u00fe 1 2 @ui @xj \u00fe @uj @xi \u00fe 1 2 @n @xj @n @xi 1 2 @2n @xi@xj z, where e0 ij is the initial strain, ui denotes the in-plane lattice displacement in the i direction, and f is the deflection. Denoting the lattice displacements as u\u00f0x; y\u00de; v\u00f0x; y\u00de; and w\u00f0x; y\u00de in the x, y, and z directions, the dominant strain components in the wrinkled film are exx\u00f0x; y; z0\u00de \u00bc e0 xx \u00fe @u\u00f0x; y\u00de @x \u00fe 1 2 @w\u00f0x; y\u00de @x 2 @ 2w\u00f0x; y\u00de @x2 z0; (1) eyy\u00f0x; y; z0\u00de \u00bc e0 yy \u00fe @v\u00f0x; y\u00de @y \u00fe 1 2 @w\u00f0x; y\u00de @y 2 @ 2w\u00f0x; y\u00de @y2 z0; (2) exy\u00f0x; y; z0\u00de \u00bc e0 xy \u00fe 1 2 @u\u00f0x; y\u00de @y \u00fe @v\u00f0x; y\u00de @x \u00fe 1 2 @w\u00f0x; y\u00de @x @w\u00f0x; y\u00de @y @ 2w\u00f0x; y\u00de @x@y z0; (3)a)Author to whom correspondence should be addressed", + " e0 xx, e0 yy, and e0 xy denote the initial strain of the film along various directions. Other strain components are not listed, as their magnitudes are small, considering the fact that the thickness of the film is much smaller than the length and width of the wrinkle.13 It should be noted that, unlike a flat strained film, in addition to the normal strains of exx and eyy, the wrinkled film also exhibits a shear strain of exy as a result of the stretching and bending of the film. Considering a wrinkled film with total length L and the coordinate system depicted in Fig. 1, the lattice displacement functions meet the boundary condition of u\u00f00;y\u00de\u00bcu\u00f0x;0\u00de\u00bc u\u00f0L;y\u00de\u00bc v\u00f00;y\u00de\u00bc v\u00f0x;0\u00de\u00bc v\u00f0L;y\u00de\u00bcw\u00f00;y\u00de \u00bcw\u00f0x;0\u00de\u00bcw\u00f0L;y\u00de\u00bc 0, which which will be used in a later discussion. Using the periodical structure of the wrinkled pattern (x-direction), the lattice displacement function of w\u00f0x; y\u00de can be constructed in terms of two components: Af \u00f0y\u00de and sin\u00f0kx\u00de. The function Af \u00f0y\u00de describes the wrinkled amplitude along the y-direction and can be obtained by fitting to the experimental data" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001418_s11012-010-9354-4-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001418_s11012-010-9354-4-Figure8-1.png", + "caption": "Fig. 8 Pin-disc contact and IR cameras position", + "texts": [ + " The corresponding analytical result could then be plotted in the same figure, revealing a satisfying agreement. Data reduction pointed out Biu = Bid = 0.01, these values being in the expected range. Since the above results seemed encouraging, a more complete experimental scheme, encompassing frictional power loss due to sliding contact, is being realized. A first experimental approach is then being obtained by setting up a testing device specifically designed in order to consider the influence of surface roughness and material properties, based on the typical pin-on-disc pair, see the sketch in Fig. 8. The test rig with a stationary pin specimen pressed against a disc is composed by a steel frame, an electric motor, an hydraulic speed regulator and a pulleybelt transmission for the rotational disc vertical axes. A system based on more sliders allows the pin-holder motion along radial and tangential directions. In this way the mating surfaces may become conformal and different tangential speed can be set for each rotational one. On the steel disc upper surface a weight device allows the axial pin load, while the pin holder can move tangentially suitable for load cell friction force detection" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001057_09507116.2011.581349-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001057_09507116.2011.581349-Figure4-1.png", + "caption": "Figure 4. Automobile chassis.", + "texts": [ + " Between 2008 and 2012, a medium powered European car should consume 25% less, compared with that in 1990. To achieve this objective, vehicle weights must be reduced. Nowadays, large quantities of microalloyed steels, with yield loads around 350MPa, are used for automobile production. Using the new generation steels, known as highstrength steels (HSS) in that they are characterized by yield loads of even 1150MPa, significant reductions in weight can be achieved. HSS rupture and yield loads are much higher than those of traditional steels. For this reason, the thickness of certain components (Figure 4) can be reduced without compromising their ability to support a given load. For example, if a component is made using traditional steel of 2mm thick, using a HSS with adequate mechanical characteristics, it is possible to reduce the thickness to half. The ability to make certain structural components, traditionally made using microalloyed steel, with Figure 5. Lap weld between sheets. Table 2. S355 steel welding combinations. Sample name Upper sheet thickness (mm) Lower sheet thickness (mm) Gap (mm) PR-1 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001841_0278364916679719-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001841_0278364916679719-Figure2-1.png", + "caption": "Fig. 2. (a) Partition of the friction cone Ci into its vertex Oi, punctured boundary Si, and interior Ii. (b) The plane i( \u03c6i) tangent to Si along a ray which starts at xi along the contact force fi = \u03bbiui( \u03c6i).", + "texts": [ + " , fk) \u2208 C1 \u00d7 \u00b7 \u00b7 \u00b7 \u00d7 Ck (8) In order to find critical forces that map to the boundary W , one must first understand the structure of the product set C1 \u00d7 \u00b7 \u00b7 \u00b7 \u00d7 Ck , which forms a stratified set in IR3k . That is, a set that can be decomposed into disjoint cells, each being a manifold without boundary. The cells\u2019 dimension ranges between zero (a single point manifold), and 3k (an open set in the ambient IR3k). In order to describe how C1 \u00d7 \u00b7 \u00b7 \u00b7 \u00d7 Ck decomposes into cells, consider the decomposition of each friction cone into three subsets depicted in Figure 2(a) Ci = Oi \u222a Ii \u222a Si i = 1, . . . , k where Oi = {0} is the cone\u2019s vertex point, Ii is the cone\u2019s interior, and Si is the cone\u2019s boundary surface excluding its vertex point (note that Ii and Si contain only nonzero forces). Using this notation, the set C1 \u00d7 \u00b7 \u00b7 \u00b7 \u00d7 Ck can be decomposed into 3k cells given by C =( O1 \u222a I1 \u222a S1) \u00d7 \u00b7 \u00b7 \u00b7 \u00d7( Ok \u222a Ik \u222a Sk) (9) Each cell in C1 \u00d7 \u00b7 \u00b7 \u00b7 \u00d7 Ck represents a choice of one component from {Oi, Ii, Si} for i = 1, . . . , k. The dimension of any particular cell equals the sum of its component cell dimensions", + " A cell K can be assigned d parameters according to its Si and Ii components as follows. The contact forces lying in the friction cone interior, Ii, are simply parametrized by their Cartesian coordinates: fi \u2208 IR3, where the components of fi are expressed in a reference frame ( si, ti, ni) based at xi. The contact forces on the friction cone boundary, Si, are parametrized by ( \u03bbi, \u03c6i) \u2208 IR+\u00d7IR, where \u03bbi > 0 is the force magnitude and \u03c6i is the force angle, measured by projecting fi on the terrain\u2019s ( si, ti) tangent plane at xi (Figure 2(b)). The forces fi \u2208 Si are thus parametrized by fi( \u03bbi, \u03c6i) = \u03bbiui( \u03c6i) such that ui( \u03c6i) = \u03bci cos( \u03c6i)si +\u03bci sin( \u03c6i)ti + ni (11) Where \u03bci is the coefficient of friction at xi. Since ui( \u03c6i) has a constant magnitude, u\u2032 i( \u03c6i) = \u2212\u03bci sin( \u03c6i)si + \u03bci cos( \u03c6i)ti is orthogonal to ui( \u03c6i). The pair {ui( \u03c6i) , u\u2032 i( \u03c6i) } thus spans the tangent plane to Si at the point fi( \u03bbi, \u03c6i) \u2208 Si, which will be denoted i( \u03c6i) (Figure 2(b)). Note that i( \u03c6i) is the tangent to Si along a ray which starts at xi. Finally, the vector \u03b7i( \u03c6i) = ui( \u03c6i) \u00d7u\u2032 i( \u03c6i) = \u2212\u03bci cos( \u03c6i)si\u2212\u03bci sin( \u03c6i)ti\u2212\u03bc2 i ni is normal to Si at the point fi( \u03bbi, \u03c6i) \u2208 Si, as shown in Figure 2(b). The cells of C1 \u00d7\u00b7 \u00b7 \u00b7\u00d7Ck are next partitioned into cell classes as follows. Definition 7. Each cell class in the set C1 \u00d7 \u00b7 \u00b7 \u00b7 \u00d7 Ck = ( O1 \u222a I1 \u222a S1) \u00d7 \u00b7 \u00b7 \u00b7 \u00d7( Ok \u222a Ik \u222a Sk) is associated with nS choices of Si components, nI choices of Ii components, and k \u2212 nS \u2212 nI choices of Oi components which represent zero contact forces. Each cell class in C1 \u00d7\u00b7 \u00b7 \u00b7\u00d7Ck can be represented by a formal word of k letters from the alphabet {S, I , O}, having nS letters S, nI letters I and k \u2212 nS \u2212 nI letters O" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001301_1.52410-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001301_1.52410-Figure2-1.png", + "caption": "Fig. 2 Nontilted dipole model.", + "texts": [ + " The unit vector along the tether in the orbital frame is given by l\u0302 cos cos i sin cos j sin k (3) To simplify the analysis, only gravity and Lorentz force are treated as the external forces affecting the attitude dynamics of the tethered satellite system. The gravity gradient torque is given by TG 3 r3 Is l\u0302 i l i 3 r3 Is sin cos cos j cos2 sin cos k (4) To calculate the Lorentz torque affecting the EDT system, it is necessary tomodel themagneticfield. In this study, a nontilted dipole model, as shown in Fig. 2, is used to describe the Earth\u2019s magnetic field. Assuming the origin of the geocentric inertia frame is the center of the mass of the Earth, the magnetic field is given by using the orbital elements shown in Fig. 3 as follows: B m r3 2 sin i sin $ cos i sin i cos $ 0 @ 1 A (5) In this paper, it is assumed that the magnetic field is constant along the tether and equal to that of the center of the mass of the tethered system, and that the electric current is a time-varying parameter, but is constant along the tether" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001575_robot.2010.5509764-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001575_robot.2010.5509764-Figure7-1.png", + "caption": "Fig. 7 Front view of shaft: Pulley positions", + "texts": [ + " The energy used for this depends on the mass of the motor unit and its direction relatively to gravity. In a system like a robotic arm, this energy consumption changes with the current position of the whole system. Because first tests show, that this is a value below 5 % of overall energy consumption, it is for now left aside. According to [4] and equation (4.1), the contracting force Fc within the rope becomes shaft ropeshaft c M dd F \u22c5 + = )( 1 (4.3) with dshaft being the diameter of the shaft and drope being the diameter of the rope. In Fig. 7 the different pulley positions for the rope guiding are shown: One double pulley on the left side and two single pulleys on the right side of the shaft. Using the Euler-Eytelwein-Formula [13] for rope friction the force Ftw at the turning wheel is estimated to \u03b1\u03bc \u22c5\u2212\u22c5= 0eFF ctw (4.4) Mechanical tests to estimate the friction constant \u03bc0 delivered an average value for the used pulleys of roundabout 0.1. Within the prototype for one direction there are four pulleys used to translate the force with ~ 90\u00b0 enlacement each, so the enlacement adds up to 360\u00b0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003175_icosc.2018.8587627-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003175_icosc.2018.8587627-Figure1-1.png", + "caption": "Fig. 1: Schematic representation of a quadcopter.", + "texts": [ + " The exchanged power is tftetP . In BG modeling, unified BG elements are used to model any energetic process. They can be classified, according to their energy characterization: active elements (source of effort SE and source of flow SF), passive elements (inertial I, capacitive C, and resistive R), transducers (transformer TF and gyrator GY), and junctions (common effort 0 and common flow 1). The reader can refer to [9] for more details about the BG methodology. 978-1-5386-8537-2/18/$31.00 \u00a92018 IEEE 435 In Fig.1, the schematic representation of quadcopter is shown. The quadcopter have four rotors installed on the ends of a symmetrical cross frame, normally made of carbon fiber. The electronics control system is usually placed in the center. The propellers (rotors) rotate with angular velocities 1 , 2 , 3 , and 4 which generate thrust forces. The pair of propellers on the same branch of the frame rotate in opposite direction of the other pair. This neutralizes the reactive couples and makes the flight possible without turning out of the command" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001530_iros.2010.5648862-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001530_iros.2010.5648862-Figure1-1.png", + "caption": "Fig. 1. Cell manipulation system with optical tweezers.", + "texts": [ + " Second, a visual servo feedback scheme, containing a multi-particle detecting process, is proposed to offer the position information of particles for the flocking control. A. Operation with Holographic Optical Tweezers The optical forces exerted by a few milli-Watts of laser light can dominate the dynamics of nano- and micro-particles. Optical tweezers exploit these forces to precisely manipulate particles or cells without mechanical contact. Trapping of a micro particle is achieved by focusing a laser beam on it, which can be further extended to multiple traps using holography. Fig. 1 shows a robotic manipulation setup that is equipped by a holographic optical tweezers system (BioRyx 200, Arryx) in City University of Hong Kong, which includes an executive module, a sensory module, and a control module. The executive module consists of an X-Y-Z motorized stage (ProScan, Prior Scientific) with microscopic objectives mounted on the Z-axis and a holographic optical trapping device (HOT). Biological samples are contained in a home-built chamber, which is placed on the two-dimensional X-Y motorized stage with a positioning accuracy of 15 nm", + " The well-known particle tracking algorithm [16], [17] is utilized to process the images. The algorithm consists of two main parts: particle location detection and particle tracking. In this paper, we assume that every trapped particle can be tightly trapped by the optical tweezers, and thus the time-consuming particle tracking can be omitted. Fig. 3 illustrates four sub-steps of the particle location detection. Fig. 4 shows an example of the particle detection. The source image is captured by the camera of the optical tweezers system as shown in Fig. 1. Although the source image has much noise, all true particles labeled with small circles are detected successfully, as shown in Fig. 4. Furthermore, the sub-pixel accuracy of detection can be achieved. Based on the flocking control algorithm and the particle detecting algorithm, the flow chart of flocking of micro-particles manipulated by optical tweezers is shown in Fig. 5. There are three main steps in the process. 1) Detecting non-trapped particles. Once the particles are detected, comparison between positions of the detected particles and trapped particles is done to check whether there exist non-trapped particles", + " The trapped particles are then moved to their respective destinations by relocating positions of the optical traps. Finally, positions of all trapped particles in database are updated. 3) Loop checking. To automatically stop the flocking process, a force variable is defined as \u2211 = \u2206\u2206=\u2206 n i i T in 1 1 \u03b5\u03b5\u03b5 (18) where i\u03b5\u2206 is defined in (10). If min\u03b5\u03b5 \u2206<\u2206 , where min\u03b5\u2206 is the minimum force, the flocking of the system is treated to reach a stable state. When min\u03b5\u03b5 \u2206<\u2206 or the maximum loop count is reached, the flocking process stops. Experiments were performed on the established cell manipulation system (Fig.1) to demonstrate the effectiveness of the proposed flocking approach. Silica micro-scale particles with the size of m\u00b52 were selected, and dissolved into water solution. Particle images were captured using a CCD camera (640\u00d7480 pixels) under 40X microscope. For the BioRyx 200 optical tweezers, positions of any optical traps must not exceed the diffractive element device (DED) coordinate range (512 \u00d7 512). Therefore, a calibration is needed to map a 640\u00d7480 particle image to the diffractive element device by using following equations \u00d7= \u00d7= 512 480 512 640 I d I d yy xx (18) where [ ]Tdd yx , and [ ]TII yx , denote the positions with respect to the DED frame and image frame, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000546_j.1460-2695.2010.01540.x-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000546_j.1460-2695.2010.01540.x-Figure8-1.png", + "caption": "Fig. 8 Crack propagation point radial and angular coordinates.", + "texts": [ + " The prediction model computational efficiency becomes even more important when considering the required BE model preparation. In order to refine the prediction model, first-order interactions between two-factor combinations are analyzed in this section. Parameter cross-influences are studied by means of the crack tip location at a specific stage of the propagation. Relative radial (r/hr) and angular (\u03b8/ \u03b8 t) c\u00a9 2011 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct 34, 470\u2013486 crack tip positions are introduced instead of the x and y coordinates (Fig. 8). For a fixed relative crack length of a/m = 0.355, Fig. 9 shows the two-factor interaction plots: the upper left plot group concerns the r/hr crack tip coordinate, whereas the lower right plot group refers to the \u03b8/ \u03b8 t coordinate. Abscissa factors are identified at the top and bottom of Fig. 9, while factors corresponding to each three-level curve are specified on the sides. Figure 9 reveals unequal influences of the factor. The plots demonstrate that factor Rv has a low degree of influence on the crack location, while significant interactions between factors \u03b10, hr and np appear" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001093_icra.2012.6225142-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001093_icra.2012.6225142-Figure6-1.png", + "caption": "Fig. 6. Prototype Module with the Omnidirectional gear Concave Version", + "texts": [ + " We made three types of gear mechanisms, each of which is described from the next section. In this section, we describe the configuration of the planar version of the omnidirectional driving gear mechanism. The actuators that drive the mechanism are placed in world coordinates so the plate that moves omnidirectionally on the flat plane can be configured in a lightweight manner. The overall view of the manufactured prototype of the convex version is shown in Fig. 5. The driving unit of the prototype of the concave version is shown in Fig. 6. To activate this prototype of concave version of the omnidirectional gear driving unit, two pairs of spur gears are deployed inside of the tube structure of the concave omnidirectional gear. One pair is deployed to its rotational direction, and the other is deployed to its translational direction. Only one spur gear is enough to activate this omnidirectional gear, however, to cancel any undesired outer torque, a pair of spur gears is deployed to move its one axis. IV. CONCLUSIONS In this paper, the principles and the actual configurations of the omnidirectional driving gear mechanism that was implemented by the authors were described" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000028_s12206-009-1166-x-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000028_s12206-009-1166-x-Figure3-1.png", + "caption": "Fig. 3. Vector geometry of the parallel-link section.", + "texts": [ + " The position of the TCP along the i-j plane is changed by the parallel-link arm section. Also, the position of the TCP along the k-axis is changed by the serial-link arm section. The pose around the yawing direction and pitching direction is changed by the serial-link arm section. The positions and pose of the TCP are controlled individually in each section. Therefore, to calculate the kinematic analysis, we divided this hybrid robot into each section. The geometrical model of each section is defined in Fig. 3 and Fig. 4 individually. Especially, the pose of the constant vector C8 depends on the length of the linear motion vector L3 and L4. The position of the TCP on the i-j plane is changed by the pose of the constant vector C8. So, the constant vector C8 is included in the parallel-link arm section. At the serial-link arm section the origin point is defined as Os. The forward kinematic analysis calculates the position and pose of the TCP for this hybrid robot [3]. At this analysis, each length of the motion vectors L3,L4 and L7 is defined as l3,l4 and l7. At the parallel-link arm section, the pose around kaxis is changed by the difference in length of l3 and l4. This pose is defined as the rotation \u03b81. At the serial-link arm section, the pose is changed by the rotation angle \u03b84 and \u03b85. So, the pose of TCP is defined as Eq. (1) [3]. 51 4k k \u03c9\u03b8\u03b8 \u03b8=E E E E I (1) In this equation, the E is the pose matrix; the vector \u201cI\u201d shows the initial position of the TCP. The position of the TCP on the projected i-j plane, based on Fig. 3, can be calculated as Eq. (2). 2 6 6 8 1 4 6 6 8 1 cos sin sin cos c p cx l p cy \u03b8 \u03b8 \u03b8 \u03b8 + +\u239b \u239e\u239b \u239e = \u239c \u239f\u239c \u239f + +\u239d \u23a0 \u239d \u23a0 (2) In this equation, c2 shows the length of the constant vector C2; p6 shows the length of the polar P6; c8 shows the length of the constant vector C8. And, the position along the k-axis, based on Fig. 4 can be calculated as Eq. (3). 7 9z l g= + (3) In this equation, g9 shows the length of the constant vector G9. Also, the position at the intersection of the polar vector P5 and P6 moves on the center-line as shown in Fig. 3. Therefore, \u03b81, \u03b86 and p6 are shown as Eq. (4). 3 4 1 2 6 1 2 6 1 2 2 cos l l c cp \u03b8 \u03c0\u03b8 \u03b8 \u03b8 \u2212 = = \u2212 = (4) The inverse kinematic analysis calculates the length of each link and rotation angle from the position and pose of the TCP [3]. At first, to calculate the rotation angle \u03b85 and \u03b81+\u03b84, the pose of TCP is defined as Eq. (5). 51 4k k \u03c9\u03b8\u03b8 \u03b8=E E E E I (5) To calculate the length of l7, the position of the TCP which is along the k-axis, as shown in Fig. 4, is defined as Eq. (6). The length of l7, based on Eq. (6), can be calculated as Eq. (7). = +k k 7 9P L G (6) 7 9l z g= \u2212 (7) The polar vector P shows the position of the TCP which is along the k-axis. The position of the TCP on the projected i-j plane, as shown is Fig. 3, is defined as Eq. (8). And the length of l4, based on Eq. (7), can be calculated as Eq. (9). 0 2 4 6 8= + + +P C L P C (8) 2 8 1 0 0 4 0 0 8 1 6 sin sincos cos tan c c pl p c \u03b8 \u03b8\u03b8 \u03b8 \u03b8 + \u2212 = \u2212 + (9) The polar vector P0 shows the position of the TCP on the i-j plane; the p0 shows the length of the polar vector P0; \u03b80 shows the angle between the polar vector P0 and j-axis; \u03b84 shows the angle between the linear motion vector L4 and j-axis. Also, the point at the intersection of the polar vector P5 and P6 is moved on the center-line as shown in Fig. 3. Therefore, p0 and \u03b81 are shown as Eq. (10). And, \u03b80 is shown as Eq. (11). 2 2 0 1 0 tan p x y x y \u03b8 \u2212 = + = (10) 1 8 1 2 8 tan 1 y c y c \u03b8 \u2212= \u239b \u239e \u2212 \u239c \u239f \u239d \u23a0 (11) The length of l3, as shown in Fig. 3 is defined as Eq. (12). 3 4 2 12 tanl l c \u03b8= + (12) The rotation angle \u03b84 around the k-axis, based on Eq. (5) and Eq. (11), can be calculated. In general, the robot for de-burring needs the performance that can be moved along an arbitrary continuous spatial locus. Therefore, we investigated the continuous path control algorithm and the pose control algorithm. In this part, the continuous path control method is shown. The equation of the position of TCP on the i-j plane is defined as the time variable function" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000793_17452759.2010.527010-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000793_17452759.2010.527010-Figure1-1.png", + "caption": "Figure 1. Recursive approaches for adaptive slicing (Mani et al. 1999, Sabourin et al. 1997).", + "texts": [ + " Other researchers use slightly different approaches. Sreeram and Dutta (1994) sort all vertices of the polygonal mesh according to their z value (build direction). The resulting segmentation enables the segment wise determination of optimum layer thicknesses by means of maximum allowable cusp height. A similar approach to accelerate the layered manufacturing process is presented by Mani et al. (1999) who slice a part into regions of different quality requirements (region-based adaptive slicing; see Figure 1) on the basis of user-defined specifications. The part regions are analysed separately so that optimum layer thickness can be determined for each region. Whereas the previous mentioned approaches assume that layer profile is rectangular, Pandey et al. (2003) use a parabolic profile to improve part accuracy. A tolerance for surface roughness Ra serves as error characteristic. This concept has been implemented and tested on a FDM system but is not suitable for other LM applications. Byun and Lee (2006) made similar developments to define local layer thicknesses with respect to surface roughness", + " (1996) picked up the cusp height concept for recursive refinement of part slices which have been generated with maximum allowable layer thickness. In an iterative process each single layer is examined with regard to a possible refinement in compliance with predefined maximum values for cusp height. In order to accurately take into account the changes of curvature, adjacent layers are included in geometrical examination. Further developments of the authors led to an approach where contours of the cross sections are generated with an offset inside the part (accurate exterior, fast interior, see Figure 1). In this manner, only outer regions will be refined whereas inner regions can D ow nl oa de d by [ D al ho us ie U ni ve rs ity ] at 0 6: 43 3 1 D ec em be r 20 14 be manufactured with thick slices (Sabourin et al. 1997). In doing so, surface quality can be sustained by using thin layers near part boundaries, while thick layers inside the part allow the acceleration of the part building process. Furthermore, Tyberg and Bohn (1998) presented a method for local adaptive slicing where only those layers are refined which lead to a reduction of build time and increase surface quality at the same time" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002124_s11012-017-0737-7-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002124_s11012-017-0737-7-Figure2-1.png", + "caption": "Fig. 2 FE model of the meshing spur gears (a) and the effect of tooth morphing (b)", + "texts": [ + " In order to find improvements in their mechanical performance, micro-geometry modifications of tooth profiles were applied to the model, using the so-called morphing technique, described in detail in [43, 44]. The performed RS-based robust optimization enabled finding of Pareto-optimal designs, improved in terms of the observed characteristics. 5.1 Gear FE model and simulation methodology The optimization methodology was implemented in order to find improvements in a pair of spur gears, for which the geometrical parameters are listed in Table 1. The FE representation of the gears is shown in Fig. 2a. The model was prepared using three-dimensional hexahedral mesh. Element size for the area in which teeth contact was foreseenwas equal to 0.11mm, measured along the tooth profile. The total number of elements was equal to 148019 and 526321 active DoF. As described in detail in [21], evaluation of each modification was done by performing a series of nonlinear static analyses of meshing gears being rotated by a fixed angle D/ \u00bc 0:21 . Constraints and loads were applied in the gear rotation center points, as shown in Fig. 2a. Based on these discrete values calculated for each gear position, STE, bending stress and contact stress curves were estimated for every tested solution. For a deeper description of the simulation method, please refer to [21]. 5.2 Definition of the objectives The objectives of the described optimization are defined as minimization of the peak-to-peak value of the STE (PP STE) and minimization of the maximum value of CS, calculated in a whole meshing cycle for a candidate modified solution. The former is a measure of internal gear vibration source and, as shown in [38, 39], is strongly correlated with the DTE", + " Tooth microgeometry modification has been proved as an efficient way to significantly improve static and dynamic gear performance [15, 32\u201335]. Despite the fact that most frequently this type of teeth modification is applied to tip and root simultaneously, researchers report a similar effect when an equivalent (i.e., being a sum of tip and root reliefs) modification is imposed only to the tip [36, 37]. Similarly, in the study described in this paper, only tip micro-geometry modification was performed. These alterations were imposed to the FE model by the so-called mesh morphing technique, as shown in Fig. 2b. The allowable modification length and amplitude was limited and controlled by two parameters, i.e. starting radius rt and depth d, for which the following ranges were defined: rt 2 \\74:10mm; 77:25mm[ and, for the depth parameter: d 2 \\0:00mm; 0:04mm[ . In contrast to the procedure described in [21], the optimization strategy presented in this paper assumes application of both linear and parabolic gear modifications, which have a non-negligible influence on the observed quantities. To control this, a discrete parameter indicating the type of imposed modification was included into the set of optimization variables" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000041_cae.20391-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000041_cae.20391-Figure4-1.png", + "caption": "Figure 4 A 3D model of the proposed MTMPS unit.", + "texts": [ + " Initial conditions required to start the MTMPS unit can be described as follows: the system pressure is adequate (B12 sensor), transfer cylinder is at home position (S33 and B2 sensors), transfer cylinder is up position (B4 sensor), and new part at part magazine (B6 and/or 7 sensors) and cycle is not active. For the required processes described above, standard mechanical, electrical/electronical, and pneumatical elements were determined and a 3D model of the proposed system was developed as shown in Figure 4. A FPC 101 type Festo PLC, which has 24 digital inputs and 16 digital outputs, was chosen to control MTMPS by considering the inputs and outputs of the system [13]. Standard mechanical, pneumatical, and electrical/ electronical components were assembled on the MTMPS platform. PLC application was performed to accomplish the desired processes by using a Festo FST 100 interface. A Festo CPV18 Valve Terminal was used to carry out the controls of the pneumatic actuators on the unit [11]. Finally, SCADA application on the MTMPS was realized" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003252_robio.2018.8665152-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003252_robio.2018.8665152-Figure1-1.png", + "caption": "Figure 1 The structure of two-segment continuum manipulator. Each cross shaft consists of two mutually perpendicular joints.", + "texts": [ + " The simulation results are shown in section 4. Finally, the conclusions are is drawn in Section 5. 978-1-7281-0377-8/18/$31.00 \u00a9 2018 IEEE 1777 The PCC continuum manipulator that we consider in this paper is cable-driven and contains multiple segments, each of which consists of 6 subsegments. Each subsegment is composed of a cross shaft and its subsequent link. Each cross shaft consists of two mutually perpendicular revolute joints. We illustrate two-segment manipulator, it contains 12 cross shafts and has 24-DoFs. Fig.1 shows the structure of the two-segment PPC continuum manipulator, , {1,2,...,12}, {1,2}ij i j represents the joint angle of the corresponding joint, i means the label of cross shaft and j means the label of joint number. The cable-driven system is that motors pull the driving cables to generate tension in the cables to drive the arm to bend. There are two kinds of mapping relationship. One is from the motor position and speed to the tension in the cables, and the other is from the tension in the driving cables to torque or speed on the joint of the robot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001389_j.ijthermalsci.2010.11.001-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001389_j.ijthermalsci.2010.11.001-Figure1-1.png", + "caption": "Fig. 1. Braking tribometer: general view (a), pi", + "texts": [ + "7% of the total surface of the pin) in order to minimize its influence on the nature of the pinedisc contact. Local temperatures are evaluated at the pin-disc interface thanks to the optical fiber of the two-color pyrometer inserted in the pin. In addition, disc temperature at the exit from the contact area is estimated by a monochromatic pyrometer. Tests were performed on a pin-on-disc braking tribometer developed to reproduce at reduced-scale the contact stress observed between disc and pad at full-scale during railway braking [20]. Fig. 1 shows the braking tribometer: general view (a), pin and disc view (b) and schematic drawing (c). A motor rotates the disc with an initial Nomenclature a thermal diffusivity, m2 s 1 A amplification constant, V W 1 m3 sr b thermal effusivity, J m 2 K 1 s 1/2 C1 first Planck constant, W m2 sr 1 C2 second Planck constant, m K C Torque, N m h heat convection coefficient, W m 2 K 1 Jm Bessel function of the first kind of order m k thermal conductivity, W m 1 K 1 K uncertainty coefficient L luminance, W m 3 sr 1 p pressure, Pa q heat flux density, W m 2 r radius, m R mean friction radius, m S detector signal, V t time, s T temperature, K z cylindrical coordinate, m Greek symbols a heat partition coefficient bn roots of transcendental equation 3 emissivity l wavelength, mm q angle, rad s relative value of the detector signal s transmittivity u rotational speed, rpm Subscripts l1 detector 1 of the two-color pyrometer l2 detector 2 of the two-color pyrometer bichro fiber optic two-color pyrometer d disc p pin model analytical solution pyro Impac monochromatic pyrometer sliding velocity comprised between 0 and 15 m s 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003864_s11012-019-01081-5-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003864_s11012-019-01081-5-Figure3-1.png", + "caption": "Fig. 3 Model of the 3-CPCR manipulator: a parameters of the kinematic chain; b resulting manipulator", + "texts": [ + " The loop equation can be written as: rP \u00bc rA \u00fe a\u00fe b\u00fe c\u00fe d \u00fe e\u00fe f \u00fe rGP \u00f01\u00de By expressing each vector associated with a variable-length link as the product of its module (length) times the unit vector of the corresponding P joints, the following expression is obtained: rP \u00bc rA \u00fe a ua \u00fe b ub \u00fe c uc \u00fe d ud \u00fe e ue \u00fe f uf \u00fe rGP \u00f02\u00de Now, by taking the derivative with respect to time, the velocity equation is obtained: vP \u00bc _a ua \u00fe _b ub \u00fe _c uc \u00fe _d ud \u00fe _e ue \u00fe _f uf \u00fe xb b\u00fe xc c\u00fe xd d \u00fe xe e\u00fe xf f \u00fe x rGP \u00f03\u00de where vP is the linear velocity of the coupler point P. In Eq. (3), the angular velocity of each element and that of the moving platform are given by: xb \u00bc _a ua xc \u00bc xb \u00fe _b ub xd \u00bc xc \u00fe _c uc xe \u00bc xd \u00fe _u uu xf \u00bc xe \u00fe _h uh x \u00bc xf \u00fe _w uw \u00f04\u00de 3.2 From the generic chain to the 3-CPCR parallel manipulator The CPCR chain, shown in Fig. 3a, is easily derived by simply particularizing the following parameters of the generic chain depicted in Fig. 2: d \u00bc constant e \u00bc f \u00bc 0 u \u00bc h \u00bc w \u00bc 0 \u00f05\u00de Once the parameters of the CPCR chain have been determined, the whole manipulator can be modeled as shown in Fig. 3b by taking into account that the two first P and R joints form a C joint and that similarly the third P and the second R joints form another C joint. Based on the kinematic diagram in Fig. 3a, the loop equation is: rP \u00bc rA \u00fe a\u00fe b\u00fe c\u00fe d \u00fe rEP \u00f06\u00de Then, by differentiating with respect to time, the velocity equation becomes: vP \u00bc _a ua \u00fe _b ub \u00fe _c uc \u00fe xb b\u00fe xc c\u00fe xd d \u00fe x rEP \u00f07\u00de where xb \u00bc _a ua xc \u00bc xb xd \u00bc xc \u00fe _b ub x \u00bc xd \u00fe _c uc \u00f08\u00de Subsequently, the components of the unit vectors associated with the translational motion ua, ub, uc and the rotational motion ua, ub, uc are obtained. For this purpose, the mathematical relations established in [21] are used. First, the unit vectors of chain 1 are obtained", + " 1 and 3a, the first vector is: ua1 \u00bc ua1 \u00bc 0; 0; 1\u00f0 \u00de \u00f09\u00de Because ub1 is perpendicular to ua1 : ub1 \u00bc cosa1; sina1; 0\u00f0 \u00de \u00f010\u00de The vectors uc1 and ub1 are parallel to each other and perpendicular to ub1 and ua1 . Applying the cross product yields: uc1 \u00bc ub1 \u00bc ub1 ua1 \u00bc sina1; cosa1; 0\u00f0 \u00de \u00f011\u00de Next, the components of the unit vectors ud1 and uc1 are obtained. Both are parallel to each other and perpendicular to uc1 . To obtain their components, an additional equation is needed. In this case, the simplest approach is to carry out two consecutive rotations: first, rotating a1 around Z, and next, b1 around uc1 . The result is: ud1 \u00bc uc1 \u00bc cosa1sinb1; sina1sinb1; cosb1\u00f0 \u00de \u00f012\u00de According to Fig. 3b, chain 2 is obtained by first rotating chain 1 by 180 around the Z-axis in the positive direction and next rotating it by 90 around the X-axis in the negative direction. On the other hand, chain 3 is obtained by rotating chain 2 by 90 around the Z-axis in the positive direction. By applying the corresponding homogeneous transformation matrices, the unit vectors of chains 2 and 3 are given by: ua2 \u00bc ua2 \u00bc 0; 1; 0\u00f0 \u00de \u00f013\u00de ub2 \u00bc cosa2; 0; sina2\u00f0 \u00de \u00f014\u00de uc2 \u00bc ub2 \u00bc sina2; 0; cosa2\u00f0 \u00de \u00f015\u00de ud2 \u00bc uc2 \u00bc cosa2sinb2; cosb2; sina2sinb2\u00f0 \u00de \u00f016\u00de ua3 \u00bc ua3 \u00bc 1; 0; 0\u00f0 \u00de \u00f017\u00de ub3 \u00bc 0; cosa3; sina3\u00f0 \u00de \u00f018\u00de uc3 \u00bc ub3 \u00bc 0; sina3; cosa3\u00f0 \u00de \u00f019\u00de ud3 \u00bc uc3 \u00bc cosb3; cosa3sinb3; sina3sinb3\u00f0 \u00de \u00f020\u00de As previously stated, the 3-CPCR manipulator has 6- DOF, the output variables of the velocity problem being the components of vP and x", + " (25) yields: 0 0 1 \u00f0rE1P\u00dey \u00f0rE1P\u00dex 0 0 1 0 \u00f0rE2P\u00dez 0 \u00f0rE2P\u00dex 1 0 0 0 \u00f0rE3P\u00dez \u00f0rE3P\u00dey 0 0 0 sina1 cosa1 0 0 0 0 sina2 0 cosa2 0 0 0 0 sina3 cosa3 2 6666664 3 7777775 vP x \u00bc 1 0 0 d1sinb1 0 0 0 1 0 0 d2sinb2 0 0 0 1 0 0 d3sinb3 0 0 0 1 0 0 0 0 0 0 1 0 0 0 0 0 0 1 2 6666664 3 7777775 _a1 _a2 _a3 _b1 _b2 _b3 2 6666664 3 7777775 \u00f026\u00de 4.2 Relation among input variables _b and _a Next, the equations that relate the inputs _a and _b will be obtained. For this purpose, the relation that the moving platform imposes among the vectors uc1 ,uc2 , and uc3 will be considered. As can be observed in Fig. 3a, b, the vectors uc2 and uc3 are parallel to each other and perpendicular to uc1 . Thus: uc2 \u00bc uc3 ; uc1 uc2 \u00bc 0 \u00f027\u00de Substituting the vectors\u2019 components into Eq. (23), differentiating with respect to time, and solving gives the following result: _b1 \u00bc\u00f0 _a1 cosa1=tana3 sina1=tana2\u00f0 \u00de _a2 cosa1 1\u00fe tan2a2 = tan2a2 _a3 sina1 1\u00fe tan2a3 = tan2a3\u00de= 1\u00fe tan2b1 cosa1=tana2 \u00fe sina1=tana3\u00f0 \u00de2 \u00f028\u00de _b2 \u00bc _a2 tana3=tana2 \u00fe _a3 1\u00fe tan2a3 = \u00f0\u00f01\u00fe tan2b2\u00desina2\u00de \u00f029\u00de _b3 \u00bc _a2 1\u00fe tan2a2 \u00fe _a3 tana2=tana3 = \u00f0\u00f01\u00fe tan2b3\u00desina3\u00de \u00f030\u00de 4", + " The three components of vP depend on the platform\u2019s position and the input velocity _a1. In this case, the moving platform rotates around a fixedorientation axis that translates in space. This axis is parallel to the unit vectors uc2 and uc3 . Particularizing the previous case by substituting a2 \u00bc a3 \u00bc p=2 into Eq. (35), the axis of rotation is parallel to the Z-axis, and there is no displacement of the moving platform in the direction of the axis of rotation, but only a translation of the axis of rotation in the XY-plane (Fig. 3): _a1 \u00bc _a2 \u00bc _a3 \u00bc 0; _a2 \u00bc _a3 \u00bc 0; a2 \u00bc a3 \u00bc p=2 ) xx \u00bc 0;xy \u00bc 0;xz \u00bc _a1 vPz \u00bc 0 \u00f036\u00de In the case that all translational and rotational inputs are cancelled except for _a2, the moving platform rotates around an axis with a variable orientation that translates in space. Moreover, if a1 \u00bc p=2: _a1 \u00bc _a2 \u00bc _a3 \u00bc 0; _a1 \u00bc _a3 \u00bc 0; a1 \u00bc p=2 ) xx \u00bc 0;xy \u00bc xtana3 ;xz \u00bc x \u00f037\u00de where x \u00bc sina3cosa3= cos2a3cos2a2 1\u00f0 \u00de\u00f0 \u00de _a2. In this case, the moving platform rotates around a fixed-orientation axis that translates in space", + " This is the case with the 3-CPPU parallel manipulator represented in Fig. 6(a), where the italicized PU refers to the parallelism between the uc translational vector and the ub rotational vector (the first revolute pair of the U joint), as shown in Fig. 6(b). Moreover, the revolute pairs that connect chains 2 and 3 with the moving platform are located in such a way that uc2 and uc3 are parallel to each other and perpendicular to uc1 [Eq. (27)]. Figure 6 shows a 3-CPPUmanipulator in which the particularization based on the generic chain is also that of Fig. 3b. Its system of velocity problem equations is given by Eq. (31), but substituting d \u00bc 0 in Eq. (5) because points D and E are coincident. All the conclusions derived from the multioperationality analysis of the 3-CPCR described in Sect. 6 are applicable to the 3-CPPU manipulator. With the aim of validating all these theoretical results and verifying that the sets of group displacements established in Sect. 5.2 can be obtained in practice, a prototype of the 3-CPCRmanipulator has been built, and is shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000017_09544062jmes1943-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000017_09544062jmes1943-Figure4-1.png", + "caption": "Fig. 4 Stress\u2013strain diagram and corresponding information describing the pullout process", + "texts": [ + " Mechanical Engineering Science Tensile, flexural, and compression tests were conducted using aWP300 PC Aided Universal MaterialTest machine at room temperature (28 \u00b1 2 \u25e6C). Meanwhile, a hardness tester (TH210 Shore D durometer) was used to determine the Shore D hardness of the composite. The hardness was measured perpendicular to the fibre orientation of the polyester composite. As a result of repeating all the experiments three times, the standard deviation of the measurements was determined and is listed in Table 1. A sample of the single fibre pullout result is displayed in Fig. 4(a) that shows the stress strain/diagram. In general, as the strain increases, the applied force increases. This indicates that there is no pullout process during the test. In the case of the pullout, there should be a dramatic decrease in the force when the strain increases. This has been reported when the oil palm fibre was tested [17, 19]. For the current work, treated betelnut fibres showed good interfacial adhesion characteristic with the matrix. In other words, the fibres carried the load and transferred it to the matrix without slipping, see Fig. 4(b). The damage mechanism that took place during the pullout test was fibre breakage (Fig. 4(b)). The presence of trichomes (Figs 1(b) and 4(c)) on the surface of the fibres locked the fibres in the matrix (i.e. generated high interfacial adhesion with the matrix). From Fig. 4(a), it can be determined that the maximum stress during the pullout test reached up to 280 MPa, which approximately equals the tensile strength of the fibres [25]. Stress/strain diagram of neat polyester (NP) and its composite (T-BFRP) is shown in Figs 5(a) and (b), respectively. In general, both materials show brittle behaviour. No plastic zone can be observed. From this figure, it can be seen that reinforcing the polyester with the betelnut fibre improves its tensile strength and strain. The improvement in the tensile strength and strain is about 75 per cent and 72 per cent, respectively", + " In other words, betelnut fibre has a high potential to replace the glass fibres for mechanical applications. Micrographs of the tensile samples after tests at the breakage surface are presented in Figs 8(a)\u2013(d). The general view on the surfaces shows two different damages in the fibrous regions, where some of the fibres were pulled out or detached, while the rest were broken. For the current work, the single fibre pullout tests showed that treated betelnut fibre has high interfacial adhesion characteristics (see Fig. 4(b)). However, the tensile samples showed that some of the fibres JMES1943 Proc. IMechE Vol. 224 Part C: J. Mechanical Engineering Science were pulled out or detached. In the magnified images, Figs 8(b)\u2013(d), it is clearly shown that the trichomes adhered on the polyester matrix. This indicates very high interfacial adhesion of betelnut fibres with the polyester matrix. However, in some situations, it seems that the inner fine fibres pulled out, which could explain the current phenomenon observed in the T-BFRP composite" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003803_s10846-019-01103-0-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003803_s10846-019-01103-0-Figure6-1.png", + "caption": "Fig. 6 Graph structure: red edges denote the trajectory edges, with measurement wP ij potentially merged from wKF ij and wLL ij , all \u2208 SE(3), cyan edges encode the star tracker-based constraints wR 0j \u2208 SO(3), and green edges represent constraints based on scan matching employed during loop closing, i.e., wLL jk \u2208 SE(3)", + "texts": [ + "2, which draws on the trajectory estimate and the respective measurements in a mapping with known pose scheme [91, Chap. 9]. The joint estimation of a trajectory and a map may be formulated as p(x0:t , Y |z0:t , u1:t ) , (1) where measurements for the single timepoints are denoted as z0:t , while u1:t represent the respective control measurements. Graph SLAM approaches can be roughly divided into a backend (see Section 4.1), where the optimization takes place and a frontend (see Section 4.2), where the structure of the graph G is constructed (see Fig. 6). The construction itself is conducted by inserting nodes v, each representing a pose estimate, into G and constrain them with edges w, which represent offset measurements. A star tracker\u2013based offset measurement wR 0j (see Section 3.4.2) constraints the attitude with respect to the navigation frame (see Section 4.2.1). In the graph, the latter is represented in terms of a node v0 with aligned pose. In addition, it is common to employ some kind of recursive filtering technique to foster the convergence speed of the optimizer backend in Graph SLAM approaches", + " The latter is used to validate a scan matching attempt wLL ij between point clouds obtained at the nodes vi and vj , respectively (see Section 4.2.4). Thus, we account for the scenario-specific observation that a successful scan matching attempt is more precise than the filter result, while a failed attempt may result in arbitrarily bad results. If validation succeeds, the scan matching result wLL ij is merged with the relative filter- estimate wKF ij , resulting in a pose constraint wP ij between vi and vj (see Fig. 6a). If validation fails, wKF ij is directly used as the constraint wP ij and scan matching is re-attempted in the next time step. After the node vj has been inserted into the graph, the EKF is reset by setting the position part r of the state vector to zero and the attitude part R to identity. In addition, the position and attitude related parts of the covariance are set to zero as well. Thus, the latest node vj effectively becomes the new reference pose x0 of the filter. Accordingly, the backend is provided with a good initial guess by the filter and in return provides the filter operating in the frontend with an improved reference pose x0 smoothed over all preceding measurements", + " Finally, if a loop was closed or the optimization has not been carried out since some \u03b5optimze, the optimization backend is called (see Section 4.1 and cf. lines 26, 32 and 33). The measurement zR t \u2208 SO(3) is taken by the star tracker and is represented as a DCM (see Section 3.4.2). It is used in the filtering as well as in the smoothing step. In the former, it serves as a correctional measurement in the Kalman filter (see Section 4.2.3), while in the latter it serves as a differential measurement wR 0j for SO(3) edges constraining the attitude part of two vertices defined in SE(3) (see the cyan edges in Fig. 6a). However, as the EKF estimate is treated as a relative transformation it does not reflect the absolute nature of the star tracker measurements. For that reason and due to the fact that only a small portion of the star tracker measurements are actually added to the graph, the negative effects of considering them twice can be neglected. For graph construction, we initially insert a fixed node v0, which can not be altered in the optimization process and whose pose is aligned with the navigation frame", + ") EKF estimate 20 Gt \u2190 add node(Gt\u22121, j, \u03bc P t ) ; // insert node vj with pose \u03bcP t 21 Gt \u2190 add edge ( Gt , 0, j, zR t , ( QR t )\u22121 ) ; // insert attitude edge wR 0j 22 Gt \u2190 add edge(Gt , i, j, \u03bc P t , P t ) ; // insert pose edge wP ij 23 rt \u2190 [0, 0, 0] , Rt \u2190 I3\u00d73 ; // reset \u03bcKF t relative to vj 24 KF t \u2190 [ 06\u00d76 03\u00d76 06\u00d73 KF t,7:9,7:9 ] ; // reset KF t relative to vj 25 if scan validated then 26 Gt , loop closed \u2190 validateNearest(Gt , vj ); // attempt loop closing 27 end 28 else 29 loop closed \u2190 False; 30 end 31 if last optimization > \u03b5optimize \u2228 loop closed then 32 G\u2217 t \u2190 optimize(Gt ); // call optimization in backend 33 last optimization \u2190 t ; 36 #steps \u2190 #steps + 1; 37 return G\u2217 t ; // updated and potentially optimized graph into the vector space R3 and thus enables the local linearization of the error in the optimizer backend (see Section 4.1). The pose edge wP ij \u2208 SE(3) fully constrains the relation of the vertices vi, vj \u2208 SE(3). (see red edge in Fig. 6a). The error function e : S \u2032 \u00d7 S \u2032 \u00d7 S \u2032 \u2192 R 6 for poses is thus defined as e(vi, vj , w P ij ) = wP ij \u2212 w\u0304P ij (vi, vj ) = ([ Rvj rvj 0 1 ]\u22121 [ Rvi rvi 0 1 ]) S \u2032 \u03bcP t , (23) where S \u2032 denotes the compound S -operator defined on the manifold but without the velocity part of the state (cf. Section 3.1). Again, to enable a local linearization of the error, it is mapped to vector space R6 by this operator. Rv\u25e6 and rv\u25e6 denote the attitude and position part of the i-th and j -th node, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001020_1.3652862-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001020_1.3652862-Figure7-1.png", + "caption": "FIG. 7. Finite element geometry of a Tourte-tapered violin bow stick with attached head and frog having effective lever heights from the neutral axis of 18 mm at the head of the bow and 24 mm at the frog.", + "texts": [ + " 6, December 2011 Colin Gough: Violin bow flexibility 4111 diameters and slightly shorter lengths of viola and cello bows, though the general form of the taper remains much the same. Vuillaume derived a very slightly larger thicker section for the bows he measured illustrated by the solid line. The main difference between violin, viola and cello bows is their increasing diameter, hence weight. The Vuillaume expression for the tapered Tourte bow will now be used to describe the taper of our finite-element bow model illustrated schematically in Fig. 7. A uniform elastic constant of E\u00bc 22 GPa along the grain is assumed. The bow stick was divided into 14, equal-length, conical, subdomains describing the tapered length of the stick with a crudely modeled rigid bow head and frog at its ends. The bow hair was simply represented by the hair tension between its points of attachment on the underside of the frog and head of the bow. Because the structure is relatively simple, at least in comparison with the violin, a 3-dimensional analysis of the bending modes was used" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000110_ac60286a023-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000110_ac60286a023-Figure1-1.png", + "caption": "Figure 1. Electrode assembly", + "texts": [ + " i( I\" i SALT BRIDGE ,TYGON SLEEVE 1 (*LASS FRIT *ELECTRODE CONTACT TEFLON SPACER GLAssgb N SLEEVE Harvard Apparatus Model 600-1200 peristaltic pump. By placing the pump downstream from the electrode, the solution can be pulled from the sample through a short piece of glass capillary tubing, minimizing the entrance of oxygen into the solution. The tubular carbon electrode ( W E ) is constructed by press-fitting a tubular piece of wax-impregnated spectroscopic-grade graphite into a specially designed holder (Figure 1). Several factors were considered in arriving at this cell design. The TCE is constructed so that it can be assembled easily and can be quickly replaced if the electrode surface becomes damaged or fouled. In addition it is important to obtain a smooth bore to and through the electrode so that a parabolic velocity profile with laminar flow is established before the solution enters the electrode which would not be disturbed by turbulence as the solution enters the electrode in order that derived equations be valid", + " This paper describes a n extension of the above two systems. The new system is a n optical ring-disk electrode (ORDE) in which a n optically transparent ring surrounds the disk electrode. This arrangement allows recording of optical spectra of unstable materials generated at the disk electrode. The system has been tested by recording the spectra of a stable species (IrC16*-) and a n unstable species (monocation of triphenylamine) which were generated electrochemically. EXPERIMENTAL Apparatus. The arrangement used for the O R D E studies is shown in Figure 1. The shaft of the rotating electrode assembly was machined from a length of 6i8-in. o.d., 3/16-in. i d . , stainless steel tubing (final 0.d. was 9116 in.). The bearing assembly and contacts for electrical connection to the working electrode were duplicated from a rotated disk electrode system which was designed by Stonehart (3) . The shaft was driven through a pulley and belt arrangement with a n Electrocraft Model E500-M electric motor and speed control system. The Electrocraft E500-M is capable of speeds up to 5000 rpm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000830_iet-cta.2010.0622-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000830_iet-cta.2010.0622-Figure3-1.png", + "caption": "Fig. 3 Vector defined for computing the dynamic model", + "texts": [ + " Furthermore, it is possible to write the dynamical equation as function of the actuated joints, ui, that is t + J utext,p = Muu\u0308+ Cuu\u0307+ gu This mathematical expression has two discouraging facts. The first is that it is still required to perform forward kinematics to calculate the position of the end effector because both Jacobians, Jp and Ji, are often expressed in terms of unit vectors in each leg. The second is about the control of the movement of the robot, it is preferable to calculate the position and velocity errors depending on the end effector than calculate the variables function of the joints. The nomenclature used for computing the dynamic model is addressed in Fig. 3. Here, the vector rp defines the position of the end effector and the angles ai and aj define the position of the spherical joint, in the upper ring, and the universal joint, in the lower ring, respectively. The vectors ri and rj define the location of the spherical and universal joint in the upper and lower rings. The unit vector u1,i is along the first rotational union in the lower universal joint, u3,i is a unit vector parallel to the translational actuator and u2,i is IET Control Theory Appl", + "2010.0622 & The Institution of Engineering and Technology 2011 unit vectors along the linear motion and angular velocity, respectively; and w is the periodic motion frequency. The equations for velocity and acceleration are obtained after differentiating the previous one r\u0307 = Arwur cos(wt) (85) r\u0308 = \u2212Arw 2ur sin(wt) (86) v = Aawuv cos(wt) (87) v\u0307 = \u2212Aaw2uv sin(wt) (88) In Table 2 the values used in the path planning simulations are grouped. The robot model was built using the nomenclature shown in Fig. 3. The external disturbances, td, are applied to the end effector and they were modelled with the function given in (81) td = Ad random(1, \u22121) (89) where Ad is the amplitude and it has a value of 1 N for force and 10 mN m for torque. In Table 1 are the values of the variablesai, aj, which define the position of each joint in the upper and lower rings. This table also contains the radius of the rings, rlower and rupper, and the length of the two links located in a leg, L1 and L2. Table 2 has the constant values used in the controller, which are L in (72), K in (78) and G in (79)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000138_2009-01-2632-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000138_2009-01-2632-Figure3-1.png", + "caption": "Figure 3. Automatic transmission test rig configuration", + "texts": [ + " The impact of the VI difference can be seen in the significant differences in KVs at the lower temperatures. See Table 3. Figure 2 shows the measured KV with respect to temperature for these two fluids. At lower temperatures the KV of the high VI fluid is increasingly lower than its low VI counterpart. An in-house, custom built AT efficiency test rig was used to measure relative torque losses in a 6 speed rear wheel drive type passenger car automatic transmission unit. The transmission was connected to a drive motor and two absorbing motors in a T-shape configuration, as shown in Figure 3. The transmission was fully drained and flushed a total of four times between fluids to achieve a complete flush with no measurable carry-over. The rig was modified to allow both the transmission and the fluid to be cooled to approximately -10\u00b0C as shown in Figure 4. Fluid temperature, measured in the sump, was allowed to rise naturally as the test progressed. SAE Int. J. Fuels Lubr. | Volume 2 | Issue 222 The efficiency testing was conducted at four different speed/load combinations as detailed in Table 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001949_12.2257550-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001949_12.2257550-Figure2-1.png", + "caption": "Figure 2: Hardware implementation: (a) the main components of the RUMHMS, and (b) the ground mobile infrastructure testbed with a sphere-shaped target for testing and validation purposes.", + "texts": [ + " Platform B at the site of control center similarly consists of a Wi-Fi access point, directional antenna, and heading control system for directional communication. Experimental studies suggest that the throughput of directional air-to-air (A2A) communication link can reach 48Mbps at a distance of 300m and 2Mbps at a distance 5000m at 5Ghz.23 An air to ground (A2G) is established in Platform B for the control center to access the remote monitoring video stream. In the section, we describe the hardware components that we choose for RUMHMS in our implementation. The prototype system is shown in Figure 2. Proc. of SPIE Vol. 10169 101690A-4 Downloaded From: http://proceedings.spiedigitallibrary.org/pdfaccess.ashx?url=/data/conferences/spiep/92321/ on 04/26/2017 Terms of Use: http://spiedigitallibrary.org/ss/termsofuse.aspx UAV Platform UAV is a vital component for RUMHMS, because it carries the flight control system, camera, communication system, heading control system, batteries, and all other necessary devices. We consider the following factors when selecting a UAV platform: frame weight, takeoff weight, battery life, lift force, flexbility of software development kit (SDK), and stability in strong winds. With such considerations, we select DJI Matrice 100 carbon fiber quadcopter frame as our UAV platform,24 as shown in Figure 2. It weighs 2431g with batteries. Its takeoff weight achieves up to 3.6 kg. Two 5700 mAh lipo batteries can offer the power for over 40-minute flight time. Matrice 100 communicates with many operational systems, such as windows, linux and embedded systems, using a universal asynchronous receiver/transmitter(UART) interface. In addition, Matrice 100 offers a flexible onboard SDK, which allows microcontrollers to easily access and control UAV attitude and set way-points. Microprocessor We choose Raspberry Pi 3 (RP3)25 as the microprocessor for image processing, and UAV and gimbal tracking control" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001242_1.3670079-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001242_1.3670079-Figure1-1.png", + "caption": "Fig. 1. Geometry of a ball rolling on a horizontal surface at speed v in the x direction while the spin axis remains nearly vertical, viewed (a) from above and (b) from the front of the approaching ball. The z-axis is vertical.", + "texts": [ + "org/termsconditions. Downloaded to IP: 129.49.251.30 On: Wed, 10 Dec 2014 21:09:50 DOI: 10.1119/1.3670079 The Physics Teacher \u25c6 Vol. 50, January 2012 25 a bird\u2019s-eye view of a ball projected horizontally in the x direction while spinning about a vertical axis in the z direction. The ball curves in the negative y direction if it spins counterclockwise viewed from above, the curvature increasing with the initial spin. In practice, it is difficult to spin a ball by hand so that the spin axis is exactly vertical. Figure 1(b) shows the spin axis inclined at an angle A to the horizontal, as viewed front-on with the ball approaching the viewer. The contact point on the horizontal surface is directly below the center of the ball, and is located a perpendicular distance r from the spin axis. The contact point therefore rotates at speed rw away from the viewer, relative to the center of mass, while the center of mass itself approaches the viewer at speed v. If v = rw then the contact point is at rest on the surface. The ball can therefore roll forward along a circular path of radius r around the ball", + " A literature search failed to provide explanations, so I filmed the ball at 300 frames per second with a Casio EX-F1 video camera to obtain quantitative data. A side-on view was filmed in order to measure both the linear and angular velocity of the ball. The data presented me with another surprise. The friction force on the ball was quite small, the same as that for a rolling ball. The ball therefore rolled to a stop, despite the fact that the ball kept spinning after the forward motion stopped and despite the fact that the ball appeared visually to slide on the carpet without rolling in the conventional manner. Figure 1 shows the geometry of the experiment and a plausible explanation for the low friction force. Figure 1(a) is This article is copyrighted as indicated in the article. Reuse of AAPT content is subject to the terms at: http://scitation.aip.org/termsconditions. Downloaded to IP: 129.49.251.30 On: Wed, 10 Dec 2014 21:09:50 26 The Physics Teacher \u25c6 Vol. 50, January 2012 tom of the ball was sliding backward or forward) so that the ball commenced rolling almost immediately. From then on, the spin axis slowly tilted into a vertical position to maintain the rolling condition. An additional experiment was performed where the ball was given a relatively small initial spin about a near vertical axis, a situation that is commonly encountered in billiards, golf, and ten-pin bowling" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001138_j.mechmachtheory.2011.07.015-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001138_j.mechmachtheory.2011.07.015-Figure6-1.png", + "caption": "Fig. 6. Partial worm gear drive tooth profiles of (a) standard proportional teeth (b) semi RA teeth and (c) full RA teeth.", + "texts": [ + " The tooth surface equation of the generated RA worm gear proposed herein can be verified by plotting the worm gear profile with computer graphics. Since the tooth surface equation is non-linear, therefore, solving by numerical analysis method is needed. Table 1 lists some design parameters of the hob cutter and generated RA worm gear. Based on the developed mathematical model of the worm gear tooth surface, three-dimensional tooth profiles of semi RA, full RA and standard proportional tooth worm gears are plotted in Fig. 6. A series of worm gears, partial (15/337) teeth are plotted, are of the same pitch circle and throat height, but with different throat circle and outside circle. Points D and E indicate the penetration points of pitch circle with partial (15/337) teeth, while points F and G are the intersection points of the throat circle with generated worm gear. For a full RA worm gear, pitch circle and throat circle pass through the throat of the worm gear, as shown in Fig. 6(c). In the other words, pitch circle and throat circle are identical. But for the standard proportional tooth worm gear (i.e. special case of RA worm gear with dx=0), pitch circle passes through the middle of the tooth height, as shown in Fig. 6(a). 9. Tooth profile comparisons of the generated RA worm gears with double-depth teeth The worm gear design parameters are chosen the same as those listed in Table 1. A series of worm gears, semi RA, full RA and the standard proportional tooth, are generated by using varying pitch lines (refer to Figs. 2 and 3), dx=1.97 mm, 3.94 mm and 0 mm, respectively. Axial cross sections of this series of worm gear teeth are shown in Fig. 7. Figs. 8, 9 and 10 show the tooth profiles of the generated RA worm gears with double-depth teeth at cross sections of Z2=0 mm, Z2=8 mm, and Z2=\u22128 mm (refer to Fig. 6), respectively. Fig. 8 shows the tooth profiles of a series of worm gears at cross section of Z2=0 mm. It is noted that for the full RA worm gear, its pitch circle and throat circle become identical (also refer to Figs. 6(c) and 7), only dedendum part of working height exists, and addendum part shrinks to zero, because of full recess design. Additionally, the fillet curves q1q2 and q3q4 of the full RA worm gear that generated by hob cutter are larger than those of two others, thus theworking height is smaller" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000753_10407790.2011.630949-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000753_10407790.2011.630949-Figure5-1.png", + "caption": "Figure 5. Cantilever plate subjected to shear force at the free end.", + "texts": [ + " Equation (42) can also be represented as Md\u00fe R \u00bc 0 \u00f043\u00de where R is the residual and d is given as d \u00bc xk\u00fe1 xk \u00f044\u00de Here xk is the solution at the kth iteration. Due to implementation of a fully implicit procedure, the system of discretized algebraic equations can be solved within one outer iteration if the convergence criterion for inner iterations in the linear solver is set to a low enough value. For transient problems, this procedure is repeated for each time step. 6.1. Verification Test 1: Cantilever Plate Subjected to Shear Force at Free End The first verification test consists of a cantilever plate, shown in Figure 5, subjected to shear force at the free end. Two aspect ratios are used, L=h\u00bc 10 and L=h\u00bc 100, to study the robustness of the method for both thick and thin plates. D ow nl oa de d by [ N ip is si ng U ni ve rs ity ] at 1 9: 46 0 9 O ct ob er 2 01 4 The deflection of the plate at the free end may be obtained from beam theory with shear effects included: w \u00bc PL3 3D 1\u00fe u 4 \u00f045\u00de where u \u00bc 2 k h L 2 Here P is the shear force on the free end, h and L are the thickness and length of the plate, respectively, and k is the lateral shear correction factor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002009_icmsao.2017.7934926-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002009_icmsao.2017.7934926-Figure3-1.png", + "caption": "Fig. 3. Double ellipsoidal heat source model.", + "texts": [ + " To model the asymmetry profile, parabola conventional mathematics function is used. The parabola model function is given by (1). 2y ax h (1) The experimental result is shown in Fig. 2(a). The value of the height is 3.13 mm and the width is 11.4 mm, which can be used to calculate the parabola model parameters, namely a=0.0963 and h=3.13. The modeling weld-bead profile is shown in Fig. 2(b) A moving heat source model, namely Goldak double ellipsoidal, is employed in the thermal model [15]. The heat distribution is given by (2), as show in Fig. 3. 2 2 2 2 2 2 3( / ) 3( / ) 3( / ) 3( / ) 3( / ) 3( / ) 6 3 ( , , ) 6 3 ( , , ) f r x af y b z d f f x a y b z dr r r f Q q x y z e e e a bd f Q q x y z e e e a bd , 2f rQ VI f f (2) where the parameters of a, b, c, f are obtained from the experimental welding pool shape, which is shown in Fig. 2(a). The simulating welding pool shape is presented as the red region of Fig. 2(b). The simulation result is suitable for the experimental result. III. FINITE ELEMENT MODELS The three-dimensional finite element model is composed of 9 clamping, a substrate and single-pass 6 layers component, which is shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000222_j.rcim.2009.09.003-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000222_j.rcim.2009.09.003-Figure1-1.png", + "caption": "Fig. 1. The 3-SPR parallel manipulator and its simulation mechanism.", + "texts": [ + " Among them, a 3-SPR parallel machine tool has attracted more attention due to its simple configuration, larger workspace, and easy control. In this paper, a novel CAD geometric variation approach is proposed for machining complex workpiece or carving letters on a 3D free-form surface or a 2D plane by using a 3-SPR parallel machine tool. 2. A 3-SPR parallel machine tool and workspace 2.1. The 3-SPR parallel manipulator A 3-SPR parallel manipulator is composed of a moving platform m, a fixed base B and 3 identical SPR (spherical jointactive prismatic joint-revolute joint) active legs ri (i=1, 2, 3) with linear actuators, see Fig. 1a. Here, m is an equilateral triangle link with 3 sides li= l, 3 vertices ai and a central point o. B is an equilateral triangle link with 3 sides Li=L, 3 vertices Ai and a central point O. Each of ri connects m to B by a revolute joint R at Nomenclatures Symbol description S the 3D free-form surface B, m the base and the moving platform M the milling machine tool P0 the guiding plane of tool path curve P, R the prismatic joint and the revolute joint S the spherical joint pw the point of free end of tool z {m} the coordinate of o-xyz fixed on m {B} the coordinate of O-XYZ fixed on B ex the distance from pw to vertical side ey the distance from pw to horizontal side Xo, Yo, Zo the 3 translation components of m xl, xm, xn direction cosine between x & X, x & Y, x & Z yl, ym, yn direction cosine between y & X, y & Y, y & Z zl, zm, zn direction cosine between z & X, z & Y, z & Z ri the active leg i=1, 2, 3 z the cutter tool of machine tool w the tool path curve of letters li the side of m o1 central point on P0 ai the vertices of m Ai the vertices of B e the distance from ai to o E the distance from Ai to O xO, yO, zO the 3 components of O on B in {m} sj prescribed spline of S, j=1, 2, y k xai, yai, zai the 3 components of ai in {m} xAi, xAi, xAi the 3 components of Ai in {m} ", + " Each axis of revolute joints R at ai on m is parallel with li, thus there are the structure constraints ri?li. In fact, the 3-SPR parallel manipulator is an inverse mechanism of the 3-RPS parallel manipulator with 3 DOFs [18]. Therefore, the 3-SPR parallel manipulator has 3 DOFs. When a milling tool T is attached onto the platform m at o, and the axis of tool and z are collinear, a 3-SPR parallel machine tool is constructed. 2.2. The simulation mechanism By using the geometric variation function of Solidworks [25], a simulation mechanism of the 3-SPR parallel manipulator is created, see Fig. 1b. The creation procedures are explained as follows. 1. Construct a base B in 2D sketch. The sub-procedures are: a. Construct an equilateral triangle DA1A2A3 by the polygon command. b. Coincide its center point O with origin of default coordinate, set its one side horizontally, and give its one side a fixed dimension Li=L=120 cm in length. c. Transform DA1A2A3 into an equilateral triangle plane by the planar command. 2. Construct a moving platform m in 3D Sketch. The sub- procedures are: a. Construct 3 lines li (i=1, 2, 3), and connect them to form a closed triangle Da1a2a3 by the point-point coincident command", + " (1a), (1b) and (2a), Ai m and ai are derived as follows: Am 1 \u00bc 1 2 aExl Exm\u00fe2xO aEyl Eym\u00fe2yO aEzl Ezm\u00fe2zO 2 64 3 75; Am 2 \u00bc Exm\u00fexO Eym\u00feyO Ezm\u00fezO 2 64 3 75; Am 3 \u00bc 1 2 aExl Exm\u00fe2xO aEyl Eym\u00fe2yO aEzl Ezm\u00fe2zO 2 64 3 75; a1 \u00bc 1 2 aexl eyl\u00fe2Xo aexm eym\u00fe2Yo aexn eyn\u00fe2Zo 2 64 3 75; a2 \u00bc eyl\u00feXo eym\u00feYo eyn\u00feZo 2 64 3 75; a3 \u00bc 1 2 aexl eyl\u00fe2Xo aexm eym\u00fe2Yo aexn eyn\u00fe2Zo 2 64 3 75: \u00f03\u00de The constraint equations of (xl, xm, xn, yl, ym, yn, zl, zm, zn) can be obtained from Ref. [1,2]. From the 3-SPR parallel manipulator in Fig. 1a, it is known that Ai m of B in {m} are constrained to move in the 3 planes x= ay, x=0 and x=ay, respectively. Thus, the 3 plane constraint conditions of Ai m are yielded as follows, xA1 \u00bc ayA1; xA2 \u00bc 0; xA3 \u00bc ayA3 \u00f04\u00de From Eqs. (3) and (4) and the constraint equations of (xl, xm, xn, yl, ym, yn, zl, zm, zn), the 3 constraint equations are derived as follows: aExl Exm\u00fe2xO \u00bc a\u00f0aEyl Eym\u00fe2yO\u00de \u00f05a\u00de xO \u00bc Exm \u00f05b\u00de aExl Exm\u00fe2xO \u00bc a\u00f0 aEyl Eym\u00fe2yO\u00de \u00f05c\u00de From Eqs. (5a\u2013c), three equations are derived as follows: yO \u00bc E\u00f0ym xl\u00de=2 Exm\u00fe2xO \u00bc 3Eyl xm \u00bc yl \u00f06\u00de Let a rotation transform matrix RB m be defined by 3 Euler rotations of (XY1X2), namely, a rotation of a about X-axis, followed by a rotation of b about Y1-axis, and a rotation of l about X2-axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003565_iciaict.2019.8784855-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003565_iciaict.2019.8784855-Figure1-1.png", + "caption": "Fig. 1. An AUV six degree of freedom [1]", + "texts": [ + " The organization of this paper is as follows: In Section II, the nonlinear dynamics of an AUV is presented, followed by a brief description of interval type-2 fuzzy control in Section III. The adaptation mechanism of our fuzzy parameters is discussed in Section III. In Section IV, the results and discussion are presented. Lastly, in Section V, a conclusion is derived. II. AUVS NONLINEAR DYNAMICS An AUV has six degree of freedom (DOF), and the equation of motion can be described with respect to a fixed frame (x,y,z) and body frame (X ,Y,Z), as shown in Fig 1. The forward, lateral and vertical speeds can be represented by [u,v,w]T across the (x,y,z) axis, while the moments acting on 978-1-7281-3745-2/19/$31.00 \u00a92019 IEEE 19 the (x,y,z) axis can be described as [L,M,N]T , respectively. The angular rates are described by [px, py, pz] [25], [26]. The dynamics can be written as follows [27]: M\u03b7\u0308 +C (\u03b7\u0307) \u03b7\u0307 +D(\u03b7\u0307) \u03b7\u0307 +G(\u03b7) = \u03c4\u03b7 (1) where M is the inertia matrix, which is a symmetric and positive definite (6x6) matrix, C(\u03b7\u0307) is a matrix that describe the Coriolis and Centripetal forces resulted due to added mass, D(\u03b7\u0307) is the matrix of the hydrodynamic damping term, G(\u03b7) is a (6x1) matrix, which represents the vehicle\u2019s buoyancy and gravitational forces and \u03c4\u03b7 denotes the control efforts" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001647_pi-c.1962.0054-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001647_pi-c.1962.0054-Figure1-1.png", + "caption": "Fig. 1.\u2014Stator and rotor angular references.", + "texts": [ + " The method of approach here is to regard any air-gap flux-density distribution as the superposition of components having this simple form. Attention is directed initially to the general form of the fluxdensity component as follows: It is assumed that the air-gap flux density is uniform in the axial direction, and end effects are ignored. Thus this flux density, b, is a function of angular position, d, and of time, /, i.e. b = b(d,t) (1) 6 may be measured with respect to a reference in the stator, 6S, or in the rotor, dr, as shown in Fig. 1. For reasons which will appear later these reference lines are taken at right angles to the magnetic axes of the stator and rotor windings in the single-phase case. For polyphase machines this relation is preserved for one reference phase, as shown in Fig. 1 for the 3-phase stator. The angle 6 is taken in electrical radians for the basic number of poles of the machine. If o\u00bbr is the rotor velocity (electrical radians per second), assumed constant, then \\ = BS- wrt (2) where a is the separation of the two reference lines when f is zero, and may be made to vanish by a suitable choice of initial time. Since this choice is arbitrary in the cases considered, it will simplify expressions to make a zero; then er = es-u>rt (3) The general form of the flux-density component can now be defined as b = B sin {kojt \u00b1 qds + ) ", + " Taking a single-phase stator winding, for example, we shall define the variable n as the number of conductors of the winding passed over in going from the stator reference to position 6S, due account being taken of the sense in which the conductor is connected into the winding. Thus n is a periodic function of 6, and can be expanded as a Fourier series. This series will have only odd harmonics in virtually all cases, owing to symmetry. Further, only sine terms will occur, provided that the reference for 6S is taken at the centre of the winding, as shown in Fig. 1. Thus n= YA NJ sin j6. (7) \u2022 = 1 , 3 , . . . 382 ROBINSON: HARMONICS IN A.C. ROTATING MACHINES Such a winding, if carrying a current / = /(/), will produce an m.m.f. / = \u00ab / = , - 2 N: sin j9s . . . . (8) y\u20141,3 In this respect, therefore, the actual winding can be considered as equivalent to a series connection of windings having conductor distributions Nt sin 9S, N3 sin 39S, etc. The computation of the harmonic-winding coefficients NJt for an actual winding is described in Appendix 9.1. (4.2) Voltage Induced in Winding This voltage is given by v = \u2014 dip/dt volts, where ip is the flux linkage with the winding, in weber-turns", + " wave being in one direction. It is easily shown that average torque is produced in this case only with j = q and with m = k, as before, and with the further condition that the time sequence of currents and space sequence of windings, are such as to produce ROBINSON: HARMONICS IN A.C. ROTATING MACHINES 383 an m.m.f. rotating in the same direction as the component bwave. Then L a v = \\ T r l r B q N q I k c o s (cf> - e k ) . . . ( 2 2 ) where ek now refers to the reference phase, located as in Fig. 1. The question might arise here as to the validity of this result for torque, depending upon the source of the flux-density component, i.e. whether it arises from the same side of the airgap or not. The flux-density wave set up by a winding carrying current produces no average torque when reacting with its own current, since cos ( \u2014 e) = 0 in this case. Only if a displacement occurs between m.m.f. and flux waves (due, for example, to saliency, as in a reluctance motor) will torque be developed", + " In general, owing variously to winding distribution, saliency and saturation, it will not be sinusoidaU and can be regarded as the sum of a number of components. The general form of any flux-density component, viewed from the rotor, is given by eqn. (6): b = B sin [{k(x> \u00b1 qoir)t \u00b1 qdr + ] In this case, the ^-distribution is time-invariant, as seen from the rotor, and assuming symmetry about the winding or pole axis, takes the form b= S Bj sin (70,) . . . . (23) y - 1 . 3 . . . . the phasing being taken to accord with Fig. 1. To examine its effect upon the stator, we must transform this equation from 0r to ds, using eqn. (3): b = HBj sin (jiot ~jds + IT) (24) taking co = w,., since this situation arises in synchronous machines. Comparing this expression with the standard form, eqn. (4), k = q = j for the 7th harmonic term. This term therefore has/ times the basic number of poles, and rotates at velocity co (which is more directly obvious in this case). It will therefore induce voltage in the stator windings only to the extent that space harmonic windings exist of the /th-harmonic order" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000758_s11465-012-0317-4-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000758_s11465-012-0317-4-Figure9-1.png", + "caption": "Fig. 9 The 3-UPU/2-PRUC redundantly actuated parallel mechanism", + "texts": [ + " Therefore, we rewrite the configuration rules for type synthesis of redundantly actuated limbs, Rule 1 For actuated joint, either revolute joint or prismatic joint can be oriented arbitrarily. Rule 2 For passive joint, the revolute joint axes must be perpendicular to the actuated couple, while the prismatic joint axes can be orientated arbitrarily as long as they are linearly independent. The possible configurations can be referred to Table 2 in Ref. [29]. Now we take an example to demonstrate the possible mechanism with redundantly actuated limbs, as shown in Fig. 9. Figure 9 illustrates a 3-UPU/2-PRUC parallel mechanism constructed with two redundantly actuated limbs. Two actuated couples provided by these two redundantly actuated limbs, $r4,a1 and $r5,a1 , satisfy the conditions: sr4==\u00f0s43 s44\u00de and sr5==\u00f0s53 s54\u00de. When the actuators are locked, these two actuated couples constrain the possible parasitic rotations, which belong to the flat pencil defined by the two actuated couples, $r4,a1 and $r5,a1 . The parasitic rotation evaluation and avoidance of the 3- UPU parallel mechanism is performed and the bound of instability of SNU parallel mechanism is obtained in this paper" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000239_sii.2010.5708334-Figure13-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000239_sii.2010.5708334-Figure13-1.png", + "caption": "Fig. 13. A side view and a front view of UAV.", + "texts": [ + " We selected the motor and the propeller of which the static thrust amounting to 200 [%] of the fuselage weight at a continuous maximum motor load. The measurement results of actual static thrust are plotted in Fig. 11. A lithium polymer battery is mounted on the UAV, which produces output power as plotted in Fig. 11. The battery capacity is 2100 [mAh], the voltage rating is 11.1 [V] and the weight is 0.16 [kg]. The calculated maximum cruising speed of the UAV is about 13 [m/s]. A configuration diagram of the electronic system is shown in Fig. 12. A setup of equipments on the fuselage is shown in Fig. 13. A commercially available attitude sensor module (Microstrain Co., 3DM-GX1) is employed. In addition to the attitude, three-axis angular velocity and acceleration are obtained from this module. The sensor\u2019s datasheet gives its attitude angle accuracy as \u00b12 [\u25e6]. An Ultrasonic distance sensor measures altitude from ground. Resolution of the ultrasonic distance sensor is 0.025 [m]. The ultrasonic sensor range is 6.45 [m]. A global positioning system (GPS) receiver module (Garmin Co., GPS 18-5Hz) which incorporates an antenna and an arithmetic chip outputs absolute position on the earth and velocity of three axes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001381_kem.490.237-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001381_kem.490.237-Figure4-1.png", + "caption": "Fig. 4. Generating system of the gear", + "texts": [], + "surrounding_texts": [ + "Due to simultaneous cutting of the both sides of the tooth space (pinion and gear), the bearing contact bias occurs after assembling of a transmission made in that way. Such an effect is undesirable since it causes loud work of the transmission and non-uniform transfer of the motion which leads to the faster fatigue wear. In order to remove the bias the helical motion is applied during pinion finishing. It is realized by the axial offset of the fixed headstock of the workpiece connected with the generating gear rotation. The generating system of the pinion is the technological hypoid gear obtained thanks to the hypoid offset of the workpiece\u2019s axis by the E value with respect to the cradle\u2019s axis. The cradle with inclined toolhead (tilt) creates the bevel generating gear. For the pinion cutting the head\u2019s axis inclination is applied in order to compensate the difference between the pressure angle and tool profile so in the initial position the c S system is rotated with respect to the d S system by the angle of j around d X axis. The head\u2019s axis inclination angle (in degrees and minutes) and the value of the hypoid technological offset E is determined on the basis of the gear geometry analysis and calculation cards. Positive values mean down-shift by the pinion with left inclination line direction of a tooth or up-shift by the pinion with right inclination line direction of a tooth. The negative value means upshift by the pinion with left inclination line direction of a tooth or down-shift by the pinion with right inclination line direction of a tooth. Mathematical model of tooth flank surface Mathematical notation of the tooth flank is apparent from the equation of the surface of action of the tool, kinematics, and accepted treatment technological system. The following discussion presents the side of the tooth surface obtained with the use of a technological system with a bevel generating gear - this is the case more generally in comparison with a ring generating gear. While processing the envelope, the equation of the tooth flank, which is the bounding surface of the utility, is determined from the system of equations [1, 2, 6]. This system includes the equation of the family of tool surfaces and the equation of meshing, resulting from the method for determining the kinematic envelope: ( ) ( ) , , , , 0 t t t t t t s s \u03b8 \u03c8 \u03b8 \u03c8 \u22c5 = 1 t1 1 1 r n v (1) where: ( ), ,t t ts \u03b8 \u03c81r - determining the vector function of the family tool surfaces system bounded with treating pinion ( 1S ), 1 n - the unit normal vector defined in 1S , ( ), ,t t ts \u03b8 \u03c8t1 1v - the relative velocity vector defined in 1S . Based on the defined technological model, the family of tool surfaces is determined as follows: ( ) ( ) ( ), , ,t t t t t ts s\u03b8 \u03c8 \u03c8 \u03b8= \u22c51 1t tr M r (2) where: ( ),t ts \u03b8tr - vector equation of the tool surface referred to the system associated with the tool, tS , ,t ts \u03b8 - curvilinear coordinates surface form, ( )t\u03c81tM - the transformation matrix being the product of a transformation matrix representing the rotations and translations of homogeneous coordinate systems included in the technological gear model t\u03c8 - parameter of motion (in this case - angle of the cradle rotation). Vector equation of the tool surface as a function of curvilinear coordinates ,t ts \u03b8 shows the relationship (3), which involves the processing of the active side of the tooth. ( ) ( ) ( ) cos sin , sin sin cos t wk t wk t t t wk t wk t wk r s s r s s \u03b8 \u03b1 \u03b8 \u03b8 \u03b1 \u03b1 + = + \u2212 tr (3) where: wk\u03b1 - the angle of the external blades, wkr - the radius of the cutterhead. Based on the model of technological gear it is designated a family of the tool surfaces according to equation (2), for which it determines conversion matrix equation (4). ( ) ( )( ) ( )1t t t\u03c8 \u03c8 \u03c8 \u03c8= \u22c5 \u22c5 \u22c5 \u22c5 \u22c5 \u22c5 \u22c51t 1w wr rh hm mk kc cd dgM M M M M M M M M (4) where: ( )t\u03c81tM - conversion matrix, ij M - the elementary transformation matrices representing the rotations and translations of homogeneous coordinate systems, ij - subscript indicating the direction of the transformation from system jS to system iS . Technological gear model is also used to determine (in the chosen system, for example 1(S or )mS a unit normal vector and relative velocity. The normal vector to the surface of the tool sets at any of the predefined layouts. In this way we can get the components of meshing equation. Solving equations (1) by eliminating one of the variables such as: ts from the meshing equation and then substituting into the equation of the family of tool surfaces ( )1 , ,t t ts \u03b8 \u03c8r we can obtain the tooth flank surface equation in two-parametric form (5). ( ) ( )( ), , , ,t t t t t t ts\u03b8 \u03c8 \u03b8 \u03c8 \u03b8 \u03c8=1 1r r (5) where: ( ),t t\u03b8 \u03c81r - the equation of the pinion tooth surfaces in the two-parametric form, ( ),t t ts \u03b8 \u03c8 - the variable ts in the function of other parameters. Model of technological gear is designed to create the flank surface of the gear and pinion teeth, which will be used for the analysis of meshing for constructional spiral bevel gear. In order to obtain a tooth surface as the family of tool surfaces it should be designated in the system rigidly bounded with cutting pinion S1 , which represents the envelope to family of surfaces of the cut gear tooth ( )\u03a3 1 . Family of tool surfaces is shown in the following way: ( ( )) ( ( ( ) ( )) )\u03c8 \u03c8 \u03c8\u2190 \u2190 \u2190 \u2190 \u2190 \u2190 \u2190 \u2190 \u2190= \u22c5 \u22c5 \u22c5 \u2212 + \u22c5 \u22c5 \u22c5 \u2212 \u2212(t) (t) 1 1 w 1 t w r r h h m m k t c d d t t k c r hr L L L T L L L r T T (6) In this equation the vector record (t) 1 r , (t) m r , (t) gr , concerns a family tool surfaces ( )\u03a3 t (as evidenced by the superscript ( t )), set respectively in the cutting gear system S1 , the basic system (stationary) and the tool system. Other signs are bounded with the transformations of coordinate systems used in the model of technological gear. The applied rotation and translation matrixes of the coordinate system are as follows: cos i 0 sin i 0 1 0 sin i 0 cos i , \u2190 = \u2212 d g L (7) cos j sin j 0 sin j cos j 0 0 0 1 , \u2190 = \u2212 c d L (8) U cos q U sin q 0 , \u2190 \u2212 \u22c5 = \u2212 \u22c5 k c T (9) cos sin 0 ( ) sin cos 0 0 0 1 , \u03c8 \u03c8 \u03c8 \u03c8 \u03c8\u2190 \u2212 = t t m k t t t L (10) 1 B1 1 0 A X p ( ) ,\u03c8 \u2190 = \u2212 + h m t" + ] + }, + { + "image_filename": "designv11_33_0001093_icra.2012.6225142-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001093_icra.2012.6225142-Figure2-1.png", + "caption": "Fig. 2. Convex Version of the Omnidirectional gear", + "texts": [ + " Versions of the omnidirectional driving gear mechanism 1) Planar version The planar version of the omnidirectional driving gear mechanism is defined as a gear structure with gear columns that are perpendicular to each other and configured and deployed on the surface with zero curvature (Fig. 1). The whole structure is a flat plane. 2) Convex version The convex version of the omnidirectional driving gear mechanism is defined as a gear structure with two crossed gear columns that are configured and deployed on the outer surface, the curvature of which is positive (Fig. 2). 3) Concave version The concave version of the omnidirectional driving gear mechanism is defined as a gear structure with two crossed gear columns that are configured and deployed on the inner surface, the curvature of which is negative (Fig. 3). Study on the Omnidirectional Driving Gear Mechanism Kenjiro TADAKUMA, Riichiro TADAKUMA, Kyohei IOKA, Takeshi KUDO, Minoru TAKAGI, Yuichi TSUMAKI, Mitsuru HIGASHIMORI, Makoto KANEKO, Member, IEEE I May 14-18, 2012 978-1-4673-1405-3/12/$31.00 \u00a92012 IEEE 3531 III" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003259_s00158-019-02237-3-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003259_s00158-019-02237-3-Figure7-1.png", + "caption": "Fig. 7 Structure of the supporting ring for locating on double sides", + "texts": [ + " Through the deformation analysis and the contact status analysis of the supporting ring, it was displayed that the radial deformation of the disc hub was slightly larger than one of the supporting ring at working stage; such result made the inside locating surfaces of the supporting ring ineffective. Therefore, the single side locating mode in the patent of Cairo (1999) seems not to be applicable with respect to TWD in this paper. Therefore, a kind of double sides locating mode for the supporting ring is proposed in this paper, shown in Fig. 7. Off the working state, the inside locating surfaces take responsibility to locating, while the outside locating surfaces are designed to separate initially off disc hub, with a reserved clearance, the magnitude of which is equivalent to the radial deformation difference between the supporting ring and the disc hub. Then, at the working state, the locating responsibility will be replaced by outside locating surfaces with deformation. Furthermore, the contact pressure has been validated meeting requirements due to the reserved radial clearance" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002158_gt2017-64151-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002158_gt2017-64151-Figure3-1.png", + "caption": "Figure 3. Radial and tangential forces acting on the cage by hydraulic force.", + "texts": [ + " However, after passing through the seal in the liquid turbo-pump, cryogenic fluid passes quickly between the cage and the race land in the axial direction. Thus, it is appropriate to approach this phenomenon using seal theory with axial flow through the differential pressure between the inlet and outlet rather than using short bearing theory. 3 Copyright \u00a9 2017 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 08/23/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use The radial and tangential forces on the cage surface generated by the fluid passing between the cage and the race land are depicted in Fig. 3. The cage not only rotates with respect to the mass center of the cage but also revolves around the center of the inner race with the whirling amplitude. At this time, the radial and tangential forces act on the cage because the fluid passes quickly through the small clearance between the cage and race land. Thus, the dynamic motion of the cage is affected by hydraulic force, not just the interaction between the ball bearing elements. The clearance between the inner race land and the cage is larger than the clearance between the outer race land and the cage, because the outer race\u2013guided cage was applied" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000451_978-3-642-29329-0_6-Figure6.3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000451_978-3-642-29329-0_6-Figure6.3-1.png", + "caption": "Fig. 6.3 Forces acting on the vehicle model", + "texts": [ + " At steady-state the derivatives vanish and we have ~ax \u00bc vr; ~ay \u00bc u2 r : In the following, we will use the standard notation du=dt \u00bc _u, dv=dt \u00bc _v. We consider the motion of the vehicle (a flat rigid body) in the x y plane. We have three state variables, namely u, v, and r, and we have to write three equations describing the dynamic equilibrium, namely the two equations describing the balance of the forces acting along the x-axis and y-axis, respectively, and the equation describing the balance of the moments of the forces around the z-axis. The forces acting on the vehicle model are depicted in Fig. 6.3. The three balance equations read m\u00f0 _u vr\u00de \u00bc \u00f0Fx11 \u00fe Fx12\u00de cos\u00f0d\u00de \u00f0Fy11 \u00fe Fy12\u00de sin\u00f0d\u00de \u00fe \u00f0Fx21 \u00fe Fx22\u00de 1 2 rSCxu 2; m\u00f0 _v\u00fe ur\u00de \u00bc \u00f0Fx11 \u00fe Fx12\u00de sin\u00f0d\u00de \u00fe \u00f0Fy11 \u00fe Fy12\u00de cos\u00f0d\u00de \u00fe \u00f0Fy21 \u00fe Fy22\u00de; Iz _r \u00bc a\u00bd\u00f0Fx11 \u00fe Fx12\u00de sin\u00f0d\u00de \u00fe \u00f0Fy11 \u00fe Fy12\u00de cos\u00f0d\u00de b\u00f0Fy21 \u00fe Fy22\u00de\u00fe t 2 \u00f0Fx11 Fx12\u00de cos\u00f0d\u00de \u00fe \u00f0Fx21 Fx22\u00de cos\u00f0d\u00de \u00f0Fy11 Fy12\u00de sin\u00f0d\u00de : The above equations can be simplified by assuming that the force at the left side of an axle is equal to the corresponding one at the right side. Additionally we will add the forces at the left and right side of an axle, obtaining Fx1 \u00bc Fx11 \u00fe Fx12 ; Fx2 \u00bc Fx21 \u00fe Fx22 ; Fy1 \u00bc Fy11 \u00fe Fy12 ; Fy2 \u00bc Fy21 \u00fe Fy22 ; Moreover, assuming that the steering angle d is small (typical values are between 15 and +15 degrees) we can set cos\u00f0d\u00de 1 e sin\u00f0d\u00de d" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001306_20110828-6-it-1002.03784-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001306_20110828-6-it-1002.03784-Figure1-1.png", + "caption": "Fig. 1. Inertia and body frames", + "texts": [ + " The rest of this paper is organized as follows: Section 2 describes the introduced aircraft flight mechanism. Section 3 presents the complete dynamical model of the aircraft. The control algorithm is presented in section 4. The performance of attitude and altitude control at the hovering condition is demonstrated in section 5. Finally, some concluding remarks will be given in section 6. 978-3-902661-93-7/11/$20.00 \u00a9 2011 IFAC 10385 10.3182/20110828-6-IT-1002.03784 In this section, we describe briefly the configuration of the proposed hybrid aircraft. The hybrid UAV scheme is depicted in Figure 1. This UAV is a special aircraft model which is constructed from a combination of quadrotor and single rotor fixed-wing platforms. It has two pair of counter-rotating motors (M1,M3) and two single rotors (M2,M4) in a cross configuration. Let\u2019s consider M2 and M32 rotating clockwise, M4 and M31 anticlockwise and M11, M12 rotating similar to the front motors. The rear counter-rotating rotor is able to tilt towards (backwards) the head (tail) to convert from helicopter (airplane) to airplane (helicopter) mode", + " In the airplane mode, the rear counter-rotating motor is parallel to the wing, the left, right and front motors are completely stop and the aircraft is controlled by airplane actuators. In this mode the hybrid UAV is similar to the single engine fixedwing airplanes and has the same control mechanism. The airplanes flight control is a well-known concept and has been well studied in the literatures, see e.g., Etkin and Reid [1996]. In this section, the mathematical dynamic model of the hybrid aircraft in helicopter, transition and airplane modes will be derived. The inertial and body frames are shown in Figure 1. The inertial frame F I is an north-eastdown (NED) coordinate system. The orientation of body frame with respect to the inertia frame is represented by euler angles \u03b7 \u2206 = (\u03c6, \u03b8, \u03c8) roll, pitch, yaw (see Figure 2). Let\u2019s define \u03be \u2206 = [x, y, z] T position in the inertia frame, vb \u2206 = [u, v, w] T body frame velocities and the body frame angular rates as \u03c9b \u2206 = [p, q, r] T . The relation between \u03be and vb is obtained by rotational matrix R(\u03c6, \u03b8, \u03c8), i.e. \u03be\u0307 = Rvb, R = ( c\u03b8c\u03c8 c\u03c8s\u03b8s\u03c6 \u2212 s\u03c8c\u03c6 c\u03c8s\u03b8s\u03c6 + s\u03c8c\u03c6 s\u03c8c\u03b8 s\u03c8s\u03b8s\u03c6 + c\u03c8c\u03c6 s\u03c8s\u03b8c\u03c6 \u2212 c\u03c8s\u03c6 \u2212s\u03c8 c\u03b8s\u03c6 c\u03b8c\u03c6 ) (1) where ck = cos(k), sk = sin(k)", + " This subsection aims to derive a dynamic model of a generic 6 DOF rigid-body equation by Newton-Euler approach. In order to develop the equation of motion, both translational and rotational motion must be considered. Equation (3) describes the dynamic motion of a 6 DOF rigid body. [ mI3\u00d73 03\u00d73 03\u00d73 J ]( v\u0307b \u03c9\u0307b ) + [ \u03c9b \u00d7 (mvb) \u03c9b \u00d7 (J\u03c9b) ] = ( F \u03c4 ) (3) where m[kg] is the mass of aircraft, J is the inertia matrix, I is a identity matrix, F is a total force applied to the aircraft body and \u03c4 is the aircraft torque vector with respect to the body frame. From Figure 1, it can be observed that the aircraft is symmetric along the y-axis in the xy-plane and yz-plane. In addition, we can approximate the aircraft CG to be on the xz-plane. Thus, it is reasonable to assume that the corresponding inertia matrix has a diagonal form J = diag(Jxx, Jyy, Jzz) Edwards [2010]. Let\u2019s denote F = ( fx fy fz ) T as the forces acting in x, y and z directions and \u03c4 = ( \u03c4\u03c6 \u03c4\u03b8 \u03c4\u03c8 ) T as the aircraft torques in different orientations. Then, by substituting equation (1) and (2) in (3) and expanding it, the aircraft motion dynamic can be derived" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002278_j.ifacol.2017.08.160-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002278_j.ifacol.2017.08.160-Figure2-1.png", + "caption": "Fig. 2. Arm and Leg Patterns", + "texts": [ + " For the purpose of controller design, the individual DOF of the simulated skydiver were organized into movement patterns : combinations of limbs that move synchronously in order to achieve a specific flight movement. Two such patterns were defined: one associated with moving the legs allowing for forward and backward flying, and another one associated with the arms making it possible to turn right and left. Each pattern is defined by Proceedings of the 20th IFAC World Congress Toulouse, France, July 9-14, 2017 373 372 Anna Clarke et al. / IFAC PapersOnLine 50-1 (2017) 369\u2013374 one control signal - the signed amplitude of the deflection angles of the limbs defining the pattern relative to a neutral pose. Fig. 2 shows the arm pattern defined by rotations about [Xleft, Xright, Zleft, Zright]shoulder = [\u2212\u03bd, \u03bd, \u03bd, \u03bd], where \u03bd is the control input u1(t); and the leg pattern defined by rotations about [Xleft, Xright]knee = [\u00b5, \u00b5], [Xleft, Xright]hip = [\u03c3, \u03c3], where \u00b5 is the control input u2(t) and \u03c3 = \u03c0 180 (\u22120.0035(\u00b5 180 \u03c0 )2\u2212 0.0335(\u00b5 180 \u03c0 )+ 17). The patterns were constructed empirically: close to the movements observed in humans. The controller design objective is to track linear and angular velocity commands by the means of two inputs to the non-linear skydiver plant: \u2019legs\u2019 and \u2019arms\u2019 patterns defined above", + " 7, is that executing the arm pattern induces a backward slide in addition to turning. This effect is well known among skydivers and constitutes a serious challenge for novice jumpers. The design specifications included a zero steady-state error requirement, a servo specification providing the required maneuver agility, and reasonable closed-loop sensitivity and cross-coupling specifications. Proceedings of the 20th IFAC World Congress Toulouse, France, July 9-14, 2017 Anna Clarke et al. / IFAC PapersOnLine 50-1 (2017) 369\u2013374 373 Fig. 2. Arm and Leg Patterns Guidance Algorithm F11 F22 \u03a3 \u03a3 G11 G21 G22 \u03a3 Plant desired path yaw rate com + speed com + + arm pattern leg pattern position orientation + yaw rate speed \u2212 \u2212 Fig. 3. Block Diagram: The plant is given by equations (3),(4) which include equations (9),(10) governed by arm and leg patterns; The controller is given by equations (16),(17). one control signal - the signed amplitude of the deflection angles of the limbs defining the pattern relative to a neutral pose. Fig. 2 shows the arm pattern defined by rotations about [Xleft, Xright, Zleft, Zright]shoulder = [\u2212\u03bd, \u03bd, \u03bd, \u03bd], where \u03bd is the control input u1(t); and the leg pattern defined by rotations about [Xleft, Xright]knee = [\u00b5, \u00b5], [Xleft, Xright]hip = [\u03c3, \u03c3], where \u00b5 is the control input u2(t) and \u03c3 = \u03c0 180 (\u22120.0035(\u00b5 180 \u03c0 )2\u2212 0.0335(\u00b5 180 \u03c0 )+ 17). The patterns were constructed empirically: close to the movements observed in humans. 3.2 Design Strategy The controller design objective is to track linear and angular velocity commands by the means of two inputs to the non-linear skydiver plant: \u2019legs\u2019 and \u2019arms\u2019 patterns defined above", + " This effect is well known among skydivers and constitutes a serious challenge for novice jumpers. The design specifications included a zero steady-state error requirement, a servo specification providing the required maneuver agility, and reasonable closed-loop sensitivity and cross-coupling specifications. Proceedings of the 20th IFAC World Congress Toulouse, France, July 9-14, 2017 Two different \u2019arm\u2019 patterns were considered for actuating the lateral control loop: one employed by experienced skydivers, defined in Fig. 2, and another one usually observed in novices. The \u2019novice\u2019 pattern is also defined by four angles: rotation about [Xleft, Xright, Yleft, Yright]shoulder = [\u2212\u03bd, \u03bd,\u2212\u03bd, \u03bd], where \u03bd is the control input. The analysis of the two patterns from the control theory standpoint reveals an interesting phenomenon, known from skydiving practice but not having until now a theoretical explanation. The \u2019novice\u2019 transfer function from arm pattern to yaw rate has a resonance - anti-resonance pair around 4 rad/sec (see Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000777_s0025654411040017-Figure17-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000777_s0025654411040017-Figure17-1.png", + "caption": "Fig. 17.", + "texts": [ + " Thus, the problem of studying stability of the system in the case where only positional (potential and circular) forces are present is an ill-posed problem in the sense that an arbitrarily small perturbation in the form of added dissipative forces can change the result arbitrarily strongly. The stability threshold varies jump-wise. The above presentation can be illustrated as follows. As d \u2192 0, the characteristic polynomial (5.6) permits calculating the real part of the root \u03bb as a function of n. In turn, the norm of circular forces n is proportional to the rolling friction coefficient (1.1). The graph of Re \u03bb(v) is shown in Fig. 17. Then note that if the instability caused by the introduction of dissipative forces is \u201cslack\u201d (an exponential increase in the vibrations is determined by the exponent that is of the order of d), then the exponent of the vibration increase exponential is already not small outside the stability region (5.8). Since the stability threshold strongly depends on the structure of the matrix of dissipative forces, to suppress the instability, it is expedient to use adjustable dampers of the fore and rear wheels so as to MECHANICS OF SOLIDS Vol" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002938_icelmach.2018.8507253-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002938_icelmach.2018.8507253-Figure4-1.png", + "caption": "Fig. 4. Thermal analysis for transient operating point 60 Nm \u2013 10,000 rpm during 20 seconds: (a) FEA + correlations, and (b) MotorCad.", + "texts": [ + "64 \u2212= \u22c5 \u2212 (8) Using these different correlations, and adopting a water jacket cooling system with 6 L/min flow rate (considered as constant), the main heat transfer coefficient have been calculated for 10,000 rpm and 60,000 rpm and are reported in Table V. For these considered assumptions, the flow is turbulent. The temperature increase has been analyzed using a FEA thermal software Ansys\u00ae (Mechanical). The air-gap and water jacket convective heat transfer coefficients have been calculated using the aforedescribed correlations and reported in Table V. The maximum torque is applied during 20 seconds. The results in Fig. 4.a. shows that the maximum temperature is located in the slots (116 \u00b0C). Another software (MotorCad) has been used in order to validate the proposed approach. The obtained results in Fig. 4.b. confirm the obtained results. The maximum temperature is located in the winding (hot spot in the endwinding at 105 \u00b0C). The maximum temperature in the slots is 100\u00b0C. The difference is mainly due to the conservative equivalent slot thermal conductivity that has been considered. ( ) ( ) ( ) ( ) ( ) 2 3 2 3 Nu 7.49 17.02 H W 22.43 H W 9.94 H W 0.065 D L Re Pr 1 0.04 D L Re Pr = \u2212 \u22c5 + \u22c5 \u2212 \u22c5 \u22c5 \u22c5 \u22c5 + + \u22c5 \u22c5 \u22c5 (5) For the nominal operating point, the same approach has been used. The results obtained by thermal FEA (see Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002466_tvt.2018.2800777-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002466_tvt.2018.2800777-Figure7-1.png", + "caption": "Fig. 7 Force analysis of front and rear wheels of rover vehicle", + "texts": [ + " equations for We, FDPe, and Te are defined as functions of s and \u03b8, and Eqs. (11)-(13) can be written as follows e e 1 DPe DPe 1 e e 1 , , , W W s F F s T T s (14) The four wheels of the manned lunar rover are driven by four independent motors. When the astronaut driving the vehicle accelerates in a straight line on the lunar surface, the front and rear wheels produce the traction force necessary to accelerate. Due to the symmetry of the vehicle, the forces of the front and rear wheels on only one side of the vehicle are shown in Fig. 7. Based on the previously established mechanical model of a single wheel, the vertical load Wf, the net traction force FDPf, and the driving torque Tf of the front wheel can be obtained by Eqs. (14) as follows, f e 1f f DPf DPe 1f f f e 1f f , , , W W s F F s T T s (15) where \u03b81f is the entrance angle of the front wheel; sf is the slip ratio of the front wheel. Similarly, the vertical load Wr, net traction force FDPr, and driving torque Tr of the rear wheel can be obtained by Eqs. (14) as follows, r e 1r r DPr DPe 1r r r e 1r r , , , W W s F F s T T s (16) where \u03b81r is the entrance angle of the rear wheel; sr is the slip ratio of the rear wheel" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001114_045102-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001114_045102-Figure1-1.png", + "caption": "Figure 1. Experimental setups for (a) developing microgrooves on the DR1/PMMA film using a fast-writing holographic method, and (b) measuring the T \u2013V curves of the hybrid LC cells.", + "texts": [ + "06 wt% POSSdoped-PI solution [19] is coated onto the DDP-layer-coated and uncoated substrates. The two coated substrates are then baked at 110 \u25e6C for 30 min to form cured POSS/PI\u2013DDP dual and POSS/PI single layers. The thicknesses of the DDP and POSS/PI layers are approximately 1.4 \u00b5m and 0.06 \u00b5m, respectively. Before combining the dual- and single-layer-coated substrates to form the empty cell, the photosensitive DDP layer on the dual-layer-coated substrate is under the exposure of an interference pattern produced by two coherent pulsed laser beams for inscribing the microgrooves. Figure 1(a) illustrates the schematic holographic two-pulsed-beam writing setup for rapidly inscribing the microgrooves on the DDP film. The two pulsed beams are propagated on the y\u2013z plane and originate from a pulsed Nd : YAG laser beam operated at a wavelength of 532 nm, a pulsed duration of 6 ns and a repetition rate of 10 Hz via a beam splitter. An interference pattern, produced by the overlapping of the two pulse beams with the same s-polarizations and the same intensity at two incident angles of 8", + " The grating vector of the formed microgrooves, G, is along the y-axis and the pitch (spatial periodicity) of the grooves is around 1.8 \u00b5m, obtained based on the following formula associated with two-beam interference: = \u03bb 2 sin \u03b8/2 (1) where \u03bb and \u03b8 are the wavelength and included angle of the interfering beams, respectively. The empty cell is filled with the LCs (E7; \u03b5 = 14.1, \u03b5\u22a5 = 5.2; from Merck), pre-assembled by the two substrates, one with and the other without the formed microgrooves, and a hybrid structure is generated. Figure 1(b) shows the experimental setup for measuring the transmittance versus applied voltage (T \u2013V ) curve of the hybrid LC cell. The probe beam from the He\u2013Ne laser (\u03bb = 632.8 nm) propagates through the hybrid LC cell, which is placed between two crossed polarizers oriented at \u00b145\u25e6 with respect to the grating vector G. The transmittance of the hybrid LC cell is measured by applying an ac voltage of 1 kHz. Each measured T \u2013V curve of the LC cell of different microgroove-writing conditions (tW and IW) is fitted with the simulated result obtained from a commercial software package (LCD master, Shintech)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001529_icphm.2011.6024357-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001529_icphm.2011.6024357-Figure3-1.png", + "caption": "Fig. 3 Planetary gearbox test rig", + "texts": [ + ", ; 1, 2,..., k kjr n r n j jT ki ki b j k i k j i j jk i j ki ki k j jk i j ki ki k C s t x b for k j i n x b for k j j r i n \u03b5 \u03b5 \u03b5 \u03b5 \u03c6 \u03b5 \u03b5 \u03c6 \u03b5 \u03b5 \u2212 = = = = = + + \u22c5 \u2212 \u2264 \u2212 + > = = \u22c5 \u2212 \u2264 + + > = + + = \u2211 \u2211\u2211 \u2211 \u2211w w w w w (10) where j runs over 1, 2, \u2026, r-1 and nk is the number of samples in the rank k. By solving the optimization problem in Eq. (10), the optimal w and bj will be found, and thus a ranking model (Eq. (7)) will be built. III. PLANETARY GERABOX TEST RIG A planetary gearbox test rig shown in Fig. 3 was used to perform fully controlled experiments. It includes a 20 HP drive motor, a bevel gearbox, a two-stage planetary gearbox, two speed-up gearboxes and a 40 HP load motor. The transmission ratio of each gearbox is listed in Table 1. 4/8 Fig. 4 shows the schematic diagram of the two-stage planetary gearbox. The 1st stage sun gear is connected to the driven bevel gear by shaft #1. The 1st stage planet gears are mounted on the 1st stage carrier which is connected to the 2nd stage sun gear by shaft #2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001575_robot.2010.5509764-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001575_robot.2010.5509764-Figure3-1.png", + "caption": "Fig. 3 First idea for an improved actuator", + "texts": [ + " Certain values of durability and of bearable duty cycles have to be reached, 978-1-4244-5040-4/10/$26.00 \u00a92010 IEEE 3254 e.g. most rope actuated robot systems, which use steel cables instead of HMPE ropes for actuation, have certain defined intervals to change their rope components, e.g. 40.000 cycles [12]. To address these challenges some new approaches were necessary. So the first step was to try to build a more compact system. To improve the actuator, mechanical models were needed. The first iterative model combined two DoHelix-shafts that were using only one motor unit, as shown in Fig. 3. This approach reduced system weight, but also reduced efficiency by integration of three gear wheels. Furthermore, the incorrect coiling problem still was not solved with that and so the maximum rotation angle around the axis still was limited. The second approach was to use one shaft to do both DoHelix-coilings (5 and 6) on it, as shown in Fig. 4, A. For that, the motor unit (1) was fixed and a worm gear (2) on the shaft translated the whole pulley-rope-guiding-mechanism on a linear axis (4). This system was much more compact than earlier ones" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002503_1350650117753915-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002503_1350650117753915-Figure3-1.png", + "caption": "Figure 3. Mesh used in the calculation. (a) Oil supply pipe, (b) Oil films and (c) Detailed view of oil films.", + "texts": [ + " The calculation domain is shown in Figure 2(a). The parameters are shown in Table 1. The bearing parameters used in the simulation are exactly the same as with those in the experiment by Hatakenaka and Yanai.1 Given the complicated shape of the fluid domain, the entire domain should be divided to keep the mesh in good quality. In this work, the domain is divided into two parts, namely, oil films and oil supply pipe, as shown in Figure 2(b) and (c). These parts are discretized into a structured mesh as shown in Figure 3. This work employs the Grid Convergence Index (GCI) method for grid independence validation as recommended by American Society of Mechanical Engineers (ASME). This method provides a dependable method to estimate the uncertainty error of the grid. The ring-to-shaft speed ratio is regarded as an important parameter in FRB calculation. The torques on the inner and outer surfaces of the ring are the variables that determine the floating ring rotation speed; thus, they are selected as parameters to be verified" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000112_jphysiol.1965.sp007798-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000112_jphysiol.1965.sp007798-Figure3-1.png", + "caption": "Fig. 3. Effect of folic acid (A-A) and ascorbic acid (I-_) on the histamine excretion (i4g/2 hr \u00b1 S.E.) of guinea-pigs given oral histidine (1000 mg/kg). Histamine excretion of untreated animals given only the histidine load (e_-) is also shown.", + "texts": [ + " The order of potency of the three inhibitors in vivo was similar to that found in vitro when the non-specific enzyme of guinea-pig kidney was used. Doses of az-methyldopa, dopa and 5-HTP alone did not influence the basal histamine excretion by guinea-pigs. Effect ofpre-treatment with ascorbic acid orfolic acid. The total amount of histamine excreted by guinea-pigs given a histidine load was markedly 803 804 W. DA WSON, D. V. MAUDSLEY AND G. B. WEST reduced by pre-treatment with ascorbic or folic acid (see Fig. 3). Although aminoguanidine (100 mg/kg), a potent inhibitor of histaminase, raised the basal histamine excretion after histidine about twofold, it did not influence the percentage reduction shown by ascorbic or folic acid. These two acids, therefore, do not reduce the histamine excretion by modifying histaminase activity. Further, they probably do not alter the activity of the non-specific histidine decarboxylase in vivo, as no inhibition of enzymatic activity was found in vitro. Effect of pre-treatment with folic-acid antagonists" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003183_0954406219843954-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003183_0954406219843954-Figure9-1.png", + "caption": "Figure 9. Schematic construction of a hypoid gear according FZG/Wirth et al.21", + "texts": [ + " For practical reasons, a standardized calculation approach based on the local calculation method presented in \u2018\u2018local calculation method\u2019\u2019 section has been developed. This calculation approach allows an estimation of the scuffing load-carrying capacity at a very early stage of designing a bevel or hypoid gear. The developed simplified calculation method of scuffing load capacity is based on the standard calculation method ISO 1030019 in terms of, e.g. geometry, stresses, and sliding velocities. The standard calculation method is based on a virtual cylindrical geometry,20 which is structured according to the schematic construction in Figure 9 according to FZG/Wirth et al.21 Detailed information regarding the representation of a hypoid gear as a virtual cylindrical gear are available in Annex A of ISO 10300-1.19 FZG/Klein et al.9 considers hypoid-specific influence factors on the scuffing load-carrying capacity, inter alia, by the newly developed approach for the calculation of the friction coefficient. The virtual cylindrical gear is derived from the reference cone of the bevel gear or rather the hypoid gear. The transverse path of contact between pinion and wheel of the virtual cylindrical gear is divided in a number of sections (e" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002138_gt2017-63815-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002138_gt2017-63815-Figure2-1.png", + "caption": "Figure 2: BEARING OIL SHEDDING TEST RIG.", + "texts": [ + " Key results involving high speed image visualization are discussed in the following section. A theoretical approach is discussed to determine critical conditions at which break-up modes transits from one mode to another. Conclusions are drawn in last section. The bearing oil shedding rig utilized in the present experimental investigation is a test rig specifically designed and built to investigate the nature of oil shedding from an aeroengine ball bearing. A cross-sectional view of the rig is shown in Figure 2. The rig is powered by a 49kW direct-drive motor designed to operate through a controlled speed range of 0 to 13,000 rpm. However, the range of shaft speeds considered in the present study is only 0-7000 rpm, limited due to rig vibration issues. This preliminary study systematically characterizes the flow through the bearing assembly over the attainable speed range. Recommissioning of the test rig to enable running up to 13,000 rpm will enable future investigation at more engine representative speeds to take place" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002977_icstcc.2018.8540739-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002977_icstcc.2018.8540739-Figure1-1.png", + "caption": "Fig. 1. Quadcopter Rotor forces and moments", + "texts": [ + " As far as we know, this is the first time this has been applied for the modelling of quadcopter dynamics. The estimation of quadcopter rotor flapping dynamic using RBF neural modelling is investigated in this paper. In Section II, the quadcopter rotor dynamics are described. Section III describes the identification methodology followed. The CFA algorithm for RBF modelling is introduced in Section IV. Section V discusses the results obtained followed by conclusions and future work in Section VI. 978-1-5386-4444-7/18/$31.00 \u00a92018 IEEE 292 The equations of motion of a quadcopter (shown in Figure 1) can be expressed with respect to the body-fixed reference frame [17]: mv\u0307 +m(\u03c9\u0304 \u00d7 v) = F (1) I \u02d9\u0304\u03c9 + (\u03c9\u0304 \u00d7 I\u03c9\u0304) = M (2) where v = [u v w]T and \u03c9\u0304 = [p q r]T are the vehicle velocities and angular rates in the body-fixed frame respectively. F = [X Y Z]T is the vector of external forces on the vehicle center of gravity and M = [L M N ]T is the vector of external moments. I = [Ixx Iyy Izz] are the mass moment of inertia measured using a bifilar pendulum method, m is the measured mass of the vehicle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003397_s00170-019-03894-w-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003397_s00170-019-03894-w-Figure2-1.png", + "caption": "Fig. 2 Definition of the all-purpose end mill", + "texts": [ + " For the common involute helicoid [12], if the tooth surface parameters are (u, \u03b8), the position vector and the normal vector of the tooth flank at the cutting point can be expressed as follows: rg u; \u03b8\u00f0 \u00de \u00bc xg yg zg 1 h i ng u; \u03b8\u00f0 \u00de \u00bc nxg nyg nzg 1\u00bd ( \u00f01\u00de where xg yg zg 2 4 3 5 \u00bc rbcos \u03c30 \u00fe u\u00fe \u03b8\u00f0 \u00de \u00fe rbusin \u03c30 \u00fe u\u00fe \u03b8\u00f0 \u00de rbsin \u03c30 \u00fe u\u00fe \u03b8\u00f0 \u00de\u2212rbucos \u03c30 \u00fe u\u00fe \u03b8\u00f0 \u00de p\u03b8 2 4 3 5, nxg nyg nzg 2 4 3 5 \u00bc prbusin \u03c30 \u00fe u\u00fe \u03b8\u00f0 \u00de prbucos \u03c30 \u00fe u\u00fe \u03b8\u00f0 \u00de r2bu 2 4 3 5, rb is the radius of base circle, \u03c30 is the half angular tooth thickness on the base circle, and p is the screw parameter of helical gear. For the all-purpose end mill, as shown in Fig. 2, note that Ot is the cutter location point, P is the cutting point on the surface of tool, H and L are the coordinate parameters of the cutting point, \u03b1t and \u03c8 represent the slope angle of cutting edge and the rotation angle of the tool rotating around xt-axis, respectively. Hence, the position vector and the normal vector of the endmill at the cutting point can be expressed as follows: rt \u03b1t;\u03c8\u00f0 \u00de \u00bc xt yt zt 1\u00bd nt \u03b1t;\u03c8\u00f0 \u00de \u00bc nxt nyt nzt 1\u00bd \u00f02\u00de where xt yt zt 2 4 3 5 \u00bc H L\u22c5 cos\u03c8 2 L\u22c5 sin\u03c8 2 2 664 3 775, nxt nyt nzt 2 4 3 5 \u00bc sin\u03b1t \u2212cosatcos\u03c8 \u2212cosatsin\u03c8 2 4 3 5, \u03b1t is the slope angle of cutting edge, and \u03c8 is the rotation angle of tool rotating around xt-axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003861_peami.2019.8915294-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003861_peami.2019.8915294-Figure5-1.png", + "caption": "Fig. 5. Magnetic field calculation", + "texts": [ + "3 shows a three-phase rectifier and voltage regulator that controls the excitation winding to stabilize the output voltage of the generator. To solve the problem, the program splits the model into a large number of finite elements. For each element, a system of equations for the local matrix is formed. A global matrix is formed from the local matrices ifor solution of the field problem . The results of splitting the model into finite elements are shown in fig. 4. The results of the calculation of the magnetic field are shown in fig.5 The fig.6\u201310 show the results of the calculation of the main parameters and characteristics The analysis of curves shows that the created model is close to the real generator in terms of calculation accuracy. VI. DISCUSSION The proposed generator has a complex magnetic system, which contains two sources of magnetic field. The ways of closing the magnetic flux from the permanent magnets and the excitation winding have a complex shape. The magnetic system has a large flow of scattering. For the exact calculation of characteristics it is impossible to use the simplified method" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003864_s11012-019-01081-5-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003864_s11012-019-01081-5-Figure5-1.png", + "caption": "Fig. 5 Planar (a) and Scho\u0308nflies (b) motion of the 3-CPCR manipulator", + "texts": [ + " As explained in the previous section, only in configurations where one of the rotational inputs is actuated and the remaining two rotational inputs are kept constant and equal to p=2 does the axis of rotation of the moving platform translate in a direction perpendicular to that axis. This axis is always parallel to one of the coordinate axes. By conveniently selecting the two translational inputs in such a way that they define two translational motions in accordance with the coordinate axes and perpendicular to the axis of rotation, planar motion is achieved. The parallel manipulator that results in this case is the 2-PPCR1-RPCR. The only possibilities are: (a) Actuated inputs a1, a2, and a3; a2 = = p=2, a1 = constant; xx \u00bc 0;xy \u00bc 0;xz \u00bc x; Translation plane XY (see Fig. 5a) (b) Actuated inputs a2, a1, and a3; a1 = = p=2, a2 = constant; xx \u00bc 0;xy \u00bc x;xz \u00bc 0; Translation plane XZ (c) Actuated inputs a3, a1, and a2; a1 = = p=2, a3 = constant; xx \u00bc x;xy \u00bc 0;xz \u00bc 0; Translation plane YZ 5. Displacements of dimension 4: \u2022 Scho\u0308nflies motion {Xe}; this is achieved by actuating all the translational inputs, ai, and combining themwith a unique rotational input, ai. The resulting manipulator is the 2-PPCR1CPCR. This manipulator has three possibilities according to a Scho\u0308nflies motion with an axis of rotation in the direction (s,t,v): (a) Themost general case. Actuated rotation a1; blocking a2 and a3 with any value: s \u00bc 1=\u00f0tan a2\u00de, t \u00bc 1=\u00f0tan a3\u00de, v \u00bc 1 (see Fig. 5b) (b) Actuated rotation a2; a1 blocked with value p=2, and a3 blocked with any value: s \u00bc 0, t \u00bc sin a3, v \u00bc cos a3 (c) Actuated rotation a3; a1 and a2 blocked with value p=2. s = 1, t = 0, v = 0 Therefore, if the axis of rotation is intended to be oriented in the general direction of vector (s, t, v), this will be possible only in case (a). In case (b), the direction of x depends only on the value of a3 and is parallel to the YZ-plane. In case (c), the axis of rotation is always parallel to the X-axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003693_j.aca.2019.09.052-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003693_j.aca.2019.09.052-Figure4-1.png", + "caption": "Fig. 4. Illustration of the \u201cLock-and-Key\u201d insertion structure of the optical detection system. A: Exploded view, showing the photodiode (a), the LED (b), the several PMMA layers that constitute the polymeric structure (c) and the four bolts that secure it (d). B: Reversible positioning of the device, perfectly fitting the insertion port and thus permitting an accurate and reproducible alignment of the optical detection chamber with respect to the LED and the photodiode.", + "texts": [ + " The system consisted of a LED with an emission peak centered at 621 nm (HLMP-EH1A, Avago, Digi-Key Electronics, Thief River Falls, MN, USA) and a Si photodiode with an effective area of 33mm2 (S1337-66BR, Hamamatsu Photonics, Hamamatsu, Japan), both mounted into a compact poly(methyl methacrylate) (PMMA) structure where the microanalyzer was inserted into. The insertion structure was based in a \u201cLock-and-Key\u201d concept [38] for allowing a reproducible positioning of the device with respect to the LED and the photodiode (Fig. 4). The LED and the photodiode were connected to a printed circuit board (PCB) that, in turn, was connected to a data acquisition card (DAQ) (NI USB-6211, National Instruments, Austin, TX, USA). The DAQ was responsible for the modulation of the LED and the acquisition and transference of the detected signal to a PC. A digital lock-in amplification was used for processing the raw data, increasing the signal-to-noise ratio and permitting the operation of the system in ambient light conditions without requiring any physical amplifier" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003469_j.engfailanal.2019.06.024-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003469_j.engfailanal.2019.06.024-Figure2-1.png", + "caption": "Fig. 2. The structure of mechanical power offset device.", + "texts": [ + " The sketch of the principle of wellbore trajectory control tool with a high build-up rate is shown in Fig. 1. The front end of the main shaft is equipped with a cantilever bearing, the lower end is equipped with a self-aligning bearing, and the center is equipped with eccentric mechanism. In order to achieve the build-up rate, the eccentric mechanism rotates to make the main shaft have different degrees of bending, and the axis of main shaft deviates from the original direction, so the bit can reach the deflection angle [3]. The structure of mechanical power offset device is shown in Fig. 2. It is mainly composed of two electromagnetic clutches, two reduction gears, an eccentric mechanism and others. The eccentric mechanism consists of two eccentric rings, and the main shaft power is transmitted to the electromagnetic clutch through the couplings. When the electromagnetic clutch is electrified, it meshes with the deceleration device, and the power is transmitted to the eccentric mechanism after the deceleration, thus realizing the rotation of the inner and outer eccentric rings [5]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001041_s0218127410028252-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001041_s0218127410028252-Figure1-1.png", + "caption": "Fig. 1. (a) Schematic diagram of the rotor-to-stator contact system. (b) The section of the rotor/stator system at the position of disk to show the deflections of the rotor and the stator.", + "texts": [ + " Then, the modes of the coupled linear and nonlinear rotor/stator system are derived analytically. The influences of the friction and the normalized contact stiffness on the nonlinear modes are analyzed based on constrained bifurcation theory in Sec. 3. After that in Sec. 4, the interaction of the linear and the nonlinear modes on the characteristics of the dry friction backward whirl is investigated. Finally, the conclusions of this work are drawn in Sec. 5. The rotor/stator model studied in this paper is shown in Fig. 1. A weightless shaft is supported in ideal bearings having the effective transverse stiffness kr and rotates at an angular speed \u03c9. A rigid disk of mass mr is mounted at the midpoint of the shaft and possesses a mass eccentricity of e. Concentric with the disk is an annular stator of mass ms. The stator is elastically supported by a symmetrical set of springs with isotropic radial stiffness ks. There exists a clearance \u03b4 between the rotor and the stator. During rotor/stator rubbing, a friction force with friction coefficient \u00b5 is applied tangentially at the contact point" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003169_s40516-019-0081-y-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003169_s40516-019-0081-y-Figure2-1.png", + "caption": "Fig. 2 a Typical weld bead profile in laser-arc welding b) Schematic representation of different geometric attributes of weld bead profile", + "texts": [ + " The weld plates of identical size were clamped such that the weld root remained in the same position for all the welds. The list of experiments is given in Table 2. The samples were prepared by cutting the specimen perpendicular to the welding direction. The location of the weld bead used for the analysis was mid of the weld length in order to avoid the initial and the end portion of weld that may have slightly different geometry. The weld sample were ground and polished as per standard metallographic procedure to obtain the weld macrograph. Figure 2 a and b show the actual weld and different weld geometric attributes that define the weld bead profile, respectively. The measured values of different weld attributes are given in Table 3. The weld beads were digitized in polar coordinates (r, \u03b8), as shown in Fig. 3. The radial distance (r) for different polar angle (\u03b8) is modeled using an ANN model and predicted as a function of process parameters, as presented next. The ANN is inspired by the neural system of living beings. The five senses acquire signal from environment and passes them to brain through a network of neurons", + " A schematic representation of the architecture of a feed forward neural network for the present case is shown in Fig. 4, wherein welding speed, wire feed speed, current, voltage, and laser power are taken as the neurons in the input layer and the radial distances at several polar coordinates are considered as the neurons of the output layer. Past literature suggests that simple geometrical or trigonometric functions such as a segment of a circle, parabola, ellipse, cosine or combination of them fit the inner and outer profile of symmetric beads obtained with different arc welding processes. It is clear from Fig. 2a that the weld in the present investigation is asymmetric. It is quite difficult to fit the aforementioned functions in the outer boundary because of asymmetric bead complexity that arises due to the irregularity in shape. While regressing the simple curves on a weld bead profile of laser-arc hybrid, weld bead profiles are found to be fit with three curves (AB, BC, and CA in Fig. 5a), four curves (AB, BC, CD, and DA in Fig. 5b) and some are found to be represented by five complicated curves (AB, BC, CD, DE, and EA in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003910_icems.2019.8922130-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003910_icems.2019.8922130-Figure3-1.png", + "caption": "Fig. 3. Status of Gear Failure.", + "texts": [ + " In this paper, the diagnosis of tooth damage in the gearbox is made by assuming the failure situation. Fig. 2 shows the experimental environment used in this paper [4]. The induction motor of the 200W electric motor is 'K9IP200FH' and gearbox gear ratio of 10: 1 is used. The vibration data were collected by attaching the acceleration sensor to the X-axis of the gearbox and the noise data were collected using an audio device. Data collection was performed using data collected using NI PXI. The fault conditions of the gearbox and gear used are shown in Fig. 3. A state in which one tooth of 36 gear teeth of G1 is broken is selected as the fault state Figure 4 shows the FFT waveform of the vibration data. The FFT data represent different characteristics. In the case of the noise data, it can be seen that another characteristic is shown in Fig 1. In this paper, we implemented a deep learning algorithm using CNN (Convolutional Neural Network). CNN learns data using convolution. Therefore, CNN classifies data characteristics better than other algorithms. Due to the convolution operation, it is a specialized algorithm for image analysis and classification [5~6]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001676_icra.2011.5980547-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001676_icra.2011.5980547-Figure9-1.png", + "caption": "Fig. 9. Modification of the trajectory to manipulate the knob.", + "texts": [ + " When the other finger touches the knob, the robot swings the gripper again in the direction of the longer side of the bounding box (4). If the gripper cannot move anymore while the force sensor indicates large variation in both side of swinging direction, the robot calculates the other contact position. By using these contact position data and the knob size, the robot calculates the tilt angle of the knob and changes its posture of the gripper to follow the angle (5). Finally the robot grasps the knob (6). The figure shows the process of recognizing the knob and grasping it by the minimum step of the gripper operation. Figure 9 shows the manipulation of the knob. The dashed line in Fig. 9(a) shows the ideal hand position, and we assume that the robot hand follows the desired trajectory Hc1-Hc2 to manipulate the knob. But the actual grasping point Hce1 often has a margin of position error from ideal grasping point Hc1. In this case, the robot tries to move its hand from Hce1 to Hce2 on the circular trajectory around Oke. The control force to move the hand generated by whole body of the robot is quite larger than the grasping force. Accordingly, when the center of the hand reaches Hce2, the gripper is forced to open strongly because the position and posture of the hand maintain their states by higher power. At that time, \u03b4gr shown in Fig. 9(b), the angle between Hce2Ok and Hce3Ok is calculated by the same way of the systematic touches shown in Fig. 8. Using error angle \u03b4gr, the robot estimates the actual position of the knob joint Ok, and updates the estimated position of the knob joint, Oke to make it closer to Ok. Then, the robot controls its hand posture to reduce \u03b4gr, and modifies the trajectory to follow the circle around the updated position of the knob joint. In the case shown as Fig. 9(b), the center of the robot hand moves from Hce2 to Hce3 with changing the posture of the hand to make it parallel to the knob. This control are done periodically while manipulating the knob. We here consider a swing door that is often used in an office environment. A behavior of the door is controlled by the door-closer equipped on its top. Since the door-closer is loaded by the spring and the dumper, the humanoid robot receives reactive force from the door when opening it. The restoring torque generated on the door can be described as follows [13]: \u03c4(\u03b8d) = Id\u03b8\u0308d + \u03b7d(\u03b8d){kcfd(\u03b8d) + \u03c4 \u2032 c0}, (\u03b8\u0307d \u2265 0)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003469_j.engfailanal.2019.06.024-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003469_j.engfailanal.2019.06.024-Figure5-1.png", + "caption": "Fig. 5. The coordinate system of eccentric strap.", + "texts": [ + " 2) The bearing is rigid support,ignoringthe deformation. 3) The cantilever bearing has no axial movement or rotation, so it is simplified as a moving pair. 4) The self-aligning bearing is simplified as a hinge constraint and the eccentric ring is equivalent to a displacement load. 5) Ignoring the interaction between bit and formation, and the effects of drilling pressure and torque are not considered. Taking the wellbore center at the cross section of the eccentric ring as the coordinate origin, the coordinate system is shown in Fig. 5. The thick edge of the outer eccentric ring Oout as the starting end, and the displacement vector formed at the end of the wellbore center O point is e1.The thick edge of inner eccentricity ring Oin is the beginning end, the center point of outer eccentricity ring as the end displacement vector is e2, and the inner and outer eccentricity ring composite displacement vector is. When e1 is consistent with the negative direction of x axis, e2 is in the same direction as the x axis, the main shaft axis is coaxial with the shell and the tool is in unbiased state, taken that internal and external ring state as the initial state" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001676_icra.2011.5980547-Figure10-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001676_icra.2011.5980547-Figure10-1.png", + "caption": "Fig. 10. Variation of ZMP when the robot pushes the door.", + "texts": [ + " For the further simplification, it is assumed that the height of the floor is zero, meaning that zzmp = 0, and the height of the center of mass (CoM) is constant, meaning that z\u0308G = 0. Then, the equation of the zero moment point (ZMP) is approximated as follows: xzmp = xG \u2212 zG g x\u0308G (3) yzmp = yG \u2212 zG g y\u0308G (4) where xG and yG are the x and y positions of the CoM, and g is the gravitational constant. Also, it is assumed that the direction of the reactive force is orthogonal to the door plane when the gripper exerts forces on the door as shown in Fig. 10. In the figure, based on the world frame, \u03a3W , pH = [xH , yH , zH ]T represents the position of the end of the gripper, and f = [fx, fy, fz]T represents the reactive force. fz becomes zero assuming that the reactive force is exerted on the direction perpendicular to the door plane. In this case, the ZMP is dependent on the reactive force, and then the (3) and (4) can be expressed as follows: x\u0303zmp = xG \u2212 zG g x\u0308G + zH Mg fx (5) y\u0303zmp = yG \u2212 zG g y\u0308G + zH Mg fy (6) where M is the total mass of the robot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002416_imece2017-71288-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002416_imece2017-71288-Figure8-1.png", + "caption": "Figure 8.(a) Schematic representation of Nishimura air spring with non-linear damper and (b) Bond grapgh model of Nishimura air spring", + "texts": [ + " The expressions for air spring stiffnesses and damping [Presthus, (2002)] are given as 5 Copyright \u00a9 2017 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/16/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use 2 0 1 2 0 2 2 3 0 1 0 0 0 0 ( ) 1 2 e r e b e a e r z t s s b r P K n A V P K n A V dA K P P dz A V C k A A V V (15) where Vb and Vr are the volume of air bag and reservoir after deflection, respectively, i.e. 0b b e s sV V zA z A and 0r r s sV V z A (z and zs are the deflections as shown in Fig.8(a)), kt is the total loss coefficients in the surge-pipe, and 1.8 [Berg, (2000), Mazzola and Berg, (2014)]. Other parameters in Eq. (15) are mentioned in Table 3. As assumed in most other literature, the change of area stiffness is neglected (K3=0) [Presthus, (2002); Mazzola and Berg, (2014)]. The polytropic coefficient of air in the spring is taken to be n=1.32. The air spring parameters mentioned in Table 3 are for constant preload of 80 kN and atmospheric pressure aP =101 kPa. The internal absolute pressure (P0) may be calculated as 0 80,000 a e P P A SYMBOLS Shakti bond graph modeling and simulation software [Mukherjee and Samantaray, (2005)] is used for modeling the full railway vehicle (car body and two bogies) along with bio-mechanical human body" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000010_13506501jet600-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000010_13506501jet600-Figure2-1.png", + "caption": "Fig. 2 Mid-plane cross-section and coordinate system", + "texts": [ + " The bushing of the journal bearing is modelled as a hollow cylinder with inner radius RBi, external radius RBo, and length L. The position and the dimensions of the supply groove are given in Fig. 1. The global coordinate system (xG, yG, zG) is defined in such a way that the applied load is directed towards the negative side of the zG-axis. The shaft is considered misaligned in the (yG, zG) plane. Its position is defined by the tilting angle \u03b3 about the xG-axis. The attitude angle \u03d5J at the mid-plane, the eccentricity e, and the direction of the rotational speed \u03c9J of the journal are shown in Fig. 2. Assuming that the radius of the journal is very large compared to the film thickness, the curvature may be neglected and the journal surface can be developed onto a flat surface as shown in Fig. 2, where (x, y, z) is the coordinate system of the film. The x-axis is located along the circumference of the journal, the y-axis along its width, and the z-axis towards the bushing along the film thickness. The bushing is considered stationary, while the journal rotates at constant angular velocity \u03c9J. Mechanical and thermal deformations as well as roughness and running-in effects are neglected. The film thickness h(\u03d5, y) between the journal and the bushing is obtained by the following equation h(\u03d5, y) = C(y) \u2212 (\u221a e2 zG + e2 xG ) cos[\u03d5 \u2212 atan2(ezG , exG)] (1) In the ideal case, when the journal and bushing axes are parallel, the eccentricity e is constant along the yG-axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000253_978-3-642-00644-9_48-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000253_978-3-642-00644-9_48-Figure5-1.png", + "caption": "Fig. 5 CAD drawings of the pcb\u2019s top and pcb\u2019s bottom", + "texts": [], + "surrounding_texts": [ + "The main software runs on the PC computer and receives information from the robjects and from the high-definition camera fixed over the table. From camera images we extract the location of all glasses and of the table border. Using bluetooth, the PC sends orders to the robots as function of the received information obtained by the camera and from the robots. The software running on the robots receives and executes order received from the PC. Some basic algorithms to do obstacle avoidance, table\u2019s border detection, pre-processing of force sensor data (if someone takes the glass, if the glass is empty, if someone fill up or drain the glass, if the liquid level laze) are embedded in the robots themselves." + ] + }, + { + "image_filename": "designv11_33_0001093_icra.2012.6225142-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001093_icra.2012.6225142-Figure1-1.png", + "caption": "Fig. 1. Concept of the Omnidirectional gear Mechanism", + "texts": [ + ". INTRODUCTION N this research, we propose a new omnidirectional driving gear mechanism that enhances its driving area from the normal X-Y plane to cover convex and concave surfaces, and even various combinations of both[1]-[3]. II. PRINCIPLE OF THE OMNIDIRECTIONAL DRIVING GEAR MECHANISM The concept of the omnidirectional driving gear mechanism proposed in this paper is shown in Fig. 1. It has two gear columns, which are perpendicular to each other, on the flat or curved surface. This plane exerts driving forces on the X and Y axes via the spur gears[4][5]. Kenjiro Tadakuma is with Osaka University, Department of Mechanical Engineering, Room #406, Building-M4, 2-1 Yamadaoka, Suita-shi, Osaka, 565-0871, Japan (phone: +81-6-6879-7333; fax: +81-6-6879-4185; e-mail: tadakuma@mech.eng.osaka-u.ac.jp). Riichiro Tadakuma, Kyohei Ioka, Takeshi Kudo, and Minoru Takagi are with Yamagata University, Faculty of Engineering, 4-3-16 Jonan, Yonezawa city, Yamagata Prefecture, 992-8510, Japan (phone: +81-238-26-3893; fax: +81-238-26-3205; e-mail: tadakuma@yz", + " Makoto Kaneko is with Osaka University, Department of Mechanical Engineering, 2-1 Yamadaoka, Suita-shi, Osaka, 565-0871, Japan (phone: +81-6-6879-7331; fax: +81-6-6879-4185; e-mail: mk@mech.eng.osaka-u.ac.jp). B. Versions of the omnidirectional driving gear mechanism 1) Planar version The planar version of the omnidirectional driving gear mechanism is defined as a gear structure with gear columns that are perpendicular to each other and configured and deployed on the surface with zero curvature (Fig. 1). The whole structure is a flat plane. 2) Convex version The convex version of the omnidirectional driving gear mechanism is defined as a gear structure with two crossed gear columns that are configured and deployed on the outer surface, the curvature of which is positive (Fig. 2). 3) Concave version The concave version of the omnidirectional driving gear mechanism is defined as a gear structure with two crossed gear columns that are configured and deployed on the inner surface, the curvature of which is negative (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002063_s00170-017-0656-8-Figure19-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002063_s00170-017-0656-8-Figure19-1.png", + "caption": "Fig. 19 The finite element model of the cutting force", + "texts": [ + " To increase calculation efficiency and improve the calculation accuracy, simplifications of the cutting tool and the workpiece are adopted and a finer meshing process is used during modeling, and only the outer cutting blade cutting situation within the time range of 15.985 ~ 16.0125s is simulated. Based on the purpose of positioning conveniently, the symmetry of the two outer cutting blades is reserved. In order to reduce the number of the grid, only two gear teeth are reserved and the grid is set as the automatically following pattern; the finite element model of the cutting force is shown in Fig. 19. Tetrahedral mesh is employed in this paper, with an element size ratio of 8 and minimum element size of 0.19 mm; the total grid number is 175,612. A finer mesh window is set as the follow-up mode (moving with the cutting tool). In the processing of setting of the movement condition and the boundary condition, a rotary cutting movement with the velocity of 15.7 rad/sec and a linear feeding movement with the velocity of 30 mm/min are applied on the cutting tool. The initial temperature is set at 20 \u00b0C, the friction is 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003466_j.mechmachtheory.2019.06.017-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003466_j.mechmachtheory.2019.06.017-Figure9-1.png", + "caption": "Fig. 9. Coordinate systems for face gear generation: (a) illustration of installation; (b) derivation of coordinate transformation.", + "texts": [ + " The varying tooth thickness is significant to the backlash adjusting of the face-gear drive. The face-gear is generated by a tapered involute shaper. Tooth surfaces 2 of the face-gear is obtained as an envelope set of the shaper surfaces s as shown in Section 2 . The coordinate systems used for the face-gear generation are established. The limiting conditions including undercutting, pointing, fillet intersection and involute interference of the face-gear are determined. The coordinate systems are shown in Fig. 9 . Coordinate systems S n and S m are fixed relative to the housing of the facegear drive. The moveable coordinate system S 2 is attached to the face-gear, and coordinate system S s is attached to the tapered involute shaper. Point P is the intersection of the coordinate axis O n - y n and the pitch cone of the face-gear. Point Q represents the intersection of the axis O n - y n and the addendum cone of the face-gear. Parameters L 1 and L 2 are the inner and outer radii of the face-gear, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001718_s11431-011-4499-5-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001718_s11431-011-4499-5-Figure4-1.png", + "caption": "Figure 4 The 3-RRRP mechanism.", + "texts": [ + " But at every configuration, is equal to , which also means that is not an independent parameter. So, is 1, not 0. Therefore, link group EF\u2019s movements along x, y-axes are only caused by rotation, the real constraint of link group EF should be gz IIm (, 0 0 z)=4. X IIm = 5,2 Im (, 0 0z)+ gz IIm (, 0 0 z)= X IIm (, 0 0 z)=4, FII=6nII IIP i ip1 X IIm = 6\u00d71 5\u00d72+4=0. Link group EF is an Assur group with X IIm =4 and F=0. For the mechanism in Figure 3, F=FI +FII=1+0=1, or it can be obtained from eq. (10) as F=6\u00d745\u00d76+(3+4)=1. Example 2. Figure 4 is a 11-bar 3-RRRP mechanism introduced by ref. [15]. Loop I is denoted by ABCDEFGH. In loop I, the eight axes are parallel to plane O-xy, there is only one common constraint which constrains the rotation around z-axis. So, X Im = X Im (0 0 , 0 0 0)=1, FI =6nI X I P i i mpI 1 =6\u00d775\u00d78+1=3. In addition, the two rotating pairs D and E in link 4 are always parallel to plane xOy , so link 4 has only a translation, the constraint of the generalized pair 11,4 IG formed by link 4 and the base link 11 is 11,4 Im (, 0 0 0)=3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001986_iet-epa.2016.0565-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001986_iet-epa.2016.0565-Figure1-1.png", + "caption": "Fig. 1 3D models of rotor (a) Straight rotor, (b) Skew rotor, (c) Dual skew rotor", + "texts": [ + " Thus, some stator and rotor slot numbers are forbidden to avoid the synchronous parasitic torque occurred. Stening and Sadarangani [9] and Haisen et al. [10] present an asymmetrical rotor slot distribution (without skewed slots) which the parasitic torques can be suppressed. This paper [11] illustrates that three-dimensional (3D) finite element method for modelling the skew slot rotor is a reasonable solution. Moreover, a simplified 3D modelling for skewed rotor of induction motors is proposed. The structure difference of three kinds of rotors is shown in Fig. 1. Squirrel-cage induction motor with straight rotor should be careful to apply the stator and rotor slot combination for avoiding the synchronous parasitic torque occurred during the low-speed IET Electr. Power Appl., 2017, Vol. 11 Iss. 8, pp. 1357-1365 \u00a9 The Institution of Engineering and Technology 2017 1357 region and standstill. Many squirrel-cage induction motors are commonly manufactured with skew rotors; skew rotors could attenuate the synchronous parasitic torque, but cannot remove it. It seems appropriate to investigate the dual skew further" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002579_s0001925900002341-Figure17-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002579_s0001925900002341-Figure17-1.png", + "caption": "FIGURE 17. Location of slices.", + "texts": [ + " Similar arrangements of pulleys and dead weights were arranged to provide tension and pure bending. The loading rig was attached to a plate fitted in the oven shelf runners, while the model stood on a similar strengthened base plate. Correct alignment of the model and the loading pulleys was preserved by the use of a dowel pin fitted in the base plate and locating the axis of the shouldered shaft. The slicing procedure for each of the three loading conditions was similar to that carried out on the corresponding standard models. Two slices were required from each model. These are shown in Fig. 17. Measurements to determine the maximum stress, tangential to the surface, in the small diameter shaft were made on the \"uniform\" slice, A. The maximum tangential surface stress at the fillet was found by extrapolation of a curve (obtained by plotting measurements of the relative retardation through the second slice, B) along a line normal to the shaft surface. The results of the three \"insert\" model tests are given in Table II and plotted with those of the standard models in Figs. 5, 10, and 14" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000151_iemdc.2009.5075258-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000151_iemdc.2009.5075258-Figure7-1.png", + "caption": "Fig. 7. Flux paths 2D", + "texts": [ + " To ensure the uniqueness of the vector potential, there must also always the tangential or normal component of the vector potential to be defined on the edge of the problem. n \u00b7 B = 0 \u2194 n \u00b7 \u2207 \u00d7 A = 0 on \u0393B (11) n \u00d7 H = 0 \u2194 1 \u03bc \u2207\u00d7 A \u00d7 n = 0 on \u0393H (12) And more simple: n \u00d7 A = 0 on \u0393B (13) n \u00b7 A = 0 on \u0393H (14) Usually in electrical machines, the last two equations form the condition for normal and tangential component of the field (see fig. 5b). If we execute, we receive, however, insufficient results (see fig. 7). We have different possibilities to build the geometry. One is to use the full 360 degree model another to use the smallest periodical part with even or odd periodicity. Even means the same boundaries on both sides of the edge and a phase rotation of 360 degree, odd means inverted boundary conditions and a phase rotation of 180 degree. The problem is, that we need a phase shift of 180 degree left an right to the periodicity. Using the full model this is not possible. We would need two conditions on the same line or point" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003344_s11431-018-9445-3-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003344_s11431-018-9445-3-Figure1-1.png", + "caption": "Figure 1 Kinematic model of autonomous electric vehicle.", + "texts": [ + " Section 2 displays the nonlinear dynamic model of over-actuated autonomous electric vehicle with input saturation. An adaptive neural network-based terminal sliding control law is designed for determining the generalized force/moment in section 3, and a distribution law of tire forces is designed for solving the over-actuated problem in section 4. Finally, section 5 shows the simulation results from an autonomous electric vehicle in order to verify the feasibility of the proposed control scheme. The conclusion is given in section 6. 2 Problem formulation Figure 1 shows the kinematic model of autonomous electric vehicle. xc and yc denote the current longitudinal and lateral coordinates of the vehicle with respect to the inertia frame; xd and yd denote the longitudinal and lateral coordinates of the desired position with respect to the inertia frame; \u03b8c and \u03b8d denote the current and desired orientation of the vehicle centerline with respect to the inertia frame; vd and \u03c9d represent the desired longitudinal velocity and the desired yaw rate, respectively; vx and vy denote the longitudinal and lateral velocities, respectively; \u03c9 is the yaw rate; xe, ye and \u03b8e denote the longitudinal, lateral and orientation errors of autonomous electric vehicle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002288_ipack2017-74173-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002288_ipack2017-74173-Figure1-1.png", + "caption": "FIGURE 1. CAD MODEL OF THE \u03bc-SLS SYSTEM", + "texts": [ + " Unfortunately, current commercially available metal additive manufacturing tools either have feature-size resolutions of greater than 100 \u03bcm, which is too large to precisely produce interconnect structures, or can only produce two dimensional structures for printed electronics applications. To produce micron-sized features on a variety of substrates, a new AM technique called micro-scale Selective laser sintering system has been developed[11,12]. The \u03bc-SLS system is made up of six critical sub-systems as shown in Figure 1: (1) the spreader mechanism used to generate the powder bed, (2) the optical system used to write features into the powder bed, (3) the laser system used to sinter the particles, (4) the stepper system used to scan the optical system across the 1 Copyright \u00a9 2017 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 10/28/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use powder bed, (5) a vacuum chuck used to hold the flexible substrate in place and ensure that it does not deform during the coating/writing/die bonding processes, and (6) a vibration isolation system used to reduce outside influences that could damage the part quality" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000908_0020-7403(65)90021-4-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000908_0020-7403(65)90021-4-Figure2-1.png", + "caption": "FIG. 2.", + "texts": [ + " An inflexionally dis tor ted, initially straight member cannot be par t of an elastica with several loops, see Fig. ld . This follows from equat ion (le). Since o ther possible values o f\u00a2 ( = 0 - f l ) a t which the curva tu re vanishes are - a, - 2~r + a, and since the corresponding points do not lie on the central line of thrust , equi l ibr ium under the g iven condit ions is not possible. S T A T I C A L A N A L Y S I S O F C O N T I N U O U S F R A M E W O R K S Continuous plane f rameworks m a y be divided into two main categories, (i) por ta l - type frames, see Fig. 2, and (ii) braced, t r iangula ted framcs, usually referred to as rigidly Typica l por ta l - type f rame in the init ial ly unbuckled form. jo inted trusses (see Refs. 7 and 13). Both types are cont inuous a t the joints which are assumed rigid. Occasional hinges m a y be present wi thou t a l ter ing the general behaviour . Braced frames are axial ly iso-static if the number of joints J and tha t of the members N are related by N = 2 J - 3 (2) I f N > 2J - 3, the f rameworks are axial ly hypers ta t ic ", + " The degree of sway D corresponds to tha t of the k inemat ic mechan i sm which can be obta ined from the given frame on replacing the rigid connections by pinned connections. In tha t case, N + D = 2JF (3 668 S . J . BRITVEC N is aga in t he n u m b e r of m e m b e r s and JF the n u m b e r of free j o in t s ( joints no t a t t a c h e d to the suppor t s ) . Such f rames are supposed to be loaded cen t ra l ly a t t he jo in t s by a sys t em of conse rva t ive p r o p o r t i o n a t e loads W,.j = I i } , such t h a t sway pr ior to buck l ing is e l imina ted , see Fig. 2. T h e n X~ = W E ~ . cos ~ . (In) i and )~ = W ~ ~ . s i n ct. (4b) J where ~is = ~sl are the c o n s t a n t s of p r o p o r t i o n a l i t y a n d IV is the single load p a r a m e t e r (d imensional ) . T r i a n g u l a t e d f r ameworks are also a s s um ed to be loaded cen t r a l ly a t the jo in t s bu t no res t r ic t ions , o t h e r t h a n those of p ropo r t i ona l i t y , are imposed on the loading sys tem. F o u r pr inc ipa l g roups of cond i t ions c o n s t i t u t e t he basis for t he s ta t i ca l ana lys is of c o n t i n u o u s f r ameworks in equ i l i b r i um ; (i) geomet r ica l c o m p a t i b i l i t y of t he d i s to r t ed conf igura t ion ; (ii) c o n t i n u i t y be tween the m e m b e r s a f t e r buck l ing ; (iii) s t a t i ca l equ i l ib r ium in the u n b u c k l e d a n d buckled conf igura t ions - (iv) genera l ized equ i l ib r ium equa t i ons of single m e m b e r s " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003789_apusncursinrsm.2019.8888860-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003789_apusncursinrsm.2019.8888860-Figure3-1.png", + "caption": "Fig. 3. (a) First mode excited using microstrips (b) Second mode excited using microstrips.", + "texts": [ + " The symmetry of the current distribution forces these two ports to be on opposing sides. The nulls occur at the center of the antenna indicating that the two opposing ports must be fed 180o out of phase. To excite the second mode, four ports are used. The nulls occur between 2037978-1-7281-0692-2/19/$31.00 \u00a92019 IEEE AP-S 2019 adjacent ports indicating that adjacent ports must be 180o out of phase. The modes and their corresponding excitations are summarized in Table I. When the ports are exited in accordance with Table I, the current distributions of the modes are as shown in Fig. 3. The current\u2019s nulls in the first mode appear to \u201cswirl\u201d. Furthermore, there are two, small, new nulls produced at opposing corners of the antenna. These perturbations are due to each feed being slightly offset to the right of the line of symmetry of the ring antenna. This offset is added because the fold line of the origami waterbomb structure would split the microstrip in half if they were centered symmetrically. The current distribution of the second mode appears relatively unaffected by the offset microstrip lines" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001027_1.4001680-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001027_1.4001680-Figure1-1.png", + "caption": "Fig. 1 Segments and joints of the feline hindlimb \u201elateral view of left hind limb\u2026. Joints are in bold font and segments are in italics. The metatarsalphalangeal joint is represented by the abbreviation MTP. The hip angle is defined as the anterior angle between the pelvis and thigh, the knee angle is the posterior angle between the thigh and shank, the ankle angle is the anterior angle between the shank and foot bones, and the MTP angle is defined on the plantar side of the paw and foot, from the plantar aspect of the foot to the plantar side of the digits.", + "texts": [ + " Radio-opaque reference markers were imbedded in the lium, femur, tibia, talus, and phalanges to allow registration beween CT for bone and MR for muscle images. Bone surfaces ere reconstructed from CT Mimics, Materialise, UK and imorted in modeling software SIMM, Musculographics, USA . 2.2 Kinematics. The distal hindlimb skeleton was modeled as 6 degrees-of-freedom DOF mechanical linkage with five rigid egments pelvic, femur, shank, foot, and paw and four joints hip, knee, ankle, and metatarsalphalangeal MTP Fig. 1 . The ip joint was assumed to be a 3 DOF ball and socket and the MTP nd the ankle joints were assumed to be 1 DOF hinge joints. The atellofemoral and tibiofemoral kinematics were specified accordng to experimental measurements and were coupled to the knee oint angle. Briefly, the motion of the tibia and patella were racked three animals from infrared emitting diodes fixed rigidly o each bone OptoTrak, Northern Digital, Waterloo, ON, Canada nd the center of rotation of each bone relative to the femur was etermined for the tibia tibiofemoral joint and patella patelofemoral joint " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000228_b905459f-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000228_b905459f-Figure3-1.png", + "caption": "Fig. 3 Experimental setup for electrical plucking (a) and acoustic excitation (b). The electrode gap is 6 mm. The dashed rectangle sketches the heating box. A millimetre thick glass fiber of 10 cm in length connects the speaker to a thin plate near the filament in the box. It generates plane sound waves. The support and needle can be rotated synchronously about a vertical axis, such that the filament can be inspected from different sides.", + "texts": [ + "10,15 X-Ray data show that the filaments consist of cylindrically wrapped molecular layers.8 In general, filaments of more than 5 mm diameter are composed of bundles of small fibers, which have diameters of the order of a few micrometres.10 The second material (2) has the phase sequence isotropic 180 C B7 (columnar) 120 C crystalline (for chemical composition see Fig. 2). Stable filaments of this material can be drawn in the B7 phase, we assume that their internal structure is similar to filaments of the first material. The experimental setup (Fig. 3(a) and (b)) is equivalent in its essential parts to the one used in previous investigations.9\u201311 Filaments are drawn in a nearly cubic copper heating box with a volume of about 100 cm3 that provides a constant temperature in the range from room temperature to about 200 C. The box contains a support dish which holds a small drop of liquid crystalline material as the reservoir. A needle, attached to a stepper motor outside the box, is moved toward the drop until the tip is in contact with the mesogenic material", + " We will discuss below that deviations of the filament from cylindrical geometry, which have earlier been ignored, represent a major error source in the interpretation of experimental data. Lateral observation windows in the heating box allow one to study the films in transmitted light by means of a long-range microscope (Questar QM 100). A fast camera (Citius Imaging C10) is used to record videos of a small region near the central part of the filament. The viewing field is sketched schematically by the dotted boxes in Fig. 3(a) and (b). Frame rates are between 1000 fps (frames per second) and 5000 fps. The setup in Fig. 3(a) is used for the electric measurements. The electric field is perpendicular to the observation axis, it deflects the filament in the field direction. The maximum deflection amplitude is of the order of 100 mm. This experiment is equivalent to that reported in ref. 9. When a DC electric field perpendicular to the filament axis is applied, the filament deflects, a well defined bend deformation is induced.11 The deflection shape is sinusoidal in good approximation, z(x)\u00bc z0 cos(kx) with the wave number k \u00bc p/L and filament length L", + "5 MV m 1, it is a few percent of L. After the field is switched off, we observe damped oscillations of a spatially sinusoidal ground mode, and retrieve the dependence of both oscillation frequency and damping rate upon filament length, filament diameter and sample temperature. It has been shown that the analysis of the deflection amplitude z0(t) is sufficient for the description of the essential filament dynamics, and that a linear oscillator model can be employed for z0(t).9 The acoustic setup is shown in Fig. 3(b). Sound-absorbing material covers the inner surface of the heating box. Before the acoustic experiments are performed, the two electrodes are removed without twisting or destroying the filament. A small glass plate with a diameter slightly larger than the filament length is positioned near the filament. This plate is connected to a speaker outside the heating box by a glass rod. The speaker is driven with a wave form generator in a frequency range between 100 Hz and 1 kHz. The relationship between electric voltage at the speaker and the amplitude of the plate oscillations is frequency dependent. The camera is adjusted so that the middle part of the excitation plate is recorded together with the filament (cf. dotted box in Fig. 3(b)). This allows a correction of slightly different excitation amplitudes during a frequency scan. The amplitude of the driving plate Aplate is chosen depending on the Soft Matter, 2009, 5, 3120\u20133126 | 3121 Pu bl is he d on 0 2 Ju ly 2 00 9. D ow nl oa de d by U ni ve rs ita t P ol it\u00e8 cn ic a de V al \u00e8n ci a on 2 3/ 10 /2 01 4 17 :5 1: 49 . View Article Online driving frequency fe so that the product (fe Aplate) is constant during frequency sweeps (see section III B). Aplate values are typically in the range between 10 mm and 100 mm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000525_978-3-642-23244-2_54-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000525_978-3-642-23244-2_54-Figure7-1.png", + "caption": "Fig. 7. C", + "texts": [ + " The obtained results h our work for this specific The DELTA type robots dynamics of the end effec minimum weight of the transmissions, stiffness of lash in the joint couplings The selected type of d the maximum output tors development of a highly components and their ma suitably selected. For mos construction material with a composite material on weight/stiffness ratio. For the purpose of ex joint couplings, several v components readily avail solution employing a ball solution is the need to ex between the individual str The second variant of s IGUS company (Fig. 7 le pendicular rotary couplin tion is that mutual prelo higher complexity due to and lower resistance to po l Kinematics 44 lar positions \u2013 Jacobian verification on the left tions and correctness of the calculation algorithm, w alculation with a Jacobi determinant according to L\u00f3pe ave verified sufficient accuracy of the approach used i type of kinematics. with the parallel mechanism are characterized with hig tor (TCP). This comes together with the requirements o moving components, sufficiently sized drives includin the individual components and mainly minimum back ", + " rive (a servo motor with a planetary transmission) wit ion moment of 90 Nm provides a sufficient potential fo dynamic device. In accordance with this criterion, th terial, directly affecting the overall TCP dynamics, wer t of the components, aluminium alloy was chosen as th suitable mechanical properties. The struts were made o the basis of carbon fibres due to its more suitab perimental tests focused on accuracy and durability o ariants of the solution were proposed. They are built o able on the market. Fig. 7 right shows a classic desig joint to realize the joint couplings. A drawback of th ert retention force; in this case it is realized by spring uts. olution is realized with the components produced by th ft). The joint is made in the form of two mutually pe gs ensuring sufficient DOF. An advantage of this solu ading of the rods is not necessary. A drawback lies i higher number of components needed to realize this lin ssible accidents [6]. 9 e z n h n g - h r e e e f le f f n is s e r- n k 450 Based on the described pr is currently being subject of the accuracy and dyna individual components" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003240_s10921-019-0571-z-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003240_s10921-019-0571-z-Figure5-1.png", + "caption": "Fig. 5 Images of simulated bearing faults a defects on outer race, b defect on inner race, c defect on ball", + "texts": [], + "surrounding_texts": [ + "The fault detection experiments were carried out using a test rig which consists of a centrifugal pump as shown in Fig.\u00a03. The test bearing is one of the two support bearings which support rotor of a centrifugal pump. Four sets of bearings (SKF R7 NB 62) of motor pump were considered for the experimental instigations. The motor was operated under a speed of 1200\u00a0rpm. The sound and vibration signals were acquired simultaneously from a test bearing mounted in the experimental setup under the following conditions: (a) healthy bearings (b) inner race fault (c) outer race and (d) rolling element fault. In general, about 90% of total faults are related to inner race flaw, outer race flaw or rolling element flaw, rarely a cage defect also occurs in rolling element bearings [28]. Initially, the healthy bearing (new bearing) was fixed in the test rig. The vibration and sound signals were acquired simultaneously using B&K 4322 accelerometer and B&K 4117 microphone respectively, the accelerometer was mounted in radial direction to the central axis of bearing housing. In the present work, the pits were artificially introduced on inner race, outer race and rolling elements using electrical discharge machining method. The diameter and depth of the cylindrical pit is approximately 0.7\u00a0mm. The specifications and operating conditions of these bearings are given in Table\u00a01. The position of the microphone near to the bearing housing is important in sound measurements. Initially, the sound measurements were taken at various distances and directions; finally, the position was set at a distance of 5.5\u00a0cm near the bearing housing under near field condition, this method was successful and provided promising fault diagnostic information in our previous works [7, 9, 29, 31]. DACTRON FOCUS\u2014 F100 data acquisition system was used to acquire the sound and vibration signals at a sampling frequency of 16.4\u00a0k\u00a0Hz; these signals were amplified using a B&K 2626 signal conditioning amplifier. The time domain averaging method is most commonly widely used in fault detection of rotating machinery which increases the strength of vibration/sound data relative to the noise obscured in the signals. In the present study the vibration and acoustic signals are synchronously averaged to minimize the random noise by considering 16 sets of raw (9)h( ) = T \u222b 0 H( , t)dt 1 3 time data. The sound and vibration signals were acquired from the bearing setup under various operating conditions and stored in a personal computer for further processing. The local faults simulated on the components of rolling element bearing generate a certain fault characteristic frequencies based on their location and size i.e. dimensions of inner and outer races, number of rollers, shaft speed etc. The bearing geometry and fault locations on bearing components are depicted in Figs.\u00a04 and 5 respectively. These defects mimic contact fatigue faults which are quite common in medium/high speed rotating machines used in industrial applications. When the bearing rotates, each type of bearing defect will generate a particular frequency of impact vibrations. The healthy bearing was replaced by defective bearings which consist of inner race, outer race and ball faults. The signals were acquired for all the cases separately, under the same operating conditions. The characteristic bearing defect frequencies of inner race, outer race and balls are calculated by the Eqs.\u00a0(10), (11) and (12) respectively [19]. (10)fo = Nb 2 fr [ 1 \u2212 Bd Pd cos ] Hz (11)fi = Nb 2 fr [ 1 + Bd Pd cos ] Hz 1 3 (12)fb = Pd Bd fr [ 1 \u2212 ( Bd Pd )2 cos2 ] Hz where, Nb is the number of rollers in bearing, fr is rotating frequency of shaft, \u03b1 is contact angle and Bd and Pd are the rolling element diameter and pitch diameter respectively. For this test rig, the shaft frequency is 20\u00a0Hz. Journal of Nondestructive Evaluation (2019) 38:34 34 Page 8 of 23 The characteristic bearing defect frequency values of inner race, outer race and roller elements are 149\u00a0Hz, 84\u00a0Hz and 52\u00a0Hz respectively. The conventional signal processing techniques viz. statistical parameter and fast Fourier transform analyses methods are used to extract fault related features from time wave forms of sound and vibration signals. Further, in order to enhance diagnostic information, advanced signal processing techniques such as envelope and EMD based envelope and statistical parameter analysis methods developed in MATLAB 6.5 are used to assess the incipient bearing faults. Figure\u00a06 shows the signal analysis procedure used to detect incipient faults in rolling element bearing." + ] + }, + { + "image_filename": "designv11_33_0001347_s10846-012-9725-2-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001347_s10846-012-9725-2-Figure5-1.png", + "caption": "Fig. 5 Horizontal force produced by the disbalance of the thrust forces F of each propulsor", + "texts": [ + " Written with respect to the angles [\u03b1 = \u03d5(t1), \u03b2 = \u03d5(t2)] such that 0 < \u03b1 \u2264 \u03b2 \u2264 \u03b1 + 360\u25e6, the average values of Fx and Fy become Fx,av = \u2212F1+F2 2\u03c0 \u03b2\u222b \u03b1 sin(\u03d5)d\u03d5 = F1\u2212F2 2\u03c0 [cos(\u03b2) \u2212 cos(\u03b1)] Fy,av = F1\u2212F2 2\u03c0 \u03b2\u222b \u03b1 cos(\u03d5)d\u03d5 = F1\u2212F2 2\u03c0 [sin(\u03b2) \u2212 sin(\u03b1)] (9) From Eq. 9, the interval [\u03b1, \u03b2] with duration of 180\u25e6 ensures the largest effect on the resultant force, which produces horizontal movement. The resultant force is Fav = \u221a F2 x,av + F2 y,av \u00b7 ei\u03b1shift (10) where \u03b1shift is the angle of movement in x-y plane. As depicted in Fig. 5, to achieve the movement in the \u03b1shift direction, it is necessary to decrease the thrust beneath the value of Fz within the interval [0, 180\u25e6], and increase it over Fz in the interval [180\u25e6, 360\u25e6], with respect to \u03b1shift. Fz stands for an average value of thrust, produced by the propulsors in order to achieve the desired vertical movement. In order to decrease the load on dc motors, the added thrusts which produce the shift in the x-y plane are always positive for motor 1 and negative for motor 2. Hence, the simple responses of the thrusts and angular position, as well as the horizontal velocities are depicted in Figs. 6 and 7, respectively. It is evident from Fig. 6 that the voltage response has two states. Figure 5 shows that the shift direction of \u03b1shift fits the angle of the disbalanced area\u2019s end. A more complex model that includes the influence of T\u03c9p = 0[s], is elaborated in Section 3. 2.3 Gyroscopic Stability The key aspect of gyroscopic stability in the spincopter aircraft is the law of conservation of angular momentum. The law of conservation of angular momentum states that the angular momentum of an object remains constant as long as no external torques, or moments, act on it. The derivations presented herein resembles the similar effect observed in a frisbee [19]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001385_isci.2012.6222678-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001385_isci.2012.6222678-Figure1-1.png", + "caption": "Fig. 1: Instrumentation arrangement of journal bearing [19]", + "texts": [ + " These conditions can cause an increase in bearing wear resulting in the shaft to have lower eccentricity which means shaft centre is close to bearing centre that causes reduction in stiffness, oil pressure or drop in S 978-1-4673-1686-6/12/$26.00 \u00a92012 IEEE 119 oil temperature [18]. In these cases, the oil film will push the rotor to another position in the shaft. This process continues over and over and the shaft keeps getting pushed around within the bearing. This phenomenon is called oil whirl, as shown in fig 1. And this whirl is inherently unstable because it increases centrifugal forces that increase whirl forces. Oil whirl can be minimized or eliminated by changing the oil velocity, lubrication pressure and external pre-loads. Oil whirl instability occurs at 0.42-0.48* rpm. San-Andre\u00b4s and Santiago [20] determined experimentally a journal bearing under high dynamical loading conditions inducing large orbital motion (50% of bearing clearance). The oil whirl frequency approaches the first critical speed of the shaft when the shaft exceeds more than two times its first critical speed, then it creates a resonant condition called oil whip" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003055_tmag.2018.2886434-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003055_tmag.2018.2886434-Figure1-1.png", + "caption": "Fig. 1. Stator end and cooling system of the generator.", + "texts": [ + "org/publications_standards/publications/rights/index.html for more information. the cooling medium from entering the air gap [1]. The rotor magnetic field is excited by SmCo magnets (SM-26U, its limited working temperature is 350 \u00b0C). The magnets have the remanence (Br) temperature coefficient of \u22120.03%K\u22121 and are placed piecewise on the rotor yoke surface. The non-magnetic sleeve (50Mn18Cr5C0.4) is used to protect permanent magnets (PMs). The end of the generator prototype and the cooling system are shown in Fig. 1. In this paper, the direct field circuit coupling method is used to analyze the changes in the electromagnetic field of the generator under different unbalanced levels. During the analysis of the electromagnetic field, the external circuit is shown in Fig. 2, where L A , L B , and LC are the end leakage inductances, RA, RB , and RC are the winding resistances, and R1, R2, and R3 are the load resistances. In order to ensure the accuracy of the electromagnetic analysis of the model, the electromagnetic characteristics test of HSPMG prototype under a balanced load with a different speed is tested" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002934_s10846-018-0935-0-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002934_s10846-018-0935-0-Figure4-1.png", + "caption": "Fig. 4 Dynamic Simulation of TMUBOT in MSC ADAMS Environment", + "texts": [ + " Therefore, the closed loop system will be as follows:\u23a7\u23a8 \u23a9 x\u0307i,1 = xi,2 x\u0307i,2 = fi (q,q\u0307) + bi (qi ) ( 1 b\u0302i (q\u0302) ( \u2212f\u0302i (q,q\u0307) + u\u0304i (t) )) = u\u0304i (t) . (57) To wrap it, for each active joint of the TMUBOT a TESO and a PD controller have been designed according to Eqs. 55, 56 respectively. The considered structure is shown in Fig. 3. Simulation of the plant (TMUBOT) and the proposed algorithm have been performed in MSC ADAMS software and the ground is modeled with the assumed friction coefficient \u03bc = 0.8, in order to avoid slipping (see Fig. 4). According to the inverse kinematics of the robot, the gait planer is designed in such a way to produce the desired angular positions for a normal walking of the robot on the ground. For instance, Fig. 5 shows the desired angles of the joints for a walking locomotion. By proper tracking of the gait planar signals, the physical stability of the robot is guaranteed. The first step is designing the TESO parametersli (t). For this purpose, based on the previous section, the PD eigenvalues \u03c1i (t) are considered as: \u03c1i,j (t) = \u03c1\u0304i,j\u03c9i,ob (t) , i = 1, " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001882_rnc.3789-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001882_rnc.3789-Figure2-1.png", + "caption": "Figure 2. The evolution of the formation with translational and rotational motion over the whole time. [Colour figure can be viewed at wileyonlinelibrary.com]", + "texts": [ + "0/ 150 0 0 1 3 5 ; r2 D 2 4 cos.0/ sin.0/ 150 sin.0/ cos.0/ 150 0 0 1 3 5 ; r3 D 2 4 cos.0/ sin.0/ 200 sin.0/ cos.0/ 0 0 0 1 3 5 : (52) Copyright \u00a9 2017 John Wiley & Sons, Ltd. Int. J. Robust Nonlinear Control (2017) DOI: 10.1002/rnc Besides, the control input of the virtual leader is given arbitrarily as follows: Oul0 D 2 4 0 0:01 5 0:01 0 0 0 0 0 3 5 : (53) In Figure 1, the formation evolution of agents is depicted at four different time stages, and the red one denotes the virtual leader, and Figure 2 shows that the formation has been achieved and moved in the desired formation like a whole rigid body. Figure 3 denotes the evolution of the agent\u2019s configuration and velocity and the relative configuration x; y ; with respect to the virtual leader keep constant after the terminal time tf D 25s, denoting that the formation is achieved at the terminal time, and moving with the desired formation after the terminal time. Figure 4 denotes the evolution of the control input of the agent, which keeps constant after the terminal time" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000753_10407790.2011.630949-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000753_10407790.2011.630949-Figure3-1.png", + "caption": "Figure 3. Section rotations associated with Mindlin-Reissner plate theory.", + "texts": [ + " Using the assumptions of Mindlin-Reissner plate theory and small-displacement bending theory, the displacements of a plate at any arbitrary location (x, y, z) may be obtained as u \u00bc zbx x; y\u00f0 \u00de \u00f01\u00de v \u00bc zby x; y\u00f0 \u00de \u00f02\u00de w \u00bc w x; y\u00f0 \u00de \u00f03\u00de where u, v, and w are the displacement components in the x, y, and z directions respectively, and bx and by denote the rotations of the normal to the plate middle plane in the xz and yz planes, respectively. The notations for rotations, shear forces, and moments are shown in Figure 2 and Figure 3. The bending strain is given by ex ey exy 8< : 9= ; \u00bc z qbx qx qby qy qbx qy \u00fe qby qx 8>< >>: 9>= >>; \u00f04\u00de Figure 2. Moments and shear forces on plate. D ow nl oa de d by [ N ip is si ng U ni ve rs ity ] at 1 9: 46 0 9 O ct ob er 2 01 4 The domain is divided into control volumes or plate elements. The governing equations for an individual control volume are obtained by balancing the forces in the z direction and moments along the x and y axes: X Fz \u00bc I1\u20acw \u00f05\u00de X Mx \u00bc I3 \u20acbx \u00f06\u00de X My \u00bc I3 \u20acby \u00f07\u00de where I1 and I3 are the inertia moments" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000945_ssd.2010.5585533-Figure12-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000945_ssd.2010.5585533-Figure12-1.png", + "caption": "Figure 12. 3-D flux path trough the stator yoke, the first magnetic collector and the magnetic ring.", + "texts": [ + " The meshing of both stator and rotor study domains has been automatically generated in the Ansys-Workbench environment. Figure 9 shows a mesh of the stator yoke and of the two associated magnetic collectors. A mesh of the stator lamination is illustrated in figure 10. Figure 11 shows a mesh of the rotor claws and the as sociated magnetic rings. 3.2.1. Main Flux Paths The flux moves through the magnetic circuit of the CPAES within three dimensional (3D) paths. These are described as follows: \u2022 axially in the stator yoke (see figure 12), \u2022 radially down through the collector and crossing the air gap down to the magnetic ring (see figure 12), \u2022 radially then axially in the magnetic ring (see figure 12), \u2022 axially-radially in the claws (see figure 13), \u2022 radially up in air gap facing the stator laminations (see figure 14), \u2022 radially up in the stator teeth and circumferentially in the stator core back (see figure 15), \u2022 radially down in stator teeth facing the two adjacent claws (see figure 16), \u2022 radially-axially in the adjacent claws (see figure 16), \u2022 axially then radially in the magnetic ring (see figure 17), \u2022 radially up through the air gap and the mag netic collector on the other side of the machine (see figure 17), \u2022 axially in the other side of the stator yoke (see figure 18)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002655_jifs-169558-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002655_jifs-169558-Figure3-1.png", + "caption": "Fig. 3. The constructed dynamic model for a two-stage planetary gearbox.", + "texts": [ + " Based on the above work, a coupled dynamic model for the two-stage PG is ready for simulation, and its virtual prototype is then obtained to generate dynamic responses of the planetary gearbox with and without cracks. To obtain vibration responses of the simulated PG with a varying crack, a virtual prototype model is established by using the software ADAMS. The PG components, rigid ones and flexible ones are separately imported to the software ADAMS, in which the latter was generated by using the software ANSYS. During modeling, the HASTIFF solver and integrator SI2 are chosen to ensure the computational accuracy and avoid excessive impacts. The virtual prototype model is shown in Fig. 3. Previous research with simulations and experiments indicate that vibrations provide valuable diagnostic information and a useful way to investigate its internal dynamic behaviors. It is usually modeled as a cyclo-stationary signal with a fundamental frequency that equals to the gear meshing frequency fm , as well as its sidebands whose space is the characteristic frequency fg of a faulty gear. According to its specification, both of them can be calculated as follows [32]: fm = ZrZs Zr + Zs fr, (1) fg = fm Zs N, (2) where Zr , Zs , and N are numbers of the ring gear teeth, the sun gear teeth, and planet gears, respectively; fr is the rotating frequency of the sun gear" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003466_j.mechmachtheory.2019.06.017-Figure17-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003466_j.mechmachtheory.2019.06.017-Figure17-1.png", + "caption": "Fig. 17. Applied coordinate systems for the worm thread surface generation.", + "texts": [ + " Assuming that the three related surfaces ri , s and w are meshing simultaneously during the generation of pinion surfaces (where w is the tooth surface of the generating worm). Two processes are implemented: (i) worm thread surface generation by a rack-cutter, and (ii) generation of the pinion surface by the worm ( Fig. 16 ). The worm thread surface w is generated by ri of the rack-cutter. Virtual internal meshing is performed between the worm and the rack. The generated surfaces w is in line contact with the rack. The coordinate systems for thread surfaces generation of the worm are established ( Fig. 17 ). The coordinate system S r is attached to the rack-cutter. S w is attached to the worm. Coordinate system S t is an auxiliary system. S c is fixed to the worm. \u03b3 wr is determined by the crossing angle between the axes z r of the rack and z w of the worm. r pw is the pitch radius of the worm. During the process of generation, the worm rotates around axis z w with angle \u03d5w , while the rack cutter translates along the axis x r with S c . Here, \u03d5 w S c = 2 cos \u03b2 m n N w (36) where N w is the threads number of the worm, generally N w = 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001226_s10409-011-0410-7-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001226_s10409-011-0410-7-Figure2-1.png", + "caption": "Fig. 2 The generator model. a The cross-section of the generator; b The Jeffcott rotor", + "texts": [ + " In particular for rotor ovality above approximately 10% of the air-gap if the stator triangularity is less then approximately 20% of the air-gap. The rest of the paper is structured as follows: Sect. 2 presents the mechanical model used in this paper and Sect. 3 presents the derivation of the UMP for an arbitrary disturbed air-gap through the principle of virtual work. These sections are similar to Sects. 2 and 3 in Refs. [3,4]. Section 4 gives the equation of motion and introduces some numerical values and parameters. Section 5 is devoted to the investigation of (3,2)-perturbation and finally Sect. 6 gives discussion and conclusion. Figure 2 shows the geometry of the generator model. The generator is treated as a balanced Jeffcott or Laval rotor having a rigid core with length l0, mass \u03b3 and stiffness k of the generator shaft. The rotor rotates counter-clockwise, at a constant angular speed \u03c9. Point Cs gives the location of the bearings while point Cr is the geometrical center line of the rigid rotor core. The coordinate system has its origin at Cs, r is the rotor radius and s is the stator radius. Let r0 and s0 be the average radius of the rotor and the stator, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000151_iemdc.2009.5075258-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000151_iemdc.2009.5075258-Figure9-1.png", + "caption": "Fig. 9. Flux density within the stator for the detent torque calculations", + "texts": [ + " Both the frequency and shape of the graph are not suitable for optimization. The periodicity is destroyed by the rotation of the boundary conditions within the the 2D model. The only possibility to do a 2D calculation would be to build a separate model for each rotor position. What\u2019s more surprising is that the more realistic reduced 3D model shows also significant deviations. The periodicity is correct, but the amplitude is by half too low. To explain that, the flux density distribution can be used (see fig. 9). There are variations in the distribution and the direction of the B field. The results of the reduced model with realistic air and stator geometry seem to base on too large stray losses. On the left side of fig. 9 we see the reduced model. Depending on the relative position of stator and rotor teeth the flux density is either high or low. According to the by half a pole twisted teeth the flux density is also twisted in the upper stack to the lower stack about a half periodicity. The resulting torque corresponds to our expectations. On the right side, however, the flux path is not as clear as within the ideal model. We can explore high saturations at areas which have not been saturated before. One reason is, that the flux goes in axial direction from the upper or lower part to the other without using the yoke" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000400_978-90-481-8764-5_3-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000400_978-90-481-8764-5_3-Figure2-1.png", + "caption": "Fig. 2 Coordinate frame convention and the forces and moments acting in hover (back view)", + "texts": [ + " The complete nonlinear model is described in the following sections. 3.1 Coordinate Frames, Kinematics and Dynamics As common two coordinate frames, the inertial frame J and the body-fixed frame B are defined for the model. Where as the body-fixed frame is placed in the Center of Gravity (CoG) and moves with the helicopter, the inertial frame is fixed to the original location of the body-fixed frame. The frames are defined as it is common in aviation and are shown in the schematic view of the helicopter in Fig. 2. Corresponding to these definitions the transformation from the inertial frame J to the body-fixed frame B \u23a1 \u23a3 x y z \u23a4 \u23a6 = A BJ \u00b7 \u23a1 \u23a3N E D \u23a4 \u23a6 (1) is given by the transformation matrix A BJ = \u23a1 \u23a3 c\u03b8c\u03c8 c\u03b8s\u03c8 \u2212s\u03b8 \u2212c\u03c6s\u03c8 + s\u03c6s\u03b8c\u03c8 c\u03c6c\u03c8 + s\u03c6s\u03b8s\u03c8 s\u03c6c\u03b8 s\u03b8s\u03c8 + c\u03c6s\u03b8c\u03c8 \u2212s\u03c6c\u03c8 + c\u03c6s\u03b8s\u03c8 c\u03c6c\u03b8 \u23a4 \u23a6 (2) obtained by the application of the Euler angles (c\u03b1 = cos(\u03b1) and s\u03b1 = sin(\u03b1)). Reprinted from the journal 29 It is well known that this transformation matrix is not valid for angular quantities. Therefore the angular velocities p,q, r have to be transformed using the transformation matrix R JB = \u23a1 \u23a2\u23a2\u23a2\u23a3 1 s\u03c6s\u03b8 s\u03b8 s\u03c6s\u03b8 s\u03b8 0 s\u03c6 \u2212s\u03c6 0 s\u03c6 s\u03b8 s\u03c6 s\u03b8 \u23a4 \u23a5\u23a5\u23a5\u23a6 (3) that can be used to obtain a differential equation for the time derivatives of the roll, pitch and yaw angles \u03c6, \u03b8 and \u03c8 :\u23a1 \u23a3 \u03c6\u0307\u03b8\u0307 \u03c8\u0307 \u23a4 \u23a6 = R JB \u00b7 \u23a1 \u23a3 p q r \u23a4 \u23a6 ", + " Using Newtonian Mechanics, the differential equations for the rigid body motion in the body-fixed frame become\u23a1 \u23a3 u\u0307 v\u0307 w\u0307 \u23a4 \u23a6 = 1 m F\u2212 \u23a1 \u23a3 p q r \u23a4 \u23a6\u00d7 \u23a1 \u23a3 u v w \u23a4 \u23a6 (5) 30 Reprinted from the journal and \u23a1 \u23a3 p\u0307 q\u0307 r\u0307 \u23a4 \u23a6 = I\u22121 \u239b \u239dM\u2212 \u23a1 \u23a3 p q r \u23a4 \u23a6\u00d7 I \u23a1 \u23a3 p q r \u23a4 \u23a6 \u239e \u23a0 (6) with the body velocities u, v, w, the system mass m, the body inertia tensor I and the total external force and moment vectors F and M. For simplification, the inertia tensor I has only diagonal elements. Such a simplification is feasible as a result of the symmetric design of muFly. So far the equations of motion are independent of the flying platform and can be found in literature [2]. Now the platform dependent total external force F and moment M have to be defined. 3.2 Forces and Moments In a stable hover position, as shown in Fig. 2, the thrust forces from the two rotors Tup and Tdw equal the gravitational force G caused by the mass of the helicopter and the integrated aerodynamical drag force on the fuselage Whub due to the down wash of the rotors. The moments acting on the helicopter are the two drag torques Qup and Qdw from the counter rotating rotors (incl. stabilizer bar), which, if unbalanced, lead to a yaw motion of the helicopter. In free flight, additional forces and moments result from the aerodynamical drag due to the motion through the air, but since the helicopter is small and reaches only low velocities, those can be neglected" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003833_summa48161.2019.8947532-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003833_summa48161.2019.8947532-Figure1-1.png", + "caption": "Fig. 1. Scheme of LSEMPM based on free-piston engine", + "texts": [ + " Mobile and standart power plants of micro-power engineering (up to 100 kW) based on free-piston internal combustion engines (FPICE), are widely used for producing electricity for ultra-small single consumers, as well as energy sources in hybrid units [1]. In such power plants it is rational to use linear synchronous electrical machines with permanent magnets (LSEMPM), built on a modular principle as an electromechanical energy converters [2]. This will allow one to scale the power installation for a specific task [3]. The operation of an electric machine combined in one housing with an internal combustion engine imposes special restrictions on the temperature, which can reach 150 \u00b0C. The design of this installation is shown in Figure 1. II. ADVANTAGES AND DISADVANTAGES LSEMPM According to the literature, FPICE has the following advantages in comparison with the currently operated diesel engines: -It works on any fuel grades, including fuel oil and crude oil; -It has the ability to work economically at low loads [4]; -The absence of a crank mechanism and a camshaft mechanism increases its reliability and maintainability [5]; - It has a relatively high efficiency (up to 40 %, and for ceramic-based pistons it can reach up to 50 %); - It has good size parameters, specific mass is 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003852_j.mechmachtheory.2019.103628-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003852_j.mechmachtheory.2019.103628-Figure8-1.png", + "caption": "Fig. 8. The interference checking model.", + "texts": [ + " Because the curvature of tooth surface of the FFHHG varies monotonously along the tooth flank, the distance between grinding wheel and the other end of the tooth surface is the greatest when grinding any end of the tooth surface, which is easy to interfere with the opposite tooth surface. Therefore, the interference of non-working side should be checked in the case of grinding the both ends of tooth surface.Taking the concave side as an example, when grinding the heel of the tooth surface with the outside of grinding wheel, the inside of grinding wheel is easy to interfere with the toe of convex side. The situation of grinding the heel of concave side is shown in Fig. 8. Let the position vector of the root point at the heel of concave side in the gear coordinate system be V g rh . The position vector V g wh and axis vector k g wh of the grinding wheel in grinding the heel of the gear concave side can be calculated by the method described in Section 5 . According to the method described in Section 3 , the tooth surface equation and normal vector equation of gear convex side can be obtained. Let them be V g v ( \u03b8, b ) and n g v ( \u03b8, b ) respectively. According to the parameters of the gear blank, the coordinates of two points at the toe of gear teeth on the cross section along the axis direction can be obtained" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001301_1.52410-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001301_1.52410-Figure8-1.png", + "caption": "Fig. 8 Definition of phase arguments for the Kuramoto model-based control methods.", + "texts": [], + "surrounding_texts": [ + "DFC is one of chaos control methods and has been widely applied to many applications. This control method is effective in stabilizing chaotic motions to a periodic motion, but needs to record the past data. Recording past data is undesired for space systems, because memory resources are limited in such systems. To overcome this problem, we propose Kuramoto model-based control methods to synchronize the librational motion of CEDTS. The Kuramoto model is a mathematical model that can be used to describe synchronized behavior of a large set of coupled oscillators, such as chemical and biological oscillators, and many other applications. This model usually makes the following three assumptions: 1) there is weak coupling; 2) the oscillators are identical or nearly identical, and 3) interactions depend on the phase difference between each pair of objects. The most popular form of the Kuramoto model has the following governing equations: d i dt i K N XN j 1;j\u2260i sin j i ; i 1; ; N (39) In this model, the phase difference is used to give the feedback to the systems. If the force among the systems is attractive, the phase difference among the systems becomes zero gradually, that is, the D ow nl oa de d by U N IV E R SI T Y O F O K L A H O M A o n Ja nu ar y 24 , 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /1 .5 24 10 oscillation of all the systems is gradually synchronized. On the other hand, the force among the systems is repulsive, the phase difference among the systems becomes the same, that is, the oscillation of all the systems gradually converges to the one with a bias phase difference between each other. To apply the Kuramoto model to the librational control of the CEDTS, we need to consider which component of the CEDTS corresponds to the phase argument i in the Kuramoto model. In addition, to design good controllers, we should take into consideration that DFC for librational motion of TSS presented in the past papers showed good ability to stabilize chaotic motions to a periodic motion. The DFC methods applied to the in-plane librational control of EDT are often expressed as I Kdfc _ t _ t It should be noted that librational angular velocity was used as the control signal to determine the control input of the DFC in past studies, while the control signal in the Kuramoto model is the difference between the phase arguments of system\u2019s motion. Although there is a difference between the above two control methods, if the angular velocity difference in a delayed feedback control is interpreted as a kind of phase difference corresponding to that of the Kuramoto model, the above two controllers become very similar. It is natural that the chaos controllers have similarity because their objective is basically the same, that is, to stabilize chaotic motions to a periodic motion. To consider the periodic librational motion as a simple oscillation, the phase arguments 1j and 2j, which will be used in the Kuramoto model-based controller, are defined using max, _ max, max, _ max, j, j, _ j, and _ j, as 1j : atan2 _ j= _ max; j= max (40) 2j : atan2 _ j= _ max; j= max (41) as shown in Figs. 8a and 8b. In this paper, we consider a triangle formation flight of EDT systems, that is, the number of the EDT systemsN is set to three. By taking into account that the librational motion of each tethered system has a bias phase argument of 2 =N, the following two Kuramotomodel-based control methods are considered to determine the time-varying electrodynamic parameter \"i for each electrodynamic tether which is intended to keep the phase difference of inplane/out-of-plane librational motion between each electrodynamic tethered system to 2 =3. D ow nl oa de d by U N IV E R SI T Y O F O K L A H O M A o n Ja nu ar y 24 , 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /1 .5 24 10 Kuramoto Model-Based Control-1 \"1 Kfsin 12 11 sin 13 11 g \"2 Kfsin 11 12 sin 13 12 g \"3 Kfsin 11 13 sin 12 13 g (42) Kuramoto Model-Based Control-2 \"1 Kfsin 22 21 sin 23 21 g \"2 Kfsin 21 22 sin 23 22 g \"3 Kfsin 21 23 sin 22 23 g (43) where is the bias phase argument, which should be set to 2 =3 for the case of the triangle formation flight. Note that the magnitude of the control input determined by the Kuramoto model-based control is always limited within 2K, because it is given by K sin sin ." + ] + }, + { + "image_filename": "designv11_33_0000882_srin.201000032-Figure11-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000882_srin.201000032-Figure11-1.png", + "caption": "Figure 11. Damage of plunger and 65mm shot sleeve: arrows indicate thehot spotswheresteel-fragmentswerecarriedwith theplunger.", + "texts": [ + " As previously mentioned, the system shot sleeve/plunger failed. The reason for this was the fact that the plunger did not translate the billet completely into the shot sleeve, and a small part remained in the cartridge. The water-cooled CuCoBe-piston moved over the still semi-solid metal. Due to the resulting rapid solidification, very hard martensite fragments are produced. During the plunger\u2019s reciprocating movement, these fragments were embedded into the plunger\u2019s surface (indicated by big arrows in Figure 11) thus damaging the plunger. Even by replacing the relatively www.steelresearch-journal.com 2010 W soft CuCoBe with steel (Figure 11, at the bottom left hand corner) could not prevent this damage. This problem must be resolved before the experiments can be continued. Regarding reproduction accuracy, surface appearance as well as mechanical properties, the quality of the cast parts is very good. Unfortunately, the fracture toughness is too low for the use as hand tool material. Hence an optimization of the tempering process (solution annealing coupled with hardening to dissolve the carbides) is planned to improve the elongation and the fracture toughness" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001755_ijnsns.2010.11.3.211-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001755_ijnsns.2010.11.3.211-Figure3-1.png", + "caption": "Fig. 3: Geometrical representation of the control of the radial position of the impact limiters of the mass particle elongations. Programmed control of the impact limiters positions by positions in radial directions \"on\" or \"off'", + "texts": [ + " Brought to you by | provisional account Unauthenticated Download Date | 6/5/15 2:11 PM In the beginning, it is necessary that initial total mechanical energy is not small and it is required to be enough for progressive multi rotational motion of the heavy mass particle along the rough circle and that the initial angular coordinate and the initial angular velocity satisfy the following corresponding necessary conditions: or <\u03b4 and 0 up to the left side of the right impact limiter position: / = 1,2,3,...,\u00ab, 3.3. Initial conditions of the phase trajectory branch between impacts and alternations of the friction force directions. The first impact appears at the moment: f = /\u201e,,_, to the right side of the right impact limiter of the angular elongations at angular position ) and total mechanical energy (\u0395,\u03c6) of the heavy mass particle vibro-impact motion along a rough circle with two radially moving impact limiters and with both side limited angular elongations, caused by programmed control of the impact limiters positions \"on\" or \"off' presented in Figure 3., are presented. Both graphical presentations are obtained for the case that coefficient of dray Coulomb-type friction is \u03bc = 0.05. Subfigures b* in both Figures 7 and 8 are corresponding details of the main graphical presentation in a* and they correspond to the second phase of the different kind of the motion when heavy material particle motion is one side vibro-impact motion with impacts only to the left side of the right impact limiter in position \"on\". Even the fact that second impact limiter is present in the position \"on\" during this phase of the mass particle motion does not impact into this left impact limiter of the system elongations in the case that \u03b4\\ \u00bb \u03b4", + " a graphical presentation of the power of the friction force (\u03a1\u03bc,\u03c8) work during the heavy mass particle vibro-impact motion path along the rough circle with two radially moving limiters and with both sides limited Brought to you by | provisional account Unauthenticated Download Date | 6/5/15 2:11 PM 222 K (Stevanovio) Hedrih, V. RaiCevio & S. Jovio: Vibro-impact of heavy mass particle moving along rough circle angular elongations, caused by programmed control of the impact limiters positions \"on\" or \"off presented in Figure 3. and for the case that coefficient of Dray Coulomb-type friction \u03bc = 0.05. b* is detail of the main graphical presentation in a* and correspond to the second phase of the different kind of the motion when heavy material particle motion is one side vibro-impact motion with impacts only to the left side of the right impact limiter in position \"on\" and that is > \u03b4. material particle vibro-impact motion path along the rough circle with two radially moving limiters and with both sides limited angular elongations, caused by programmed control of the impact limiters positions \"on\" or \"off' presented in Figure 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000758_s11465-012-0317-4-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000758_s11465-012-0317-4-Figure5-1.png", + "caption": "Fig. 5 Spatial atlas of norm\u00f0\u03c9\u00de when \u03b23 \u00bc 72\u2218", + "texts": [], + "surrounding_texts": [ + "As we known that the SNU 3-UPU parallel mechanism prototype will always exhibit the parasitic rotations in practice [17]. Now we will use the value of norm\u00f0\u03c9\u00de to find the bound of mechanism instability, which indicates there, will appear the parasitic rotation when the number exceeds the bound. For SNU 3-UPU parallel mechanism, the first revolute axes of universal joint attached to the fixed base intersecting at one common point. And the same is true for the last revolute axes of universal joint attached to the moving platform, as shown in Fig. 7(a). Take one limb as an example, the direction of the constraint couple applied on the moving platform is sketched in Fig. 7(b). It\u2019s easy to find that the angle \u03b2i from the direction of constraint couple to the moving platform keeps in a relative stable area. It is around \u03c0=2, and we express the angle \u03b2i as \u03b2i \u00bc \u00bd\u03c0=2 . Of course, the mechanism will be singularity when all the three constraint couples are parallel with each other. However, since the existing clearances, these three constraint couples will never truly be parallel with each other in practice. Since the value \u03b2i \u00bc \u00bd\u03c0=2 , we choose a values set of \u03b23, (80\u03c0=180, 85\u03c0=180, 87\u03c0=180, 89\u03c0=180), to get the contour atlas of norm\u00f0\u03c9\u00de, as shown in Fig. 8. Through observation of the above contour atlas of norm\u00f0\u03c9\u00de, and the obtained practical region of angle \u03b2i, \u03b2i \u00bc \u00bd\u03c0=2 , we conclude that the norm of parasitic rotation, norm\u00f0\u03c9\u00de, of SNU parallel mechanism is located in the area of around value 1.2. However, practically the manufactured hardware prototype of SNU parallel mechanism really revealed an unexpected parasitic rotation due to the joint clearance. When all the prismatic joints are locked, the mechanism behaves as if it has additional rotational degree of freedom. Hence we call the number 1.2 is the bound of instability. The parallel mechanism combined with three UPU limbs will be instability when the norm of parasitic rotation exceeds 1.2 with limited clearance \u03b5. Namely, if we still want to achieve the pure translation without parasitic rotation through the parallel mechanism with three UPU limbs, we should study the following two aspects further. One is obtaining three relative stable parameters \u03b2i by placing the universal joint at the proper position and orientation. These three parameters can make the value of norm\u00f0\u03c9\u00de be lower. Another is decreasing the possible clearance to the greatest extent." + ] + }, + { + "image_filename": "designv11_33_0003000_ece.2018.8554979-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003000_ece.2018.8554979-Figure2-1.png", + "caption": "Fig. 2. BLDC Motor [5]", + "texts": [ + " Thus the fusion of BLDC motors in ceiling fans will have a quantifiable impact on the overall domestic energy demand and promote energy efficiency culture on domestic scale. The paper introduces a five phase BLDC Energy Efficient Ceiling Fan as a substitute for orthodox single phase induction motor and three phase BLDC fans for improved performance and low power consumption [3]. Moreover the paper also discusses various technical details of five phase BLDC motor and energy saving potential that can be obtained by incorporating five phase Energy Efficient Ceiling Fans. field excitation is not required. Internal structure of BLDC is shown in Fig 2. The characteristics of PM BLDC motors are similar to shunt motors except for speed torque characteristics that show more linearity and predictability [5]. BLDC motors have similar working principle as conventional DC motors in which an energized conductor placed inside magnetic intensity is impressed by a an electromotive force. The direction of this force can be calculated from Fleming\u2019s Left hand rule [6]. The force on the conductor on armature windings tends to produce a torque resulting a rotational force" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001087_icca.2011.6138050-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001087_icca.2011.6138050-Figure1-1.png", + "caption": "Fig. 1. Kinematic chain of the MAMR\u2019s prototype.", + "texts": [ + " This kind of robot has similarities, in its kinematic chain, with multi-articulated vehicles which are widely used for load transportation in roads, seaports, airports and storage yards. In the kinematic chain of a MAMR, the configuration variables are the angles, \u03b8i, between the successive trailers, and the steering of the front wheels is characterized by the angle \u03b3. The amount of these variables defines the degree of freedom, DoF, of the system (or DoF = n+1, for constant speed, where n is the length of the configuration vector). Figure 1 shows the chain of the MAMR\u2019s prototype used in this work, with three degrees of freedom, and its geometrical parameters (Ai, Bi), where A1 is the distance between the axles of the truck, Ai (i>1), the distance between the front hitch and the axle and Bi, the distance between the rear hitch and the axle. The greatest complexity for maneuvers and navigation control of a MAMR is related to backward movements. In this situation the system behaves as an inverted multiple horizontal pendulum. The angles between the trailers (at one or more links of the composition) may increase, for any value of steering angle, assuming improper values, from which any action becomes ineffective to control the vehicle", + " The corresponding critical angles, for a given maximum value of the steering angle, can be calculated recursively, in the direct sense of the chain, as next equations. For \u03b81 For \u03b82: The MAMR's prototype, used in this work, has a severe restriction on its steering angle ( 20\u00b0). This restriction could be relaxed by modification of the mechanical hardware, which would allow a greater flexibility and maneuvering. The Table II illustrates the calculation of the critical angles, for the MAMR with geometrical parameters illustrated in Fig. 1 and for different limits of the steering angle. In other words, the table shows the geometrical limits of the configuration angles. These angles have also a physical limit, inherent to the MAMR, that is, in general, much smaller than 90\u00b0. A physical restriction about to 90\u00b0 will be considered for the prototype for analysis issues. The Table II shows some results for rear traction, according to the expressions presented earlier, and some results for front traction, according to the expressions formulated in [7]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002485_aab928-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002485_aab928-Figure3-1.png", + "caption": "Figure 3. Section view of the StarJet printhead with a stainless steel outer shell, ceramic reservoir, ROT unit, chip holder, and nozzle chip (left). Closeup of the ROT unit with the nozzle chip (right).", + "texts": [ + " Two nickelchrome/nickel (NiCr\u2013Ni) thermocouples monitor the temperature of the heater and the nozzle chip. Figure\u00a0 2 depicts a computeraided design (CAD) sketch of the printhead with its components (top) and assembly (bottom). In the first step, the ceramic reservoir is inserted into the main body. The chip holder, reservoir outlet tube (ROT) unit, and chip are assembled and attached to the ceramic reservoir. Afterward, the heater, isolation, and rinse gas adapters are added. After the printhead is filled with raw Al material, the cap is closed. Figure 3 shows a more detailed sectional view of the printhead (see figure\u00a0 2), revealing details of the connection between the printhead body and nozzle chip. The ROT con nects the liquid metal to the inner part of the nozzle chip. This part of the printhead is prone to clogging by impurities con tained in the metal melt, and has to be carefully matched in diameter to the nozzle dimensions for optimal printing results. The ROT is therefore implemented as an exchangeable part. The ROT acts as a small capillary counteracting the hydro static pressure by the molten metal because of the nonwetting contact angles of the metal melt on the ceramic surface of the ROT" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003466_j.mechmachtheory.2019.06.017-Figure16-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003466_j.mechmachtheory.2019.06.017-Figure16-1.png", + "caption": "Fig. 16. Generation of the double-crowned pinion by a generating worm.", + "texts": [ + " This section covers the doublecrowned tapered involute pinion generation by a generating worm. As shown in Fig. 15 , a tapered involute pinion with tooth surface including profile crowning and longitudinal crowning is presented. Assuming that the three related surfaces ri , s and w are meshing simultaneously during the generation of pinion surfaces (where w is the tooth surface of the generating worm). Two processes are implemented: (i) worm thread surface generation by a rack-cutter, and (ii) generation of the pinion surface by the worm ( Fig. 16 ). The worm thread surface w is generated by ri of the rack-cutter. Virtual internal meshing is performed between the worm and the rack. The generated surfaces w is in line contact with the rack. The coordinate systems for thread surfaces generation of the worm are established ( Fig. 17 ). The coordinate system S r is attached to the rack-cutter. S w is attached to the worm. Coordinate system S t is an auxiliary system. S c is fixed to the worm. \u03b3 wr is determined by the crossing angle between the axes z r of the rack and z w of the worm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001537_transducers.2011.5969179-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001537_transducers.2011.5969179-Figure4-1.png", + "caption": "Figure 4: a) Arrangement of the sensors with respect to the axis of the rolling body and PCB attached to washer. The sensor elements are tilted 0\u00b0, 22.5\u00b0 and 45\u00b0 with respect to the axis of the rolling element. b) Sensor elements (22.5\u00b0 and 45\u00b0) and protected wire bonds.", + "texts": [ + " The same parameters as in the first layers are used (Fig. 2f). A PCB was designed that provides a Wheatstone bridge for each of the ten sensors of one washer. The signal is then amplified by an instrumental amplifier. The PCB is glued to the washer. Electrical contact with the sensor is established by first removing the Al2O3 in the desired bonding areas using strong ultrasound of a wire bonder tip. Subsequently, the Ti sensor is wire-bonded to the PCB and the bonds are coated and protected by a resin. The result is shown in figure 4. To analyze the response signal of the sensor at high loads a force measuring machine was used. It consists of a steel frame that holds a hydraulic cylinder. The pressure can be controlled by a voltage signal and thus the force can be adjusted between 0 and 6.5 kN. A load cell A.S.T. KAF-S that measures up to 10 kN at 0.2 % error is attached to the hydraulic cylinder. The load cell is read out by an A.S.T. AE 903 read-out unit at 320 samples per second. The sensor-equipped washer is placed in the frame below the load cell" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002902_00423114.2018.1531135-Figure21-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002902_00423114.2018.1531135-Figure21-1.png", + "caption": "Figure 21. MacPherson suspensionmechanism: (a) the schematic diagramand (b) the kinematicmodel.", + "texts": [ + " Applications of screw axis theory to the behaviour analysis of full-vehicle models and the synthesis of suspension system, PhD Thesis, Korea University, Korea 2007. [21] Lin JS, Kanellakopoulos I. Nonlinear design of active suspensions, 34th IEEE Conference on Decision and Control, New Orleans, LA, 1997. [22] Kumar MS, Vijayarangan S. Analytical and experimental studies on active suspension system of light passenger vehicle to improve ride comfort. Mechanika. 2007;3(65):34\u201341. Appendix Kinestatic analysis of MacPherson suspensionmechanism In Figure 21(a), the wheel is connected to the vehicle body by one SP-serial (N3N2) kinematic chain of a shock absorber, one SS-serial (N4N5) and one RS-serial (N7N6) kinematic chains. Since the wheel can be viewed as a moving platform of the 1-DOF spatial parallel mechanism, it has five constraint wrenches s\u0302i (for i =1, . . ., 5) acting on the wheel, as shown in Figure 21(b). In particular, the constraint wrenches s\u03021 and s\u03022 pass through the S-jointN3 normal to the P-jointN2, respectively. The instantaneous twist of the wheel with respect to the vehicle body can be represented by a unit twist S\u0302H that is reciprocal to the five constraint wrenches s\u0302i. The forces of the spring and the damper act along the line vector r\u0302s passing throughN1 andN3 (see Figure 21(b)). The kinestatic relations of the MacPherson suspension mechanism can be obtained from Equations (8) to (12). Displacement analysis ofmacpherson suspensionmechanism Referring to Figure 22, the lengths li (i =1, 2, 3) of three rigid links N4N5, N6N71 and N6N72 are constant and three points N1, N3 and N8 keep in a same line. Five constraint equations can be Figure 22. MacPherson suspension mechanism. Figure 23. Kinestatic relations of MacPherson suspension mechanism. found as [(x4 \u2212 x5)2 + (y4 \u2212 y5)2 + (z4 \u2212 z5)2] 1 2 = l1, [(x6 \u2212 x71)2 + (y6 \u2212 y71)2 + (z6 \u2212 z71)2] 1 2 = l2, [(x6 \u2212 x72)2 + (y6 \u2212 y72)2 + (z6 \u2212 z72)2] 1 2 = l3, (x1 \u2212 x3)(y8 \u2212 y3) = (y1 \u2212 x3)(x8 \u2212 x3), (x1 \u2212 x3)(z8 \u2212 z3) = (z1 \u2212 z3)(x8 \u2212 x3)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001967_jmech.2017.23-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001967_jmech.2017.23-Figure4-1.png", + "caption": "Fig. 4 A two-stage gear dynamic model.", + "texts": [ + " 3 a schematic torsional model of a dry clutch in an engaged mode is presented where I1 represents the torsional flywheel and cover inertia, I2 is the inertia of the pressure plate, I3 is the friction disk inertia and I4 is the clutch hub inertia. The flywheel and the cover are connected to the pressing plate by a diaphragm spring, which is modeled by a torsional stiffness k12. The connection between the friction disk and the flywheel on one side and pressure plate on the other are respectively modeled by a constant torsional stiffness k13 and k23. The torsional stiffness K1 presents the liaison between the friction disk and clutch hub. 2.1.3 Gearbox Modeling Figure 4 displays a torsional dynamic model of the third block of the drivetrain system (gearbox). The transmission mechanism is achieved by two helical gear stages. I12, I21, I22 and I31 are respectively the inertia of gear 12, 21, 22 and 31. The toothed wheels are assumed to be a rigid body and the bearings are modeled https:/www.cambridge.org/core/terms. https://doi.org/10.1017/jmech.2017.23 Downloaded from https:/www.cambridge.org/core. Eastern Michigan University Library, on 16 May 2017 at 08:50:26, subject to the Cambridge Core terms of use, available at Journal of Mechanics 3 by the linear springs kxi , kyi and kzi of the i-th bearing (i = 1, 2, 3)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003948_iecon.2019.8927305-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003948_iecon.2019.8927305-Figure3-1.png", + "caption": "Fig. 3. Examined model geometries in modal analysis and harmonic response.", + "texts": [ + " The study is performed on a four-pole, aluminum squirrel-cage induction motor having Q1 = 48 stator slots and Q2 = 40 rotor slots. Rated parameters of the studied machine are listed in Table I. Taking rated machine parameters into account gives frequencies 2L = 100 Hertz and FQR = 949 Hertz. Different geometries are used in models for magnetic field and structural analyses. Magnetic transient analysis is performed using 2D geometry of the motor magnetic circuit cross-section and it is identical for all modeled cases. Structural analyses are carried out using 3D models in a three different arrangements (see Fig. 3), further denoted as Case A, Case B and Case C. The Case A exhibit model of a solely stator lamination where stator yoke and teeth are modeled as a separate objects fixed together. In the Case B, model of stator lamination is assembled into a machine frame. The frame is made of 5 millimeters thick aluminum shell with 72 cooling ribs around its circumference and it is enclosed with aluminum end-shields. Finally, in the Case C, geometry studied in the Case B is fastened to the vertical mounting and a mass point modeling the mass of a loading machine is defined on the opposite side of the mounting" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000041_cae.20391-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000041_cae.20391-Figure7-1.png", + "caption": "Figure 7 Movements of the transfer cylinder within the MTMPS.", + "texts": [ + " Transfer cylinder retrieves the processed parts from loading/retrieving position and locates to the sorting magazine according to the material type as metal and plastic. Turntable and processing modules picture is shown in Figure 6. Marking operation was modeled by a single acting pneumatic cylinder. Drilling operation was simulated with a cylinder and an electrical motor and cleaning process was simulated by a nozzle expelling compressed air to the materials. Movements and position names are given in Figure 7. Transfer times of the materials in the MTMPS unit with respect to positions are given in Table 1. Since the plastics part is required drilling operation, the average total time for plastic part is higher than metal average total time. After completion of the processes, parts were stored according to the material type as metal or plastic. Apart from the above-mentioned four main stations, this system has an additional feature. If the process is reset by the user during material handling, transfer cylinder locates the parts in a magazine" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002598_j.ymssp.2018.04.033-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002598_j.ymssp.2018.04.033-Figure5-1.png", + "caption": "Fig. 5. Vibration system split into subsystems.", + "texts": [ + " With the angle au \u00bc X t \u00feuu the vertical and the horizontal unbalance force, which act at the rotor mass, can be described as follows: f uz \u00bc f\u0302 u cos\u00f0Xt \u00feuu\u00de; f uy \u00bc f\u0302 u sin\u00f0Xt \u00feuu\u00de; with f\u0302 u \u00bc e\u0302u mw X2 \u00f01\u00de The dynamic magnetic eccentricity e\u0302m creates together with the electromagnetic stiffness cm a magnetic force f\u0302m which is also rotating with the rotor angular frequency X, acting at the rotor mass and in opposite direction at the stator mass. With the angle am \u00bc X t \u00feum, the vertical and the horizontal magnetic force can be calculated as follows: fmz \u00bc f\u0302m cos\u00f0Xt \u00feum\u00de; fmy \u00bc f\u0302m sin\u00f0Xt \u00feum\u00de; with f\u0302m \u00bc e\u0302m cm \u00f02\u00de The vibration system is split into subsystems (Fig. 5) \u2013 rotor system, journal system, bearing housing system, stator system, foundation system (left side) and foundation system (right side). The displacements of the stator mass zs, ys, us are much smaller compared to the dimensions of the motor h, b, W, therefore following linearization is possible [32]: zaL \u00bc zs us b; zaR \u00bc zs \u00feus b; yaL \u00bc yaR \u00bc ys us h \u00f03\u00de The derivation of these kinematic constraints is described in detail in [10]. Now, a linear inhomogeneous differential equation system can be derived, described by mass matrix M, damping matrix D, stiffness matrix C, coordinate vector q, and vector for excitation fe and vector for actuator forces fa" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003300_cyber.2018.8688228-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003300_cyber.2018.8688228-Figure1-1.png", + "caption": "Fig. 1. Two configurations for twin-lift rotorcraft systems (The first one: with support bar; The second one: without spreader bar)", + "texts": [ + "00 \u00a9 2018 IEEE 913 To give details about such development, the rest of this survey is organized as follows. In section II, dynamics modeing for MLR systems is briefly outlined. In section III and IV, studies of multi-lift rotorcraft systems are presented in detail, which consist of coordinated control strategies as well as experimental studies. In section V, the main challenging problems in this field and some development tendencies are discussed. In section VI, some concluding remarks as well as future work are given. Some studies put emphasis on configurations of MLR systems. Fig. 1 and Fig. 2 show the most common configurations in this research field [7], [29], [34], [35]. To make the mathematical modeling of the MLR systems, most of previous work made the following assumptions: 1) The tether is massless and the tether force is always nonnegative. 2) The rotorcrafts are rigid and symmetrical. 3) The slung load is regarded as a mass point. 4) Aerodynamics of the load/tether system is neglected. 5) The suspension point coincides with the rotorcrafts\u2019 center of gravity. Under the above assumption, a majority of relative modeling work was absorbed in twin-lift helicopter systems and considered the coupling in the lateral/vertical plane", + " It can be seen from references [4], [35] [19], [20], section II and Remark 1 that some uncertainties could be considered and some assumptions may do harm to systems\u2019 control. So how to derive a reasonable model for a MLR system is of great importance. (2) Reasonable configuration and structure The configurations of MLR systems are diverse, such as with support bars and without support bars. The length of tethers, the suspension position and the mass and shape of slung load are all extremely vital for the research of MLR systems. It is clear from Fig.1, Fig.2 and reference [29], [34], [35] that proper configuration and structure are beneficial to enhance the performance of MLR systems, yet there is no flight test comparing different configurations and structures. Designing proper configuration and structure for MLR systems may be quite interesting. (3) Perception and communication The perception and communication of MLR systems are indeed of great significance. It can be seen from this survey and the references mentioned in this paper that few teams have considered the perception or communication of MLR systems, which means that this topic may be of great interest" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001347_s10846-012-9725-2-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001347_s10846-012-9725-2-Figure8-1.png", + "caption": "Fig. 8 Roll angle control: motor axle closes a 5\u25e6 with the rotation plane", + "texts": [ + " gyroscopic precession will take effect. The same phenomenon could be caused by slight imperfections in wings\u2019 design. Any deficiency will induce different lift forces on right and left wing. In turn, torque produced by such a difference in lift forces will increase/decrease roll angle of the vehicle and spin axis will start to change its orientation, in effect describing a cone shape trajectory. Clearly, such behavior of the vehicle is detrimental. However, a slight modification to the spincopter construction, shown in Fig. 8 can stabilize the precession of the vehicle. As the lift force of the wing depends on the angle of attack, which corresponds with the pitch angle of the vehicle, a slight change of pitch angle will change the angle of attack. By varying the angle of attack of the wings, we can consequently cancel out the aerodynamic difference between the wings and in turn trim the precession down. In order to control the roll, the angle of motors\u2019 axle with respect to the rotational plane (Fig. 8) is slightly increased (in our case for 5\u25e6). This modification introduces new force in direction of z axis. New force generates torque acting on vehicle\u2019s pitch, thus directly changing the angle of attack of the wings. 2.4 Controller Structure Previously mentioned modification to the spincopter construction, allows us to control spincopter speed in a 3D space, as well as the roll angle of the aircraft. Controller structure is shown in Fig. 9, with separate loops for horizontal speed, vertical speed and roll angle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001673_sav-2010-0501-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001673_sav-2010-0501-Figure5-1.png", + "caption": "Fig. 5. Upper and lower suspension arms of a vehicle [24].", + "texts": [ + " Knowing the deformation, fatigue life can be predicted by using the Manson-Coffin relationship as in Eq. (14): \u2206\u03b5 2 = \u03c3\u2032 f E (2N)b + \u03b5\u2032f(2N)c (14) where \u2206\u03b5 = \u03b5max \u2212 \u03b5min is the interval of deformation, \u03c3 \u2032 f the coefficient of the fatigue strain, N the number of cycles to failure, E the Young modulus, C the exponent of the fatigue ductility, b the exponent of the fatigue strain and \u03b5\u2032f the coefficient of the fatigue ductility. Currently, the lower arm suspension of a vehicle is generally manufactured from steel. Figure 5 illustrates a vehicle suspension containing two suspension arms. The applicability of a weight optimization algorithm to this part already in service is verified. We propose a direct method of weight optimization to obtain a lighter automobile part with a safer fatigue life and a natural frequency away from the PSD frequency range. The optimization is achieved by removing elements of the automobile part that have the minimum energy density sum variation. We developed a strategy that determined the critical elements and their coordinates, and allowed the isolation of the element with the maximum sum of positive variation of the strain density energy" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000706_pi-c.1962.0036-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000706_pi-c.1962.0036-Figure2-1.png", + "caption": "Fig. 2.\u2014Admittance circle for polyphase machine.", + "texts": [ + " 1 + JSOJT 1 + JSOJT (22) fl - ^ n \\ 1 + jscar J coth /7\u00ab(1 - \u00a3 sech2 />\u00ab + i sech2 pue2^) where tan S = SCOT. This is the equation of a circle of radius (cofjLQbJp) cosech 2pu centred at (;cu/xo&/2/>) coth pa (2-sech2 pu) = (ja)ixQblp) coth 2/?\u00ab. This is the 'impedance circle' (Fig. 1) and, with appropriate change of scale, the voltage diagram for constant current drive. The 'admittance circle' can be similarly derived, for v 1 / T T ~JP 4. U Y= \\\\Z= tanhpa a)fjLb +JSCOT tanh/?w(l +$ cosech2 pu +% cosech2 pue~2Jf) where tan y = SCOT tanh2 pu = SCOT (1 \u2014 k2) = SCOTL. s < 0 s >o The admittance circle (Fig. 2) has radius O/O>JU,06) cosech Ipu and is centred at \u2014 (jp tanh/?w/2to/V) (2 + cosech2/?\u00ab) = \u2014 (jpIcoix0b) coth 2pu. With a suitable change of scale it is the current circle diagram for constant voltage drive. Bo \u2014 -^~- coth 2p\u00ab CO|A0& r' \u2014 tan (1' cosech 2DM WT(1 \u2014 &) \u2022\u00bb The circuit equivalent to Z can easily be shown to be as in Fig. 3. The rotor 72iMoss line on the admittance circle can be obtained as follows: The input power is proportional to G and the rotor I2R loss 274 PIGGOTT: A THEORY OF THE OPERATION OF CYLINDRICAL R2 The conditions for maximum torque a re : \"'WV* 1 Current drive: SCOT = 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003729_jfm.2019.776-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003729_jfm.2019.776-Figure1-1.png", + "caption": "FIGURE 1. Schematic of the dimensional problem.", + "texts": [ + " An infinite solid cylinder of radius a is decorated with a finite number N(> 1) of equally spaced grooves, infinitely extending parallel to the cylinder axis. Upon immersing the cylinder in a liquid (viscosity \u00b5), an infinite bubble is trapped in each groove. The boundary of the cylinder accordingly consists of N liquid\u2013gas interfaces, corresponding to the bubbles, and N solid\u2013air interfaces, corresponding to the ridges which separate the bubbles. It is assumed that the curvature of the liquid\u2013gas interfaces coincides with that of the solid ridges. The boundary of the cylinder cross-section is accordingly a circle. The geometry is portrayed in figure 1. Our interest lies in the two-dimensional flow which is driven by the imposed rigid-body rotation of the cylinder about its axis, say with an angular velocity \u2126 . 880 R4-2 ht tp s: // do i.o rg /1 0. 10 17 /jf m .2 01 9. 77 6 D ow nl oa de d fr om h tt ps :// w w w .c am br id ge .o rg /c or e. A cc es s pa id b y th e U CS F Li br ar y, o n 08 O ct 2 01 9 at 0 4: 57 :1 1, s ub je ct to th e Ca m br id ge C or e te rm s of u se , a va ila bl e at h tt ps :// w w w .c am br id ge .o rg /c or e/ te rm s", + " The mobility of a cylinder near a superhydrophobic wall has been addressed recently by Schnitzer & Yariv (2019); a natural follow-up of both that work and the present analysis would involve the calculation of the mobility tensor of a superhydrophobic cylinder near a no-slip wall (Kaynan & Yariv 2017). E.Y. acknowledges support from the Israel Science Foundation (grant no. 1081/16). M.S. acknowledges support from the National Science Foundation (grant no. DMS1909407). We thank M. Hodes and D. Kane for useful discussions and for their help in preparing figure 1. BENDER, C. M. & ORSZAG, S. A. 1978 Advanced Mathematical Methods for Scientists and Engineers. McGraw-Hill. BOCQUET, L. & LAUGA, E. 2011 A smooth future? Nature 10 (5), 334\u2013337. CASTAGNA, M., MAZELLIER, N. & KOURTA, A. 2018 Wake of super-hydrophobic falling spheres: influence of the air layer deformation. J. Fluid Mech. 850, 646\u2013673. CROWDY, D. 2010 Slip length for longitudinal shear flow over a dilute periodic mattress of protruding bubbles. Phys. Fluids 22 (12), 121703. CROWDY, D. G. 2013a Exact solutions for cylindrical slip\u2013stick Janus swimmers in Stokes flow" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003152_s12283-019-0302-9-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003152_s12283-019-0302-9-Figure5-1.png", + "caption": "Fig. 5 Representation of an arrow in free-flight and the vectorial magnitudes described in the mathematical model", + "texts": [ + " The average initial velocities corresponding to the actual archers are V0 = 57.3ms\u22121 for the type-A arrow and V0 = 56.7ms\u22121 for the type-B arrow. The velocity of the arrows reduces \u223c 15% and \u223c 12% for type-A and type-B, respectively, in their way to the target. The density of the air is = 1.20 kgm\u22123 and the gravitational acceleration is g = 9.81ms\u22122. A coordinate system x, y and z with the corresponding unit vectors i, j and k is defined. Here, the xy-plane forms the horizontal ground with z orthogonal to it (Fig.\u00a05). The arrows were modelled as rigid bodies; in consequence, no flexural oscillation is considered. The relative velocity is the vectorial difference of the velocity of the arrow\u2019s centre of mass (V) and the background wind component (U) or V\u2013U. The angle formed between V\u2013U and z is , whereas the angle formed between the projection of V\u2013U in the xy-plane and x is . Note that V = (Vsin cos , Vsin sin , Vcos ) and U = (ux, uy, uz) . The unit vector tangential to the trajectory of the centre of mass and parallel to V\u2013U is t = V \u2212 U\u2215|V \u2212 U|. The unit vector along the arrow\u2019s axis is n = (sin cos , sin sin , cos ). Here is the angle formed between n and z, whereas is the angle formed between the projection of n in the xy plane with x, as shown in Fig.\u00a05. During flight, the wind and the translation of the arrow generate a torque which induces rotation around the arrow\u2019s centre of mass, leading to a loss of alignment of the vectors n and t. Such a misalignment is characterized by the angle of attack defined as = cos\u22121[n \u22c5 t]. Besides the gravitational effects, a flying arrow is subject to the aerodynamic forces drag ( FD ) and lift ( FL ). The former is exerted in the direction of the relative motion between the arrow and the air flow, whereas the latter is perpendicular to such relative motion" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000374_39-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000374_39-Figure1-1.png", + "caption": "Figure 1. Chemical structure of the electropolymerizable monomers used in this study", + "texts": [ + " Electropolymerized membranes The electrodeposition of NiSTPc was performed on Pt UMEs as previously reported (9) by repeated cyclic voltammograms between -0.2 and 1.2 V in the aqueous electrolytic solution (0.1 M NaOH, pH = 13) of NiTSPc (2 mM). In addition, different membranes were also electrodeposited onto the electrodes in order to assess their ability to act as a selective barrier against interfering analytes while maintaining the ability of the electrodes to detect NO through its oxidation at 0.8 V. The monomers used to form these membranes are shown in Figure 1. The general principle for the electrodeposition process of these membranes is based on the application of a positive potential at the working electrode (ranging from 0 V to 0.8 V) that leads to the oxidation of monomers, and thus allows the polymerization process. The electropolymerized layer that is formed is not conductive and hinders further oxidation process of the monomer from the solution until the polymerization eventually stops by itself. This is exemplified in Figure 2 by the obtained repeated cyclic voltammograms of phenol in PBS on Pt UMEs (25 \u00b5m)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001264_vppc.2012.6422500-FigureI-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001264_vppc.2012.6422500-FigureI-1.png", + "caption": "Fig. I: Example of motor cross-section and magnetic flux", + "texts": [], + "surrounding_texts": [ + "Low levels of noise and vibration required for electric vehicles while driving are placing increasing demand on electric motors. Intensive development of reducing such unpleasant byproducts has been ongoing for the last couple of decades. However, vibration and acoustic noise still occur especially due to excitation forces with high order harmonics. A couple of mechanisms have been identified as contributing to vibration and acoustic noise of IPMSM. First, mechanical sources such as bearing and shaft misalignment are the important factors of vibration and noise [1,2]. Second, electromagnetic pressure generated along the gap due to Maxwell stress is applied on stator teeth, leading to cyclic deformation of motor housing at high frequencies [1,3]. To avoid instrumentation-intensive experiments, a considerable amount of research has been done in order to understand the mechanisms of vibration and acoustic noise by using numerical modeling methodology [4]. However, these studies have a couple of limitation. First, the motor geometry has been considered as a two-dimension plane in most numerical simulations [5,6]. Second, these studies have neglected the combined effect of fundamental and high order harmonics on acoustic noise, while both the fundamental and high order harmonics exist on the stator teeth at the same time [3]. To overcome the aforementioned problems, this paper presents a new numerical model for predicting acoustic noise accounting for both fundamental and high order harmonics simultaneously. The outline of this paper is as follows. First, to facilitate the acoustic noise model development, Maxwell stress tensor method and decomposition of forces based on harmonic frequencies were briefly reviewed for calculating radial magnetic forces on slot teeth. Modal analysis was conducted to investigate the vibration behavior of the motor. An acoustic noise prediction model was then developed by incorporating an acoustic indirect boundary element solver. Finally, the model was validated by comparing the prediction results with actual measurement of sound pressure level. A. Overview II. MODEL DEVELOPMENT The numerical model presented in this paper predicts noise level of electric motors over a wide range of operation conditions. The model combines electromagnetic force calculation and modal analysis together with an acoustic boundary element processor. First, radial electromagnetic forces are calculated by integrating motor geometry, material properties and operation conditions into an electromagnetic finite element solver. A modal analysis is then conducted to find natural frequencies and modal shape of the motor. Both fundamental and harmonic order forces are then applied on the tooth surface to predict sound pressure level due to the motor vibration at a given distance from the motor." + ] + }, + { + "image_filename": "designv11_33_0003169_s40516-019-0081-y-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003169_s40516-019-0081-y-Figure1-1.png", + "caption": "Fig. 1 Arrangement of laser and MAG arc (a) Front view (b) Side view", + "texts": [ + " Reliability of prediction in case of repetitive experiments and physics of laser-arc hybrid welding through analysis of weld dilution are also presented. The fillet-lap joint welding experiments were conducted using two overlapping steel plates. The metal active gas (MAG) version of GMAW is used for deposition. The SAPH 440(JIS G3113) of 2.3 mm\u00d7 100 mm\u00d7 200 mm and MG50 of 1.2 mm wire were taken as the base plate and electrode wire, respectively. The machine setup specifications are given in Table 1. The positional arrangement of laser and GMAW are shown in Fig. 1. A total 20 experiments were conducted taking welding speed, wire feed speed, current, voltage, and laser power as input parameters. The laser-arc distance was kept constant (i.e., L = 3 mm). The orientation of laser and arc sources were kept unchanged throughout the experimentation for all the welds. The weld plates of identical size were clamped such that the weld root remained in the same position for all the welds. The list of experiments is given in Table 2. The samples were prepared by cutting the specimen perpendicular to the welding direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000832_1.3657241-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000832_1.3657241-Figure8-1.png", + "caption": "Fig. 8 Ratio of s l iding A in complete fluid lubrication of roller bearing. Experimental data were taken by T. Sasaki and others.", + "texts": [ + " (6) into (2) Mo = 2RjnbKFoToxa M = 2RinbKFTT V m means first kind of elliptic integral. Pbo = Kri Rbr-j Pb = KpTRbT0 + ICPV &r)URb ( 4 ) Cr( 1 + e cos 6) where r is replaced by R. Kt,7, Kpr, and Kpr, include Sa and S* which are functions of h\u201e. Sa can be considered to be constant because the influence of Sa becomes very small if Sa takes sufficiently large value as shown in Fig. 8 in Part I. Substituting eq. (4) into eq. (1), and integrating it under the assumption that KOT, K\u201et, and K\u201e\u201e are constant, L0 = 0 L = Ii,. 8yURb 1 Cr V l ~ - 1 ( 5 ) where L0 corresponds to the case of very small values of IF, and L is the case of large IF. It is clear from eq. (5) that r0 does not affect the load capacity of bearing, and then the theoretical load capacity for non-Newtonian lubricants is the same as that for Newtonian lubricants, if the viscosity fi is replaced The yield stress To of Bingham plastic does not affect the load capacity, however, since it increases the frictional moment by the amount of the first term of eq", + " Next, complete fluid lubrication where fluid film is maintained even at the loaded side is discussed as follows. Because of continuity of cyclic behavior of the roller, the condition of Ur(e = $o) \u2014 Ur(e=ei+2*) should exist in this case. Then, from eq. (31) UT = kiW/3, . - . A = 2 / 3 ( 3 4 ) This simple expression means that rollers rotate always in the sliding state and the magnitude of sliding is expressed as A = 2 / j . This is confirmed with the experimental results by Sasaki and others [6] as shown in Fig. 8. In complete lubrication, the load capacity L and the frictional moment M in the case of sliding of roller are obtained as follows: L = M = 16 HkjWRbn 3 Cr 16 fikiWRibn S / R 3 Vcr Vl + (35) Journal of Basic Engineering M A R C H 1 9 6 2 / 1 7 9 Downloaded From: http://fluidsengineering.asmedigitalcollection.asme.org/ on 01/30/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use The magnitude of load capacity is shown in Fig. 2 with dashed lincs. Both load capacity and frictional moment in this case de crease by 33 per cent of theoretical values considering no sliding" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001468_s0022112010004489-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001468_s0022112010004489-Figure1-1.png", + "caption": "Figure 1. Schematic for two spherical drops settling under gravity (not to scale).", + "texts": [ + " Our method consists, at each time step, of (i) solving the hydrodynamical problem for two drops given the instantaneous surfactant distribution, (ii) updating the surfactant concentration through the transport equation and (iii) economical truncation of multipole expansions involved in (i) and (ii). These steps are detailed below. 3.1. Solution of the hydrodynamical problem It is convenient, and computationally most efficient, to work in two, moving and rotating coordinate systems (x1, y1, z1), (x2, y2, z2) centred at the drop centres O1, O2 and the z1-, z2-axes along the line of centres from sphere S2 to S1 (figure 1); these coordinate systems differ only by the translation. By definition, the x-axis lie in the plane span on the gravity vector and the line of centres. Along with (x\u03b3 , y\u03b3 , z\u03b3 ), spherical coordinates (r\u03b3 , \u03b8\u03b3 , \u03d5\u03b3 ) (index \u03b3 labels drops 1 and 2, and related quantities) are used for an observation point M (figure 1) x\u03b3 = r\u03b3 sin \u03b8\u03b3 cos \u03d5, y\u03b3 = r\u03b3 sin \u03b8\u03b3 sin \u03d5, z\u03b3 = r\u03b3 cos \u03b8\u03b3 , (3.1) where \u03d5 = \u03d51 = \u03d52 is the common angle of positive rotation about the z-axes. The velocity field ue outside the two spheres is sought as ue = u1 \u2212 + u2 \u2212, (3.2) where u\u03b3 \u2212 is a Stokes flow (to be found) regular everywhere outside the sphere with surface S\u03b3 and represented by Lamb\u2019s (Happel & Brenner 1973) singular series: u\u03b3 \u2212 = \u221e\u2211 \u03bd=1 [ \u2207 \u00d7 ( r\u03b3 \u03c7 \u03b3 \u2212(\u03bd+1) ) + \u2207\u03a6 \u03b3 \u2212(\u03bd+1) \u2212 (\u03bd \u2212 2)r2 \u03b3 \u2207p \u03b3 \u2212(\u03bd+1) 2\u03bd(2\u03bd \u2212 1) + (\u03bd + 1)p\u2212(\u03bd+1)r\u03b3 \u03bd(2\u03bd \u2212 1) ] ", + " 4\u03c0(\u03bd + m)! ]1/2 P m \u03bd (cos \u03b8)eim\u03d5, (m 0) Y\u03bd,m(r) = (\u22121)mY \u03bd,\u2212m(r), \u23ab\u23aa\u23ac\u23aa\u23ad (3.5) P m \u03bd (x) = (1 \u2212 x2)m/2dmP\u03bd(x)/dxm is the associated Legendre function (in the definition of Korn & Korn 1968) and the overbar stands for complex conjugation; a1 = k and a2 = 1 are the non-dimensional drop radii here and in the rest of the paper (unless otherwise stated). Complex coefficients in (3.4) obey A \u03b3 \u2212(\u03bd+1),m = (\u22121)mA \u03b3 \u2212(\u03bd+1),\u2212m, etc. (3.6) to make (3.4) real valued. For the gravity-induced motion shown in figure 1, there is obvious symmetry of the solution about the y = 0 plane (with constraints on the coefficients in (3.4)), which is not assumed, though, in the method description below to make it general and suitable for other problems (e.g. flow-induced motion, without such symmetry). The surface tension, \u03c3 , on the drop surfaces is assumed given by its expansions in spherical harmonics, equivalent to Ma \u0393 \u2223\u2223\u2223 S\u03b3 = \u221e\u2211 n=0 \u03c3\u03b3 n , \u03c3 \u03b3 n = n\u2211 m=\u2212n \u03c3 \u03b3 n,mYn,m(r\u03b3 ), (3.7) with complex coefficients satisfying (3", + "5) take the form A \u03b3 \u22122,0 = 2 ( \u03c0 27 )1/2 a\u03b3 cos \u03b2, A \u03b3 \u22122,1 = \u2212 ( 2\u03c0 27 )1/2 a\u03b3 sin \u03b2, (3.14) where \u03b2 is the angle between the centre-to-centre vector O2 O1 and the gravity vector. Upon convergence of iterations, the drop velocities V \u03b3 are found from (3.13). 3.2. Solution of the surfactant transport equation It is most natural, and computationally convenient, to solve the surfactant transport equation (2.3) on each sphere S\u03b3 in the reference frame coincident with the axial coordinate system (x\u03b3 , y\u03b3 , z\u03b3 ) of figure 1. This reference frame translates with the drop centre velocity, V \u03b3 , and rotates with the angular velocity, (V 1 1 \u2212 V 2 1 )/R12, (about the y\u03b3 -axis) both known from the solution of the hydrodynamical problem. The fluid velocity, u\u2217, on S\u03b3 relative to this frame is obtained from (3.2)\u2013(3.4), (3.8)\u2013(3.9) and the boundary conditions (3.12) u\u2217(x) = a2\u2207S\u03a0 + a\u2207S\u03a8 \u00d7 r, \u03a0 = \u221e\u2211 n=1 n\u2211 m=0 bn,mDm n (\u03b7) cos m\u03d5, \u03a8 = \u221e\u2211 n=1 n\u2211 m=1 cn,mDm n (\u03b7) sin m\u03d5, \u23ab\u23aa\u23aa\u23aa\u23ac\u23aa\u23aa\u23aa\u23ad (3.15) where the real coefficients are bn,m = (2 \u2212 \u03b4m,0) (1 + \u00b5\u0302) [ (2n \u2212 1) (n + 1)a2 B\u0303n,m + 1 2(n + 1) An,m \u2212 \u03c3n,m (2n + 1)a ] , cn,m = \u2212 2(2n + 1) a[n + 2 + \u00b5\u0302(n \u2212 1)] Im(Cn,m) \u2212 ( 8\u03c0 3 )1/2 ( V 1 1 \u2212 V 2 1 ) R12 \u03b4n,1\u03b4m,1, \u23ab\u23aa\u23aa\u23aa\u23ac\u23aa\u23aa\u23aa\u23ad (3", + " Two modes of mutual approach were predicted and quantified, \u2018rapid coalescence\u2019 (for a sufficiently strong driving force) on a time scale commensurate with that for clean drops, and long-time approach (for a subcritical driving force) when the film drainage may be even slower than for rigid spheres. Since |V 2 \u2212 V 1| \u223c \u03b41/2 in our simulations for figure 12, this case appears to be characteristic of \u2018rapid coalescence\u2019. Besides small, but finite diffusion, though, the present asymmetric case has additional complications. Figure 13 (solid lines) presents the distribution of surfactant along the gap for drop 1 in the plane y1 = y2 = 0, as a function of a stretched variable x1\u03b4 \u22121/2 for the near-contact stage of simulation in figures 11(a) and 12 at Pes = 100; the axis x1 is defined in figure 1 and is (for the given case) antiparallel to the direction of the doublet motion. For comparison, dashed lines show an analogous simulation with less diffusion (Pes = 500). All curves were calculated for azimuthal truncation M = 8 (\u00a7 3.3) and found to be graphically indistinguishable from those for M = 4; variation of \u03b5-parameter (controlling the order of harmonics retained, \u00a7 3.3) confirms excellent numerical convergence of the results in figure 13. The range |x1\u03b4 \u22121/2| 8 essentially limits the data to near-contact region, both for \u03b4 = 0", + " Although an analytical theory of surfactant-enhanced coalescence is still lacking, we have been able to verify the direction of relative motion for the simulation in figure 12(a) using the boundary-integral algorithm of Rother et al. (2006) at \u00b5\u0302 = 1, Pes = 100 and finite, but small deformation. Enhanced coalescence phenomenon for two equal drops sedimenting along the line of centres (figure 11b) at high Pe\u0301clet numbers presents an interesting case, where a nearly stagnant cap of surfactant and a clean spot include the lubrication area. Referring to figure 1, where drop 1 is assumed to be strictly below drop 2 (and both sediment downwards), figure 14 presents the axisymmetrical distribution of surfactant versus \u03b81 (or \u03b82) angle on both surfaces at \u03b4 = 3 \u00d7 10\u22125 and Pes = 100, 500 and 20 000. The Pes = 5000 result is almost indistinguishable from that for Pes = 20 000 and is not shown. Although the process is not quite steady state, nearly stagnant caps of surfactant are formed as Pes \u2192 \u221e, on the leading and trailing drops, with cap angles \u03b8c 1 \u2248 98\u25e6 and \u03b8c 2 \u2248 106\u25e6, respectively; the rest of each drop is free from surfactant" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001560_025701-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001560_025701-Figure6-1.png", + "caption": "Figure 6. Illustration of the liquid layer thickness on the different sizes of the filter paper. The central zone is the part of the solution in the fluorescence detection region.", + "texts": [ + " Again, in the figure, the PD saturation means that the PD detector is saturated or reached the maximum detection limit. Figure 5 illustrates the effect of the filter paper size on the fluorescence intensity. Two different sizes (5 \u00d7 5 and 10 \u00d7 10 mm2) of the filter paper were used. The volume and the concentration of the MQAE solution are 10 \u03bcL and 0.25 mM, respectively, and the volume of the chloride solution is 10 \u03bcL. The results show that, for the same chloride concentration, the fluorescence intensity is higher in the smaller paper than the larger one. This effect may be understood as follows (see figure 6). For a given volume of the MQAE solution and a given volume of the chloride solution, the larger filter paper can absorb the liquid more and spread the liquid over a larger area; the liquid layer immediately above the photodetector sensing surface is thinner; consequently, the fluorescent molecules are fewer in the detecting region and a weaker fluorescence signal is detected. On the other hand, the same amount of liquid (the MQAE solution plus the chloride solution) forms a thicker liquid layer over the photodetector sensing surface; consequently more fluorescent molecules in the detection region and a stronger fluorescence signal is detected" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003921_b978-0-12-803581-8.11722-9-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003921_b978-0-12-803581-8.11722-9-Figure7-1.png", + "caption": "Fig. 7 Meridional cross-section of the considered hydrodynamic clutch: 1 \u2013 pump, 2 \u2013 turbine, 3 \u2013 average line.", + "texts": [ + " The result is the relation for one gap between the cylinders: M\u00bc 2 p b mp r13 h o\u00fe 2 p b t0r12 \u00f015\u00de If the cylindrical viscous clutch or brake consists of a few cylinders, then the transferred torque is calculated by adding up the torques calculated with Eq. (15) for radiuses r1 of individual gaps. Model of hydrodynamic clutches with ER fluid The mathematical model of hydrodynamic clutches with ER fluid is created on the basis of the average through-flow stream (Olszak et al., 2019b) which assumes that the flow of fluids within a clutch is fixed, continuous and takes place along a single average through-flow line, Fig. 7. Due to the fact that in hydrodynamic clutches and brakes with ER fluid occurs a flow mode, the calculations are conducted with the assumption that an increase in electrical field intensity E results in an additional increase in fluid flow resistance \u2013 a pressure drop Dpt, similarly to occurrences in valves with ER fluid. For hydrodynamic clutches or brakes with flat radial blades, to calculate the transferred torque M the relation is used: M\u00bc r cmA o1 \u00f0r22 r12ik\u00de \u00f016\u00de where: r \u2013 density of the working fluid, cm \u2013 meridional velocity of the average through-flow stream, A \u2013 area of the clutch crosssection, o1 \u2013 angular velocity of the pump, r1 \u2013 radius of the average through-flow stream on the turbine output and on the pump input, r2 \u2013 radius of the average through-flow stream on the turbine output and on the turbine input, ik \u00bc o2/o1 \u2013 velocity ratio" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002966_s12161-018-1395-7-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002966_s12161-018-1395-7-Figure3-1.png", + "caption": "Fig. 3 Comparison diagrams of a half-peak widths and b peak responses of four analytes on different material electrodes. Experimental conditions: fused silica capillary, 25 \u03bcm i.d. \u00d7 19.5 cm; buffer solution, 50 mM Na2B4O7-NaOH buffer (pH 9.35); separation voltage, 3 kV; injection time, 4 s (at 3 kV)", + "texts": [ + " The peak areas of four analytes increased with increasing the content of chloroauric acid; as the ratio value was 1:1, they all achieved maximum peak area, the peak shape, and resolution were satisfactory (The corresponding electrophoretograms of four analytes were shown in Fig. S4 of BOnline Resource^). The assay results further verified above conclusion that AuNPs and rGO play an equally important role for the electrochemical responses of target analytes. Furthermore, the half-peak widths and peak responses of four model target analytes on the electrodes of AuNP/rGO, AuNPs, AuNP/Gr, rGO, and Gr were compared with the miniCE-AD system. As shown in Fig. 3, it was obvious that four target analytes had smaller half-peak widths and higher peak responses on AuNP/rGO composite electrode than the other electrodes. The assay results suggested that the synergistic effects of AuNPs and rGO can promote the electron transduction, improve the electrocatalytic activity, and enhance the responses of the analytes, which was beneficial to improve the detection sensitivity of CE-AD approach. (The corresponding electrophoretograms of four analytes were shown in Fig", + " The sample electrophoretograms indicated that the co-existing substrates in beer did not interfere with the analyses of the target analytes under the optimal conditions. To further evaluate the reliability of this analytical system, the recovery experiments were also carried out using the beer sample at three different concentrations. As shown in Table 2, the assay results showed that the average recoveries of four target analytes were between 84.5 and 108%. Experimental conditions were the same as those in Fig. 3 a ND meant the content of the target analyte in tested samples was lower than its LOQ of this method In this work, AuNP/rGO composite electrodes were prepared by in situ chemical reduction and successfully applied to the mini-CE-AD system for analyzing the main bioamines and their precursor amino acids in beer. In combination with mini-CE-AD analytical system, AuNP/rGO composite electrode could achieve high detection sensitivity and selectivity for four target analytes at a relatively low and constant oxidation potential, and was competent to detect the real beer samples without derivatization or extra pretreatment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001045_jsea.2010.312134-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001045_jsea.2010.312134-Figure8-1.png", + "caption": "Figure 8. FE model of the motor assembly.", + "texts": [ + "3739, the solution converges in the 38th generation (Figure 4) and it is found to be less than that of obtained by SA based algorithm which is $ 238.206 [29]. The values of the variables are as follows. x-base flatness x1 = 0.086439, motor base flatness x2 = 0.08 , motor shaft size x3 = 0.106116, and the motor shaft perpendicularity x4 = 0.078027. It can be concluded that the proposed hybrid methodology with BP and NSGA II can solve tolerance synthesis problem effectively. The FEA integration (Figure 8-11), helps in determining deformation due to inertia effects like gravity, velocity,acceleration, etc., resulting in decrease in the critical assembly feature. The CAD integration (Figure 5), helps in determining contribution of various tolerances towards the critical assembly feature. The assembly model of the motor assembly is shown in Figure 6. The exploded view of the motor assembly is shown in Figure 7. In this research, the proposed approach provides better formulation of cost-tolerance relationships for empirical data" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000945_ssd.2010.5585533-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000945_ssd.2010.5585533-Figure4-1.png", + "caption": "Figure 4. Rotor magnetic circuit including tewlve claws.", + "texts": [ + " The two halves of the field winding are inserted in both sides of the machine between the armature end-windings and the housing as illustrated in figure 3. They are con nected in series in such a way to produce additive fluxes. Following the transfer of the DC-excitation winding from rotor to stator, appropriate changes of the magnetic circuit have been introduced. These concerned mainly the rotor where the two iron plates with overlapped claws facing the air gap tum to be magnetically decoupled. Figure 4 shows a photo of the rotor of the CPAES. Following the removal of the field winding from rotor to stator and for the sake of an efficient flow of the flux, two magnetic collectors have been included in the stator mag netic circuit. These guarantee the flux linkage between rotor and stator. They are embedded on the two flasks of the machine. Figure 5 shows a photo of one magnetic col lector. The CPAES static and rotating components are illustrated in figure 6. One can notice that each claw plate includes six poles" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001827_s0219455417400120-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001827_s0219455417400120-Figure1-1.png", + "caption": "Fig. 1. Compliant foil bearing.", + "texts": [ + " The equations are solved by means of numerical calculations. As a result of the work, a parametric model has been created. With further developments, the model will be universally used for calculation of foil journal bearings dynamics. Moreover, computational methods o\u00aeer the possibility of deeper understanding of the phenomena occurring in the foil bearing. The foil bearing is a self-acting support of aerodynamic type, so the external pressurization is not needed. It is made up of compliant surfaces (a bump and a top foil), as shown in Fig. 1. During the foil bearing operation, the top foil is clenched on the rotating journal by means of the elastic bump foil. The aerodynamic \u00aflm of a very low thickness, theoretically close to the cylindrical one, is generated by viscosity e\u00aeects. Gas bearings have limited load capacity in comparison to liquid-lubricated bearings, because of low viscosity of the medium. The operation of foil bearings is inevitably related with friction. Miazga et al. have shown that the important application problem is related to the start-up and the shut-down in contact with the shaft surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002457_itcosp.2017.8303079-Figure10-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002457_itcosp.2017.8303079-Figure10-1.png", + "caption": "Figure 10: the simulation setup for pick and place using RobotStudio", + "texts": [], + "surrounding_texts": [ + "The communication between MATLABR2016a and the IRC5 robot controller is carried out over the Ethernet CAT5 cable. The robotics system toolbox installed in the matlab communicates with the IRC5 controller which transfers the location of the object in the workspace of the robot. The position value is then stored in the registers1, 2, 3 of the controller and the orientation values are stored in the registers 4, 5, 6 respectively. The values thus saved are then used in pick and place operation program written in the RAPID programming language in the flex pendant to modify the position value as well as the orientation value of the tool attached to the real robot. Pick and place operation using RobotStudio RobotStudio is ABB\u2019s simulation and programming software. The pick and place operation was simulated using this software. The robot has to perform different routines according to the object information. Hence different routines are developed and are executed only if the respective inputs become high. So before simulation of pick and place operation is executed it is required to attach the gripper tool to the robot tool frame. Once the tool is attached we need to define the tool and attach smart components to the gripper require for pick and place operation. After making the robot along with the gripper tool ready to perform the pick and place operation, now the robot is taught the path it has to follow using the RobotStudio. Finally the simulation setup is transferred to the real robot [8]. Hardware setup Moving object platform Other than the robot the object is also moving in this proposed method. So another moving robotic vehicle is developed for moving the object to be picked. It is controlled by 8051 microcontroller. And when the object comes in the vicinity of the camera attached to the ABB robot, instruction is send to the scrap robot to stop the movement. Then the object is picked from the robot and placed. The figure shows the robotic vehicle developed to carry the object to be picked The motion control of the robotic vehicle carrying object to be picked is done using L293D dual H-bridge motor driver integrated circuit(IC). The vehicle is made to move in a rectangular path around the robot workspace. The 8051 microcontroller is used to programme the robot. Keil\u03bcvision software was used to compile the programme." + ] + }, + { + "image_filename": "designv11_33_0003253_012016-Figure12-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003253_012016-Figure12-1.png", + "caption": "Figure 12. Result of the FE-model based modal analysis (a), functional qualification by laser scanning vibrometry (b).", + "texts": [ + " Computer tomography verifies the material bond between the actuator system and the hip stem (Figure 11c). Static properties of the completed hip stem correspond to the properties of the volume fractions of the heat-treated and the as-build parts. Dynamic measurements experimentally validated the functionality of the embedded actuator. A numerical modal analysis using a finite element model is utilized to identify the initially unknown natural frequencies. The first two modes for lateral and vertical bending of the hip stem are located at 2180 Hz and 3488 Hz, resp. (Figure 12). These frequency values are used for the excitation for the WTK IOP Conf. Series: Materials Science and Engineering 480 (2019) 012016 IOP Publishing doi:10.1088/1757-899X/480/1/012016 next investigations. The analysis of the dynamic response of the implant was carried out using a 3D laser scanning vibrometer to measure the surface velocities in three spatial directions. The experimentally determined frequencies inducing the largest displacement and velocity at 2135 Hz and 3416 Hz deviate slightly from the simulation values" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001587_iros.2011.6094899-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001587_iros.2011.6094899-Figure3-1.png", + "caption": "Fig. 3. Biped model used for simulation.", + "texts": [ + " Therefore it is placed as the first priority and torso angle velocity tracking as the second. This approach ensures that no torso control output will interfere with the desired COG motion. Once the joint speeds,~\u0307\u03b8r, are know from (12), the final joint reference angle ~\u03b8r is calculated as ~\u03b8r = ~\u03b8 + ~\u0307 \u03b8r\u2206t, (13) where \u2206t is the control step period. Finally, local joint PD controllers track ~\u03b8r to achieve the desired COG position and torso orientation. The proposed controller was tested in simulation on a 7 link biped, shown in Fig. 3, with the mass and link parameters listed in Table I. The simulation model, is based on ABL-BI an experimental biped platform designed at Carleton University [18]. For comparative purposes, two controllers were tested; a classical ZMP based control based on [4] and the CMP based controller described in this paper. Both controllers were able to generate stable walking with equivalent walking parameters, when no disturbance was applied. A second test consisted of generating a stable walk and applying a disturbance to the robot mid stride. The magnitude of the disturbance was 40 N applied for two seconds. The disturbance was also applied to the torso at a height of 0.1 m relative to Frame 0 (refer to Fig. 3). Hence, this disturbance also induced a moment of 10 Nm about the COG. For the disturbed test the CMP controller is able to compensate for the disturbance and returns to a stable walk as seen in Fig. 4. The behavior of the controller during stable walking, as well as when the disturbance is applied, is best shown by the modification and tracking of the the reference COG motion compared to the planned reference CMP as shown in Fig. 5. The disturbance, applied between 16 s and 18 s, causes reference COG to diverge from the reference CMP by up to 6 cm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000945_ssd.2010.5585533-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000945_ssd.2010.5585533-Figure6-1.png", + "caption": "Figure 6. CPAES topology. Legend: (a) stator, (b) rotor.", + "texts": [ + " These concerned mainly the rotor where the two iron plates with overlapped claws facing the air gap tum to be magnetically decoupled. Figure 4 shows a photo of the rotor of the CPAES. Following the removal of the field winding from rotor to stator and for the sake of an efficient flow of the flux, two magnetic collectors have been included in the stator mag netic circuit. These guarantee the flux linkage between rotor and stator. They are embedded on the two flasks of the machine. Figure 5 shows a photo of one magnetic col lector. The CPAES static and rotating components are illustrated in figure 6. One can notice that each claw plate includes six poles. Moreover it is to be noted that the two claw plates are magnetically decoupled. In order to reduce the computation time, the FEA study domain is limited to a one pair of poles of the CPAES. Figures 7 and 8 show the stator and the rotor study domains, respectively. The meshing of both stator and rotor study domains has been automatically generated in the Ansys-Workbench environment. Figure 9 shows a mesh of the stator yoke and of the two associated magnetic collectors" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003375_ilt-10-2018-0390-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003375_ilt-10-2018-0390-Figure7-1.png", + "caption": "Figure 7 The testing bearing system", + "texts": [ + " T0 T1 Rs s0 1 T0 Ta Rs0 a \u00bc 0 T1 T0 Rs s0 1 T1 T2 Ri s \u00bc 0 T2 T1 Ri s 1 T2 T4 Ri g \u00bc Hi T4 T2 Ri g 1 T4 T3 Rb g \u00bc 0 T4 T3 Rb g 1 T4 T5 Rg o \u00bc 0 T5 T4 Rg o 1 T5 T6 Ro h \u00bc Ho T6 T5 Ro h 1 T6 T7 Rh h7 \u00bc 0 T7 T6 Rh h7 1 T7 Ta Rh7 a \u00bc 0 8>>>>>>>>>>>>>>>>>>>>< >>>>>>>>>>>>>>>>>>>>: (18) Here, the calculation of Ri s, Ri g, Ri g, Rb g, Ro-g and Ro h are detailed in the literature (Yan andHong, 2016): Rh h7 \u00bc Zln Dh D 2khpC ;Rh7 a \u00bc Z 2hhpDhLh ; Figure 5 Bearing system nodes distribution Angle contact ball bearing Pingping He, Feng Gao, Yan Li, WenwuWu and Dongya Zhang Industrial Lubrication and Tribology D ow nl oa de d by B os to n C ol le ge A t 2 0: 42 1 6 M ay 2 01 9 (P T ) Rs s0 \u00bc 2Z kspB ;Rs0 a \u00bc 4ZLs kspd2 1 4Z hspd2 The calculation process of thermo-mechanical coupling model of the bearing is shown in Figure 6. The test platform is shown in Figure 7(a). FAG B7008C bearing and special grease(NLGI:2-3, working temperature range: 40 120\u00b0C, the viscosity of base oil at 40\u00b0C is 25 mm2/s) are used, the initial interference fit between the shaft and the inner ring is 5 mm, and the structural and material parameters of the bearing system are shown inTable I. During the testing process, the sensor layout is shown in Figure 7(b). A three-component force sensor TR3D-A-1K is used to monitor actual axial load of the bearing. Three platinum thermal resistors PT100 are mounted in the holes which are uniform distributed in the circumferential direction of the bearing housing surface matched with the outer ring. The thermocouple is attached to the surface of the bearing housing, and is in a circumference with the platinum thermal resistors. The sensors\u2019 data are obtained by multi-channel data logger HIOKI. The initial preload of the bearing is 200 N" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001452_1464419311408949-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001452_1464419311408949-Figure1-1.png", + "caption": "Fig. 1 Geometrical features of the contacting regions of ball bearings [15]", + "texts": [ + " Stiffness values associated with these contacts are calculated using the Hertzian theory. If the ball bearings are running at high operational speeds, the theory of elastohydrodynamic lubrication (EHL) must also be taken into account when modelling the dynamics of ball bearings. This type of contact is called EHL contacts [1]. The stiffness effect of the lubrication film is neglected in this study. The rolling elements in a ball bearing are in contact with the inner and the outer raceways [15]. Figure 1 shows the major geometric features of these contacting bodies and ball bearings. Proc. IMechE Vol. 225 Part K: J. Multi-body Dynamics at Glasgow University Library on December 22, 2014pik.sagepub.comDownloaded from Referring to Fig. 1, if Rre represents the radius of the ball itself, the radii of the curvatures for the inner contact are R1x \u00bc Rre \u00f01\u00de R1y \u00bc Rre \u00f02\u00de R2x \u00bc dm=2 cos\u00f0 \u00de Rre \u00f03\u00de R2y \u00bc Ri \u00f04\u00de In a similar manner, the radii of curvature for the outer contacts are R2x \u00bc dm=2 cos\u00f0 \u00de \u00fe Rre \u00f05\u00de R2y \u00bc Ro \u00f06\u00de When calculating the curvature radii of the contacting surfaces, the assumption of a 0 contact angle ( ) can give satisfactory results for deep groove ball bearings [15]. The geometric features between two contacting solids can be expressed in terms of the curvature sum R, and curvature difference Rd which are described in reference [15] as 1 R \u00bc 1 Rx \u00fe 1 Ry \u00f07\u00de Rd \u00bc R 1 Rx 1 Ry \u00f08\u00de 1 Rx \u00bc 1 R1x \u00fe 1 R2x \u00f09\u00de 1 Ry \u00bc 1 R1y \u00fe 1 R2y \u00f010\u00de Rx and Ry represents the effective radii of curvature in the x- and y-planes [16]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001589_1.4003148-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001589_1.4003148-Figure2-1.png", + "caption": "Fig. 2 Fluid thickness in the ridge or groove regions in terms of circumferential coordinates", + "texts": [ + "org/ on 01/28/2016 Terms less than the bearing radius, the fluid film can be unwrapped into a plane. For a grooved bearing, the modified Reynolds equation for the steady state is 13 x ch 3 12 g x + y ch 3 12 g y = c u 2 h x 1 where the coordinate system x ,y is fixed to the bearing, is the density ratio of a fluid density divided by a fluid density in the cavitation region c, is the bulk modulus, and g is a switch function in the flow field of the full-film and the cavitation region. Fluid thickness in the ridge or groove region in terms of circumferential coordinate is shown in Fig. 2 h = c 1 + cos in the ridge 2a or h = cg + c 1 + cos in the groove 2b The pressure can be obtained from the fluid density as follows: P = Pcav + g ln 3 However, for a smooth bearing, the film thickness changes as the grooved journal rotates. This problem can be solved by assuming that the groove journal is stationary and the smooth bearing is rotating in the direction opposite to that of the groove journal 14 . The pressure boundary conditions at bearing edges are p , L 2 = p ,\u2212 L 2 = 0 4 When pressure has been solved for in the equilibrium state, the radial and tangential loads can be obtained by integrating the pressure over the bearing area Wt = A pr sin \u2212 dyd 5a Wr = A pr cos \u2212 dyd 5b The load capacity can be derived as W = Wt 2 + Wr 2 1/2 6 In a dimensionless form, the load capacity is W\u0304 = W r2 c r 2 7 Also, the load distribution along the axial direction can be ob- tained by integrating the pressure force circumferentially Transactions of the ASME of Use: http://www" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001986_iet-epa.2016.0565-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001986_iet-epa.2016.0565-Figure6-1.png", + "caption": "Fig. 6 3D mesh models of straight rotor, skew rotor and dual skew rotor induction motor (a) induction motor with straight rotor, (b) induction motor with skew rotor, (c) induction motor with dual skew rotor", + "texts": [ + " The rotor harmonics (list k2 = \u00b1 1, \u00b1 2, \u2026) induced by stator harmonics will match up with other stator harmonic in same order number in locked rotor period. The \u221234 pole first-order rotor slot harmonics field (mark with frame in Table 5) will interact with \u221234 pole first-order stator harmonic field at rotor locked. The generated synchronous torques are inevitable during the motor start. Some important values of locked rotor torque in Fig. 8 are shown in Table 6. 3D mesh models of straight rotor, skew rotor and dual skew rotor induction motors with 36\u201336 are shown in Fig. 6. The locked rotor torques are analysed in 21 stator and rotor relative positions in the simulations. The convergence error is <10% and the adaptive frequency is 50 Hz. The average mesh number in induction motor with straight rotor is 433,824; the average mesh number in induction motor with skew rotor is 472,092; and the average mesh number in induction motor with dual skew rotor is 483,610. The spatial distributions of air-gap magnetic field in dual skew rotor induction motor are shown in Fig. 7. The locked rotor torque varies between \u2212268.5 and 419.5 N m and average torque is 95.7 N m as shown in Fig. 6a. It can found that the peak-to-peak value of the locked rotor torque in Fig. 8b is 2/5 times the torque of straight bars in Fig. 8a. The value of average torque is 80.9 N m with 36 and 36 slot numbers in Fig. 8b. The range of positive locked rotor torque in Fig. 8b is enlarged, compared with Fig. 8a. It can also be found that synchronous torque created by interaction of the first-order stator slot harmonics and first-order rotor slot harmonics still exists in skew rotor induction motor. Skew attenuates the magnitude of synchronous torques" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001070_1.4002694-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001070_1.4002694-Figure4-1.png", + "caption": "Fig. 4 A prototype of 2SPS+UPR PM", + "texts": [ + "org/ on 01/29/2 vqi = i rqi i = \u2212 rqi\u0302i i = \u2212 rqi\u0302iJ iV Vqi = vqi qi = JqiV, Jqi = \u2212 rqi\u0302iJ i J i aqi = rqi i i + rqi i i i 38 From Eq. 9 and the relative equations in Sec. 3, vpi , pi ,Vpi ,api , pi of the ith piston can be derived as follows: pi = i, pi = i vpi = i ri \u2212 rpi i + vri i = v + ei \u2212 rpi i i = v \u2212 e\u0302i + rpi\u0302i i = E \u2212 e\u0302i + rpi\u0302iJ i V Vpi = vpi pi = JpiV, Jpi = E \u2212 e\u0302i + rpi\u0302iJ i J i api = a + ei + ei \u2212 rpi i i \u2212 rpi i i i 39 5 Kinematics of Linear Legs for 2SPS+UPR PM 5.1 Characters of 2SPS+UPR PM. A 2SPS+UPR PM has 4DOF in Ref. 27 see Fig. 4 . It is composed of a moving plat- FEBRUARY 2011, Vol. 3 / 011005-5 016 Terms of Use: http://www.asme.org/about-asme/terms-of-use f t w e a L = w a a j r c R S a o t i a b p E b c H a e 0 Downloaded Fr orm m, a fixed base B, two SPS-type active legs ri i=1,3 with he linear actuators, and one UPR-type composite active leg r2 ith a linear actuator and a rotational actuator. Here, m is an quilateral triangle link with three sides li= l, three vertices bi, and central point o. B is an equilateral triangle link with three sides i=L, three vertices Bi, and a central point O" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000961_09615531111177787-Figure11-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000961_09615531111177787-Figure11-1.png", + "caption": "Figure 11. Temperature contours for (a) without heat sink and (b) with heat sink during welding", + "texts": [ + " Convergence of cost function and cooling flux due to jet with iterations are shown in Figures 8 and 9. After a sharp change in the fourth iteration, the values of cost function and cooling flux does not undergo considerable change with further iterations. The cost HFF 21,8 1060 function converges to a value of 66,326. The cooling flux for the last iteration 703,972 W/m2. The temperature history of the node 5 mm away shown in Figure 10 with the introduction of cooling estimated from IHT shows a good agreement. The temperature contours obtained with heat sink and without heat sink is shown in Figure 11. The tail effect of the hot region resulting in a elliptical contour due to the weld torch has been reduced in size due to the introduction of the heat sink. The size of the hot region in the temperature range of 1,0408C to 7908C is significantly reduced in size due to the introduction of the cooling sink. Figure 9. Heat flux vs iteration of IHT study 800,000 600,000 500,000 700,000 400,000 300,000 200,000 100,000 0 0 2 4 6 8 10 Iterations H ea t f lu x (W /m 2 ) Residual stresses and distortions in arc welding 1061 A good agreement in temperature history of a point 5 mm away from the weld line is obtained between the experimental and numerical study" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003753_chicc.2019.8866455-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003753_chicc.2019.8866455-Figure6-1.png", + "caption": "Fig. 6. Prototype of the 1-kW qZSI with minimized capacitance.", + "texts": [ + " Therefore, the switching device voltage stress is increased by 53% compared with the conventional design. If the grid voltage is increased to 240 Vrms, the inverter power rating is kept the same and the input voltage is 260\u2013340 V, and C1 is decreased from 800 \u03bcF required in conventional design to 100 \u03bcF with the proposed control, the switching device voltage stress is increased by only 15%. Therefore, there is more benefit to apply the proposed method in qZSI with higher output voltage. The picture of the 1-kW qZSI prototype with minimized capacitance is shown in Fig. 6. TPH3006PS GaN devices from Transphorm are selected as the inverter switches and the switching frequency is selected at 100 kHz. In the experimental study, the Magna Power Electronics XRii 250\u201316 was used as the PV emulator. The waveforms of vDC , vC 2 , and vIN for the qZSI system with 2 mF capacitor are shown in Fig. 7. The input voltage was 150 V and the output voltage was 120 Vrms. The inverter was operated at 300 W output power condition. The double-frequency components in vDC , vC 2 , and vIN were 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002746_s41314-018-0014-0-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002746_s41314-018-0014-0-Figure3-1.png", + "caption": "Fig. 3 RW setup", + "texts": [ + " Secondly, the methodology for dynamic characterisation of the rubber/foam material is exposed. Thirdly, the material characteristics are derived. The experiments consist in shooting real KENLW projectiles against a RW structure. A RW structure is a metallic structure equipped with a force sensor for the measurement of the impact force. It is assumed infinitely rigid with respect to the deformations that take place in the projectile. The RW experimental setup used to obtain the projectile firing data is shown in Fig. 3. The main components are: 1. A pneumatic launcher for launching the projectile at desired velocity; 2. A light screen barrier for the impact velocity measurement; 3. A set of led lights for the high-speed camera; 4. A high-speed camera for impact event recordings. Through the processing of the impact videos, a great amount of information concerning the kinematics of impact can be obtained; 5. The impacting projectile; 6. A RW structure equipped with a force sensor Kistler 9061a [26] for the impact force measurement" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003400_978-3-030-19648-6_41-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003400_978-3-030-19648-6_41-Figure6-1.png", + "caption": "Fig. 6. Applying free-form surface load on a beam during the bionic structural optimization process: (a) feature surfaces of the beam before optimization, (b) during the shape optimization the feature surfaces of loading and unloaded area are merged, (c) the load domain and fixation domain, (d) Von Mises stress in the optimized structure", + "texts": [ + " In our application we first construct the ball bearing as a sphere with 5 mm radius (Fig. 5a). With our new method, we use a small cylinder with radius 1 mm on the top of the sphere to define the loading domain (Fig. 5b). The fixation domain is defined by the same cylinder but on the bottom of the sphere. Figure 5c and d shows a plausible result of the FEM analysis of the ball bearing, which means the free-form surface load is successfully applied with our method. As already described, the surface profile of a model is always changing during the bionic shape optimization. Figure 6 shows an example of applying our bionic structural optimization on a beam. At the beginning of the optimization the load is applied on the feature surface F6 in the direction of y axis (Fig. 6a). During the optimization the sharp boundary between loading area and unloaded area has disappeared (Fig. 6b). With our new method a loading domain and a fixation domain are created for applying boundary conditions (Fig. 6c). The result of FEM analysis shows a very equal stress distribution which is reasonable according to the theory of structural optimization [5\u20137, 11]. Therefore, with the ORC method we can overcome the limits of feature surfaces to make our structural optimization method much more robust. This paper introduces a new method called Overlapping Region Concept to extend Matlab\u2019s PDE Toolbox for applying free-form surface load during FEM analysis. In our institute we have integrated the PDE Toolbox into our self-developed SG-Library in Matlab to implement and develop some bionic methods like CAO and SKO to optimize the structure of our medical robots and mechanisms" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000209_tac.2009.2020643-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000209_tac.2009.2020643-Figure7-1.png", + "caption": "Fig. 7. Torsion of the strip-like region.", + "texts": [ + " Firstly, we construct a suitable vector field in the plane such that for all by five steps as follows. Step 1) It should be pointed out that a series of smooth curves without any singular point (i.e. the curve which degenerates into a point) in the plane can be defined by a corresponding ordinary differential equation , for all , and vice versa (see [11]). Therefore, in next steps, we will construct a series of suitable smooth curves in the plane rather than construct the vector field directly. Step 2) We find a series of smooth curves (8) which cover a strip-like region as shown in Fig. 7. It is not difficult to conceive that we can twist the strip-like region such that it rotates like Archimedean spiral as shown in Fig. 7, i.e. there is a differential homeomorphism which transforms to or . Therefore, every curve in defined by (8) is transformed a corresponding curve in or . Since , we have given a series of smooth curves without any singular point which can cover the plane entirely. Step 3) We choose a trajectory of the following system: (9) and a point on . Next, we choose another trajectory of the system (9) and a point on as shown in Fig. 8. Now, we can find a curve smoothly connecting the trajectory and on the point and such that the vector field of (9) on will go into the outer side of the curve which consists of and the positive semi-trajectories of and with initial points and . The two disjoint sides separated by are denoted as and , respectively. The side including the origin is , and the other is as shown in Fig. 8. Step 4) Similar to Step 2, we can twist and properly transversely deform (namely, the direction of abscissa axis) the striplike region in Fig. 7 to become the region in Fig. 8. Obviously, the curve in the strip-like region defined by (8) cannot be distorted to a point under the proper transverse deformation. Then every curve in can be transformed to a corresponding curve in under the above metamorphism. Step 5) Similar to Step 4, we can transform in Fig. 7 to the region Fig. 8 such that the curves in defined by (8) are also transformed to the corresponding curves in . Therefore, a series of curves have been defined to cover the plane entirely. We further remark that this is possible, since smooth functions are not rigid, see [12], [13]. Hence, for any point in the plane , we have already defined a smooth curve passing through it. By Step 1, actually we also define a nonzero vector field in the plane. Let . Then for all . Now, we can consider the following control system: (10) where is the same as the foregoing definition" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002839_j.mechmachtheory.2018.06.021-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002839_j.mechmachtheory.2018.06.021-Figure7-1.png", + "caption": "Fig. 7. Mathematical sketch for establishing an outer cone CVT.", + "texts": [ + " Nomenclatures \u03b8 half cone-angle of roller \u03b3 titling angle of roller D distance ( Fig. 3 ) r radius of circle ( Fig. 3 ) C cone apex P arbitrary point ( Fig. 3 and Fig. 5 ) v movement speed of the cone apex \u03c9 angular speed of axis t time parameter \u03bb individual parameter that represents the length of CP A and B contact points on input disc and output disc A\u2019 and B\u2019 foot points ( Fig. 6 ) A\u201d and B\u201d foot points ( Fig. 6 ) \u03c9 i and \u03c9 o rotational speeds of input disc and output disc \u03c9 r rotational speed of roller i speed ratio R length ( Fig. 7 ) \u03c9 spin spin speed \u03c9 dg rotational speeds of driving component \u03c9 dn rotational speeds of driven component \u03b8dg angle between the contact area and the rotational axis of the driving component \u03b8dn angle between the contact area and the rotational axis of the driven component \u03c3 spin spin ratio \u03c9 spinT and \u03c3 spinT spin speed and spin ratio of half-toroidal CVT \u03c9 spinL and \u03c3 spinL spin speed and spin ratio of logarithmic CVT \u03c9 spinI and \u03c3 spinI spin speed and spin ratio of inner cone CVT \u03c9 spinO and \u03c3 spinO spin speed and spin ratio of outer cone CVT e the distance of the toroidal cavity from the disc rotation axis k aspect ratio Cr in creep coefficient at input contact point Cr out creep coefficient at output contact point \u03bcin tangential force coefficient at input contact point \u03bcout tangential force coefficient at output contact point \u03c7 in spin momentum coefficient at input contact point \u03c7out spin momentum coefficient at output contact point t in input traction coefficient t out output traction coefficient \u03b7 total efficiency direction at the point of contact [17,18] ", + " (29) By substituting Eq. (29) into Eq. (28) , we can derive i = \u03c9 i \u03c9 o = cos (\u03b8\u2212\u03c9t) sin \u03b8 cos (\u03b8+ \u03c9t) sin \u03b8 = cos (\u03b8 \u2212 \u03c9t) cos (\u03b8 + \u03c9t) , (30) where i is the speed ratio, as defined above. Thus, Eq. (30) is the formula of the speed ratio of the inner cone CVT without considering the slip. In this section, the parameters v and \u03c9 will be discussed separately and replaced with particular expressions. The influ- ence of the v / \u03c9 ratio to the convexity\u2013concavity of the envelope can then be avoided. In Fig. 7 , an outer cone roller is placed into the plane coordinate system XOY . Here, B is the assumed contact point on the roller, while O\u2019B (perpendicular to the generatrix of the roller) intersects X -axis at point O\u2019 . Furthermore, R represents the length of the segment O\u2019B . We suppose that the movement of the roller shown in Fig. 7 is the same as that shown in Fig. 5 (a). In Eq. (13) , the values of v and \u03c9 are assumed to be v = R t cos (\u03b8 \u2212 \u03c9t) (31) and \u03c9 = 1 t . (32) By substituting Eqs. (31) and (32) into Eq. (13) , we can obtain { x = R cos (\u03b8\u2212\u03c9t) + R cos (\u03b8\u2212\u03c9t) sin (\u03b8 \u2212 \u03c9t) cos (\u03b8 \u2212 \u03c9t) y = R cos (\u03b8\u2212\u03c9t) cos 2 (\u03b8 \u2212 \u03c9t) . (33) Thus, { x = R sin (\u03b8 \u2212 \u03c9t) + R cos (\u03b8\u2212\u03c9t) y = R cos (\u03b8 \u2212 \u03c9t) . (34) Eq. (34) represents a family of circular curves, and it can be illustrated more clearly by calculating ( x \u2212 R cos (\u03b8 \u2212 \u03c9t) )2 + y 2 = R 2 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000928_12.978271-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000928_12.978271-Figure6-1.png", + "caption": "Figure 6: Convex lens effect of the oil flow Figure 7: Light cover on the receiver head", + "texts": [ + " When the oil flow is close to the transmitter, the oil will decrease the light amounts reaching the receiver. When the oil flow moves from the middle point toward the receiver, and the light amount reaching the receiver begins to increase, in one position the value reaches the maximum, as shown in figure 5. The reason is that oil flow forms a shape of convex lens in the cross section, when the scattered light cross the oil flow, the oil flow can concentrate the light to the receiver as shown in figure 6. This effect causes the increase of the light amount reaching the sensor receiver. There are probably other oil patterns in actual application, which makes it complex for light reflection and refraction. For the bearing structure, the surfaces of the parts are fine polished and have good light reflection. Therefore, during the bearing running, some parts properly reflect the light randomly. Considering the complex light reflection and refraction of the oil, there are probably some unexpected lights captured by the receiver" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003639_iet-cta.2019.0444-Figure14-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003639_iet-cta.2019.0444-Figure14-1.png", + "caption": "FIGURE 14. Layout of the proposed dual-band coupled-line balun.", + "texts": [ + " Due to the addition of two stubs Z5, back-feeding approach is taken for the convenience of measuring, and the corresponding photographs of the fabricated circuit are shown in Fig. 13. The impedances of source and load are fixed as 50 . Applying Equations (6, 8, 11, 12, and 15), the initial design parameters for this dual-band balun are calculated to be: R0 = 81.2 , Ze1 = 72 , Zo1 = 60 , Ze2 = 89.2 , Zo2 = 74.3 , Ze3 = 126.7 , Zo3 = 67.3 , Ze4 = 84.5 , Zo4 = 70.4 , Z5 = 85.2 , Z6 = 100 , Z7 = 51.1 , Z8 = 34.8 , and \u03b8 = 50\u25e6 at 1.0 GHz. After several optimization iterations, convert these electrical parameters into physical dimensions defined in Fig. 14. The final dimensions in Fig.14 are (unit: mm): L1 = 25.4, W1 = 1.1, S1 = 1.2, L2 = 26.5, W2 = 0.7, S2 = 1.3, L3 = 25.0, W3 = 0.4, S3 = 0.3, L4 = 25.1, W4 = 0.8, S4 = 1.2, L5 = 23.5, W5 = 1.7, L6 = 25.7, W6 = 0.6, L7 = 26.0, W7 = 0.4, L8 = 34.8, W8 = 1.7, L9 = 24.6, and W9 = 3.0. Finally, the effective size of this fabricated dual-band balun is 165\u00d7 78 mm2. The full-wave EM simulation and optimization are executed by High Frequency Structure Simulator (HFSS) and the scattering parameters are measured by the Agilent Vector Network Analyzer E5230C" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002349_s12289-017-1388-x-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002349_s12289-017-1388-x-Figure8-1.png", + "caption": "Fig. 8 Hill\u2019s slip line flat indenter solution [15] (contact geometry)", + "texts": [ + " In the following paragraph of the paper, the force model utilized in the validation case as well as in all the study cases is presented and explained, highlighting the importance of the projection of the contact arc length in the calculation of one of the force model parameter, namely the pressure factor. Slip line force model As concerns the model for the calculation of the forming force, authors\u2019 have chosen to utilize the Hill\u2019s slip line flat indenters force model detailed in [15] and developed for the case of a flat punch indenting on a flat surface, Fig. 8. Since this model is based on a straight contact surface between the workpiece and the tool, by considering the projection of the contact arc, and not its curved geometry, the contact assumption is fulfilled. According to Hill\u2019s slip line model [15], the force is proportional to a pressure factor, here expressed by the symbol \u03b3, and which represents the state of stress in the plastically deformed region. The pressure factor is proportional to the ratio s/Lc, namely the ratio between the thickness of the ring slice undergoing the deformation and the projection of contact length, assumed to be a straight line" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003124_2019-01-0028-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003124_2019-01-0028-Figure8-1.png", + "caption": "FIGURE 8 Drag torques of different oil films with spray lubrication", + "texts": [ + " Thus, the fluid along\u00a0z-axis is divided into full oil film in the continuous section and mist oil film in the blocked section (Figure 7). The boundary of the two sections is calculated by the equality of the initial flow rate and the flow rate of centrifugal force component (Equation (15)). Q Q z h nz m h mnz a c a a = ( ) = - + -\u00e6 \u00e8 \u00e7 \u00f6 \u00f8 \u00f7 + + \u00e6 \u00e8 \u00e7 \u00f6 \u00f8 \u00f7 0 0 0 2 22 1 2 2 2 2p w w w wD D . (15) The drag torque is estimated under consideration of fluids film contact with both friction surfaces. The whole drag torque is then the sum of the drag torque contribution due to full oil film and mist oil film (Figure 8). T T Tsyn v d= + , (16) T nz m h h nz m h nz m v z VF a a = - + -\u00e6 \u00e8 \u00e7 \u00f6 \u00f8 \u00f7 +( ) - + -( ) - - +( )\u00e9 \u00f2\u00f2 1 2 2 0 2 0 0 p m w w wD\u00eb \u00f9\u00fbdzdq , (17) T nz m h h nz m h nz d z bcos DF a a = - + -\u00e6 \u00e8 \u00e7 \u00f6 \u00f8 \u00f7 +( ) - + -( ) - - + \u00f2\u00f2 1 2 2 0 2 0 pa m w w wD m dzd( )\u00e9\u00eb \u00f9\u00fb q . (18) At low rotational speeds the centrifugal force has little effect on the flow rate, so that the fluid can flow through the whole gap length and the mist film does not exist. With increasing rotational speeds, the full film is gradually blocked by the increasing centrifugal force" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000263_robio.2010.5723411-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000263_robio.2010.5723411-Figure1-1.png", + "caption": "Fig. 1. Nine-link biped model", + "texts": [ + " To this end, first a stable walking path is planned for hip and ankle joints and based on that, joint angle trajectory is planned. The consumed torque is calculated using inverse dynamics. This paper consists of six sections. Dynamic modeling is discussed in second section. The third section describes the method of path generation. The indices of energy, power and torque are introduced in forth section. Fifth section presents simulation results followed by some concluding remarks in last section. Let us consider a nine-link biped with toe-joint as represented in Figure 1. Some features of the model and hypothesis are as follow: - Model is 2D and moves in sagittal plane. - All links are rigid and each two adjacent links are connected together with a revolute joint. - All eight joints of model are actuated with colocated actuators. - Both single support phase (SSP) and double support phase (DSP) are considered. - Adequate friction to prevent slipping is provided by the ground surface. The system has eight degrees of freedom, which is described by generalized coordinates vector: (1) where q1-q8 are relative angles between each two links and are shown in Figure 1. One gait step includes two single support, SSP, and double support, DSP phases. B 978-1-4244-9318-0/10/$26.00 \u00a9 2010 IEEE 697 The SSP starts when toe of the rear leg leaves the ground and ends when it again comes in contact with the ground. The equation of motion in this phase is: (2) where D is symmetric inertia matrix, C is coriolis and centripetal matrix, G is gravity vector and is torque vector. The DSP starts at the end of SSP, when the swing leg comes in contact with the ground and ends at the beginning of next SSP, when the other leg, stance leg, starts to detach from the ground", + " (3), to obtain where 6, 8 and : are as follows: * > : ;*2+3 < (7) In this section, two paths are specified for hip and swing ankle joints (Figure 2). To have smooth motion for hip joint, a trigonometric function is used as follow [5]: ?@ A % BCD % BEFG % BGFH I@ JCD K % LMN \"DOP A(Q BGR BCR % JC % AS (8) where Ls is one step length, T is one step period, Lfot is foot length, Ltoe is toe length, Lth is thigh length, Lsh is shank length, hs is vertical width of hip path curve and tl is a tolerance inserted in yH equation to prevent the model from singularity condition in knee joint. In this research torso in Figure 1 remains vertical during walking. This way Eq. 8 can be used as path of the center of mass with a vertical shift. Trajectory of swing ankle is considered to have two sections related to double support and single support phases. During DSP toe angle value (q1) begins from an initial value and linearly up to zero at the end of this phase. Then the ankle tracks a circular arc with a variable radius: ?TU %BC % BT LMN VT ITU BT NWX VT (9) where subscript \u2018d\u2019 refers to DSP, La and YZ are respectively the variable radius and its angle with the horizon, shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001997_iccre.2017.7935043-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001997_iccre.2017.7935043-Figure1-1.png", + "caption": "Fig. 1: AMB system", + "texts": [ + " As a result, the GS controller based on parameter-dependent Lyapunov function is derived without using complicated methods such as descriptor representation, LFT and the sum-ofsquares technique. The effectiveness of proposed GS controller is illustrated by its comparison robust LQ (R-LQ) controller in simulations. The notation E O stands for positive definite matrix. The notation He{E} stands for E + ET . In this section, a state-space representation of an AMB system is derived. The AMB system levitates and supports a rotor without contact by using magnetic force. The schematic diagram of the system is shown in Fig. 1. In this study, the AMB system has the four degrees of freedom. The system consists of a rotor, four pairs of electromagnets and two gap sensors attached at each end. The shape of the rotor is assumed as a circular cylinder. A flywheel is attached on the center of the rotor. The pair of electromagnets face to each other and gain scheduling, active magnetic bearing, linearKeywords- 978-1-5090-3774-2/17/$31.00 \u00a92017 IEEE generate the levitation force corresponding to control input. The levitation force is generated in vertical and horizontal direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003852_j.mechmachtheory.2019.103628-Figure10-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003852_j.mechmachtheory.2019.103628-Figure10-1.png", + "caption": "Fig. 10. 3D model of the gear.", + "texts": [ + " It is necessary to re-check the interference of non-working side according to the above method until there is a reasonable gap between the non-working side of grinding wheel and the opposite tooth surface. In order to verify the above methods, a hypoid gear set, shown in Table 1 , is used as an example for grinding simulation process. The gear is machined by Formate\u00ae face hobbing process. The machine setting parameters of gear cutting are shown in Table 2 . The mesh coordinates of both sides of the tooth surface can be obtained by solving the tooth surface equation in Section 3 , and the 3D model of the gear is shown in Fig. 10 . The calculated mesh coordinates of the tooth surface are compared with the coordinates calculated by the hypoid gears design and machine setting calculation program [21] . Taking the concave side as an example, the maximum tooth deviation (shown in Fig. 11 ) is less than 0.25 \u03bcm, which may be due to rounding errors of input parameters. It can be considered that the two tooth surfaces are consistent. According to the method described in Section 4 , the curvature characteristics of concave and convex sides are analyzed", + " According to the motion principle of CNC hypoid grinding machine shown in Fig. 7 , a five-axis tooth grinding simulation platform for the FFHHG is constructed, as shown in Fig. 16 . The distance R 0 is 350 mm, L B and L M are 102.538 mm and 65 mm, respectively. Then the grinding wheel location of tooth grinding can be calculated according to the method described in Section 5 , which are shown in Figs. 17 and 18 . At the same time, the grinding wheel model (the green part) is built according to Table 4 , and the imported gear model (the orange part) is shown in Fig. 10. There is no feeding during the first grinding, and the grinding wheel should be tangent to the tooth surface theoretically. The actual ground tooth surfaces are shown in Fig. 19(a) , which shows that the grinding wheel and the tooth surface are in intermittent contact state, and the ground area extends from the top of the tooth to the root of the tooth evenly at the contact positions. This is because the material removal simulation of machining simulation software is based on the discrete grid model, but not on Boolean operation of accurate 3D model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000777_s0025654411040017-Figure14-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000777_s0025654411040017-Figure14-1.png", + "caption": "Fig. 14. Fig. 15.", + "texts": [], + "surrounding_texts": [ + "By setting \u03b4 = 0 in Eqs. (3.6) and (3.7) and by introducing new variables z1 and z2 by formulas z1 = z + a1\u03d5 and z2 = z \u2212 a2\u03d5, we obtain the equations z\u2022\u20221 + a11(c1z1 + \u03b31z \u2022 1) + a12(c2z2 + \u03b32z \u2022 2) = 0, z\u2022\u20222 + a21(c1z1 + \u03b31z \u2022 1) + a22(c2z2 + \u03b32z \u2022 2) = 0, (4.1) MECHANICS OF SOLIDS Vol. 46 No. 4 2011 a11 = 1 + a2 1 \u2212 a1f1, a12 = 1 \u2212 a1a2 \u2212 a1f2, a21 = 1 \u2212 a1a2 + a2f1, a22 = 1 + a2 2 + a2f2. (4.2) The dimensionless variables ak and fk (k = 1, 2) in formulas (4.2) are the ratios of the corresponding variables introduced in the preceding sections to the radius of inertia \u03c1; i.e., we assume that the radius of inertia \u03c1 is equal to unity. The variables z1 and z2 have the following physical meaning. These are small deviations of the points O1 and O2 from their equilibrium positions described in Sec. 3. Equations (4.1) and (4.2) and relations (1.1) with the use of the Lyapunov theorem permit solving the above stability and instability problem for any fixed v. Let us write out the characteristic polynomial of the linear system (4.1) as p4 + p3(a11\u03b31 + a22\u03b32) + p2(a11c1 + a22c2 + \u03b31\u03b32\u0394) + p(\u03b31c2 + \u03b32c1)\u0394 + c1c2\u0394 = 0, (4.3) where \u0394 = a11a22 \u2212 a12a21. Applying the Routh\u2013Hurwitz theorem to the polynomial (4.3), we arrive at the following result. Proposition 1. For the asymptotic stability of the equilibrium in the case of rectilinear and uniform motion of the vehicle, it is necessary and sufficient that the rolling friction coefficients f1 and f2 (and hence the velocity v of the vehicle translational motion determined by (1.1)) satisfy the inequalities a11\u03b31 + a22\u03b32 > 0, a11c1 + a22c2 + \u03b31\u03b32\u0394 > 0, \u0394 > 0, (4.4) (a11\u03b31 + a22\u03b32)[a11\u03b32c 2 1 + a22\u03b31c 2 2 + \u03b31\u03b32\u0394(\u03b31c2 + \u03b32c1)] \u2212 (\u03b31c2 + \u03b32c1)2\u0394 > 0, \u0394 = a11a22 \u2212 a12a21 = a(a + f2 \u2212 f1), a = a1 + a2. (4.5) The dimensionless parameters aij (i, j = 1, 2) are determined by formulas (4.2). Consider several particular cases where Proposition 1 applies. Proposition 2. If the system is completely symmetric, i.e., if \u03b3k = \u03b3 = 0, ck = c, fk = f, k = 1, 2, a1 = a2 = a0, (4.6) then inequalities (4.4) hold for each f ; i.e., the asymptotic stability of the vehicle equilibrium considered above is preserved for any wheel rolling friction coefficients and hence for any velocities of its translational motion. The proof is by a straightforward verification of inequalities (4.4) under conditions (4.6). In this case, it follows from (4.5) that \u0394 = a2 > 0 and the first two inequalities in (4.4) are satisfied, because they are reduced to the following ones: 2\u03b3(1 + a2 0) > 0, 2c(1 + a2 0) + \u03b32a2 > 0. The last inequality in (4.4) has the form c2(1 \u2212 a2 0) 2 + 4\u03b32c2a2 0(1 + a2 0) > 0. Thus, the stability conditions (4.4) are satisfied for any f \u2265 0, and the proof of Proposition 2 is complete. Proposition 3. Assume that all conditions (4.6) except for the last are satisfied; i.e., a1 =a2 (the case of incomplete symmetry; i.e., the car body center of mass is displaced). Then, for a1 < a2, the stability conditions (4.4) are satisfied for every f \u2265 0, and for a1 > a2, the stability is violated for f > f0 = 1 a1 \u2212 a2 [ 2 + a2 1 + a2 2 + \u03b5(a1 + a2)2 \u2212 \u221a \u03b52(a1 + a2)4 + 4(a1 + a2)2 ] > 0, \u03b5 = \u03b32 c . (4.7) Remark. The quantity f0 in inequality (4.7) can be arbitrarily small for some values of the system parameters. For example, as \u03b5 \u2192 0 (small damping or large rigidity of the wheel) and for a1a2 = 1 and a1 > 0, we have f0 = (a1 \u2212 1)(1 + a2 1)/(a1 + a2 1), which can be arbitrarily small as a1 \u2192 1. MECHANICS OF SOLIDS Vol. 46 No. 4 2011 The proof is by a straightforward verification. Under the conditions given in Proposition 3, the instability inequalities (4.4) become \u03b3[2 + a2 1 + a2 2 + (a2 \u2212 a1)f ] > 0, c[2 + a2 1 + a2 2 + (a2 \u2212 a1)f ] + \u03b32(a1 + a2)2 > 0, \u0394 = (a1 + a2)2 > 0, x2 + 2x\u03b5(a1 + a2)2 \u2212 4(a1 + a2)2 > 0, x = 2 + a2 1 + a2 2 + (a2 \u2212 a1)f. (4.8) For a2 > a1, these inequalities are necessarily satisfied for any f > 0. (For the first three inequalities in (4.8), this is obvious, and for the last, this follows from the inequality x > 2 + a2 1 + a2 2 > 2(a1 + a2).) But if a2 < a1, then the first inequality in (4.8) is violated for f > f1 = (2 + a2 1 + a2 2)/(a1 \u2212 a2), the second is violated for f > f2 > f1, the third is always satisfied, and the last is violated for f \u2208 [f0, f00], where f0 is given by the formula in (4.7) and f00 is given by the same formula with the plus sign at the radical. It is easily seen that 0 < f0 < f1 < f00, which implies that all inequalities in (4.8) are violated for f > f0. The proof of Proposition 3 is complete. Proposition 4. Let f2 = 0, \u03b30 = 0, and f1 = f > 0; i.e., assume that the fore (driving) wheel does not experience any rolling resistance and there is no dissipation in it. In this case, if a1a2 < 1, then the instability arises for f > a1, and if a1a2 > 1, then the instability arises for f > (a1a2 \u2212 1)/a2. Proposition 5. Let f1 = 0, \u03b31 = 0, and f2 = f > 0; i.e., assume that the rear (driven) wheel does not experience any rolling resistance and there is no dissipation in it. In this case, if a1a2 < 1, then the instability arises for f > (1 \u2212 a1a2)/a1, and if a1a2 > 1, then the system is stable for any f > 0. The proof of Propositions 4 and 5 is by a straightforward verification of inequalities (4.4) in Proposition 1 under the assumptions stated above. Let \u03b31 > 0, \u03b32 = 0, and \u03b32 > 0. Then the following assertion holds. Proposition 6. Under the above conditions, the asymptotic stability domain D1 in the plane of the parameters (f1, f2) is given by the following inequalities. If 0 < a1a2 < 1, then 0 < f2 < 1 \u2212 a1a2 a1 , 0 < f1, c1a1f1 \u2212 c2a2f1 < c0, f1 \u2212 f2 < a1 + a2. (4.9) If 1 < a1a2 < \u221e, then 0 < f1 < a1a2 \u2212 1 a2 , 0 < f2, c1a1f1 \u2212 c2a2f2 < c0, (4.10) c0 = c1(1 + a2 1) + c2(1 + a2 2). (4.11) The proof of Proposition 6 is by a straightforward verification of inequalities (4.4) under the restrictions given in the statement. Note that if one of the inequalities in conditions (4.10), (4.11) is not satisfied, then the vibrations under study are unstable. Now let \u03b31 = \u03b32 = 0; i.e., assume that there is no dissipation in the wheels. In this case, Eq. (4.3) is biquadratic, the stability of system (4.1) can be only nonasymptotic, and the conditions for such stability are equivalent to the conditions that this biquadratic equation has only pure imaginary roots. These considerations lead to the following result. Proposition 7. For \u03b31 = \u03b32 = 0, the stability domain D0 in the plane of parameters (f1, f2) is given by the following inequalities: f1 > 0, f2 > 0, c0 > c1a1f1 \u2212 c2a2f2, a1 + a2 > f1 \u2212 f2, [c0 \u2212 (c1a1f1 \u2212 c2a2f2)]2 > 4c1c2(a1 + a2)(a1 + a2 \u2212 f1 + f2). (4.12) The proof of Proposition 7 follows from well-known necessary and sufficient conditions for the existence of pure imaginary roots of a biquadratic equation. Proposition 8. The inclusion D1 \u2282 D0 holds. MECHANICS OF SOLIDS Vol. 46 No. 4 2011 The proof of Proposition 8 is by a straightforward substitution of the extreme values in the linear inequalities (4.9) or (4.10) into inequalities (4.12) and verification that the latter are satisfied. Remark. Proposition 8 means that this nonsymmetric addition of dissipation narrows down the stability domain in the plane of the parameters (f1, f2). Graphically (schematically), the domains D1 and D0 in various possible cases are shown in Figs. 2\u201315, where digit 1 corresponds to the line f1 \u2212 f2 = a1 + a2, digit 2 corresponds to the line c1a1f1 \u2212 c2a2f2 = c0 = c1(1 + a2 1) + c2(1 + a2 2), digit 3 corresponds to the line f2 = |1 \u2212 a1a2|/a1, digit 4 corresponds to the line f1 = |1 \u2212 a1a2|/a2, digit 5 corresponds to the line f1 \u2212 f2 = \u2212|1 \u2212 a1a2|(a1 + a2)/(a1a2), and digit 6 corresponds to the parabola (c0 \u2212 c1a1f1 + c2a2f2)2 \u2212 4c1c2(a1 + a2)(a1 + a2 \u2212 f1 + f2) = 0. Figures 2, 4, 6, 8, 10, 12, and 14 correspond to the case of \u03b31 = \u03b32 = 0, while Figs. 3, 5, 7, 9, 11, 13, and 15 correspond to the case of \u03b31 = 0, \u03b32 = 0 or \u03b31 = 0, \u03b32 = 0 for all possible distinct values of the parameters a1, a2, c1, and c2. Now consider the case of weighted wheels, m1 = 0 and m2 = 0. Since we assume that the car body mass is m = 1, the quantities m1 and m2 are the respective dimensionless ratios of the wheel masses to the car body mass. Simple calculations show that the form of the vibration equations (4.1) remains the same and the coefficients aij (i, j = 1, 2) are determined by the functions a11 = \u03c12 \u2212 (a1 + am2)(\u2212a1 + f1) \u03c12 0 , a12 = \u03c12 \u2212 (a1 + am2)(a2 + f2) \u03c12 0 , a21 = \u03c12 + (a2 + am1)(\u2212a1 + f1) \u03c12 0 , a22 = \u03c12 + (a2 + am1)(a2 + f2) \u03c12 0 , where \u03c12 0 = m0\u03c1 2 and m0 = 1 + m1 + m2. Then one can readily compute that \u0394 = a11a22 \u2212 a12a21 = a(a + f2 \u2212 f1)/\u03c12 0. Proposition 1 remains valid under the above variations in aij , i, j = 1, 2, and \u0394. In a similar way, one can restate Propositions 2\u20138. Equations (4.1) and (4.2) describing the vehicle vibrations under the action of the rolling friction forces admit the following analogies (interpretations) to well-known systems in mechanics. Each of the coefficient matrices multiplying zk and z\u2022k (k = 1, 2) in (4.1) with (4.2) taken into account consists of MECHANICS OF SOLIDS Vol. 46 No. 4 2011 two parts. The first part (it is said to be unperturbed) does not contain the friction coefficients f1 and f2. The second (perturbed) part depends on these coefficients linearly. As follows from the above derivation of the equations of motion, the origin of the perturbed part significantly depends on the following two external causes. One of them is the interaction between the deformable wheels and the rough road, which causes the rolling resistance torques Mk1 and Mk2; the second cause is the torque M1 from the engine (external energy source) to the driving wheels ensuring their (and the driven wheels) permanent rolling if there are resistance torques mentioned above. If these causes are absent, then the system motion occurs for f1 = f2 = 0, and Eqs. (4.1) contain only the first parts of the above-mentioned matrices for which the solutions have the property of (asymptotic) stability (a conservative system MECHANICS OF SOLIDS Vol. 46 No. 4 2011 under the action of some forces with complete or incomplete dissipation). But if f2 1 + f2 2 = 0, then these causes result in the appearance of additional terms in Eqs. (4.1), which, according to [6], can be interpreted as positional (circular) forces (for the corresponding terms with z1 and z2); gyroscopic and nonconservative (depending on the velocities) forces (for the corresponding terms with z\u20221 and z\u20222); conservative positional forces (for the corresponding terms with z1 and z2); and dissipative (Rayleigh) MECHANICS OF SOLIDS Vol. 46 No. 4 2011 MECHANICS OF SOLIDS Vol. 46 No. 4 2011 forces (for the corresponding terms with z\u20221 and z\u20222). In the general case, such a decomposition of forces was performed in [7]. Thus, the existence of the above-mentioned external causes leads to the appearance of additional forces of the described structure, whose destructive character is well known in mechanics of gyroscopic systems, structural mechanics (in particular, in mechanics of bridges), and other adjacent fields; it is described in the literature [8\u201310]. Even if there forces do not actually lead to the vehicle instability, they still significantly change the frequencies of its vibrations in the vertical plane, and this fact cannot be ignored when solving the corresponding problems. These facts were already considered earlier by the authors in the paper [11]." + ] + }, + { + "image_filename": "designv11_33_0001375_j.fusengdes.2011.01.062-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001375_j.fusengdes.2011.01.062-Figure2-1.png", + "caption": "Fig. 2. Schematic diagram of the hybrid robot.", + "texts": [ + " The robot is composed of a 4 degrees of freedom (DOF) multi-link serial mechanism (named as Carriage) serially connected to a standard 6-DOF Stewart parallel mechanism (named as HexaWH), which aims to arrive at a compromise between a high stiffness of parallel manipulators and a large workspace of serial manipulators. In what follows, we first derive a nominal kinematic model for the proposed robot. Thereafter, based on the nominal model, a related identification model including unknown parameters is developed. 1864 Y. Wang et al. / Fusion Engineering an 2 s W r a 0 t l w m f m .1. Kinematic model The schematic diagram of the redundant hybrid manipulator is hown in Fig. 2. The connection platform frame {4} of the Hexa- H is coincident with the end-effector of the Carriage. The global eference frame {0} is located at the left rail of the Carriage. Based on this hybrid structure, a vector-loop equation is derived s: P5 = 0P4 + 0R4 4P5 = 0P4 + 0R4(lili + 4ai \u2212 4R5 5bi) = 0P4 + 0R4lili + 0R4 4ai \u2212 0R5 5bi (1) From Eq. (1), the nominal leg length, i.e., the inverse solution of he robot can be expressed as: ili = (0R4) \u22121 (0P5 \u2212 0P4 \u2212 0R4 4ai + 0R5 5bi) (2) here 0P5 and 0R5 are the nominal position vector and rotation atrix of the end-effector frame {5} with respect to the fixed base rame {0}" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002149_gt2017-63495-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002149_gt2017-63495-Figure3-1.png", + "caption": "Figure 3 Schematic of a pre-loaded three-pad AFB", + "texts": [ + "org/about-asme/terms-of-use ( , ) ( ) cos ( )sin ( , ) cos( ) j j j j j j p p h z C X Z Y Z w z r (3) , where C represents the bearing radial nominal clearance and ( , )j jw z is the elastic deflection of the underlying support structure, calculated following the independent elastic foundation model presented in [16]. pr is the hydrodynamic preload of each top foil pad. The hydrodynamic preload is the difference between the nominal clearance C and the minimum clearance, mC , i.e., p mr C C as illustrated in Figure 3. The hydrodynamic preload is achieved by making the bearing sleeve with non-circular contour with three lobes (Figure 3). The hydrodynamic preload results in different clearance along the circumferential direction with larger clearance at the leading and trailing edges. It should be noted that the hydrodynamic preload is different from a mechanical preload which is commonly generated by the manufacturing error of the foils. Typically, overall curvature of the foils (upon following a typical forming/heat treatment process) does not match to the sleeve and foils do not ideally sit on the sleeve. The mechanical preload that comes from the aforementioned manufacturing issue also exists in the single pad foil bearings designed with uniform assembly clearance" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001718_s11431-011-4499-5-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001718_s11431-011-4499-5-Figure5-1.png", + "caption": "Figure 5 The 3-PUU mechanism.", + "texts": [ + " Link group JKML is in loop II, all of its axes are parallel to z-axis, it can not rotate around x-, y-axes and its common constraint should be gz IIm (0, 0 0 0)=2, X IIm = 11,4 Im (, 0 0 0)+ gz IIm (0, 0 0 0)= X IIm (0, 0 0 0)2, IIF = IIP i iII pn 16 + X IIm = 6\u00d73 5\u00d74+2=0, F= FI+FII =3+0=3. The mobility can be obtained from eq. (10): F=6\u00d7105\u00d712+(1+2)=3. The total number of over constraints in the mechanism is L j X jm1 = X Im + X IIm = 1+2=3, the sum of the common constraints in the two loops is L j jm1 = Im + IIm =1+1=2, thus, v = L j X jm1 L j jm1 =1, there is only one virtual constraint in the mechanism. Example 3. Figure 5 illustrates the 3-PUU mechanism [22] with two independent loops. The coordinate system O-xyz is attached to the frame link J, the axes A1, A2 and A3 are parallel to plane xOy, and axis B1 is parallel to x-axis. Axis Bi is perpendicular to axis Ai (i=1,2,3). Similarly, Axis Ci is perpendicular to axis Di (i=1,2,3). Axes Ai and Bi form a universal pair (U-pair), the same as axes Ci and Di. Loop I consists of link groups 1 and 2, it has three rotations and three translations(caused by one prismatic pair and two revolution pairs)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001147_demped.2011.6063611-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001147_demped.2011.6063611-Figure8-1.png", + "caption": "Figure 8. Alternator Test Bench", + "texts": [ + " There exists a repeating pattern of a large ripple followed by two small ripples, which agree well with the analytical analysis in section III. When the short circuit in the simulation is removed, the alternator output current is shown in Fig. 7. The repeating pattern is not obvious any more. Therefore, the harmonic component at 1/3 of the rectified ripple frequency is verified to be the fault signature for stator turn fault. V. EXPERIMENTAL VALIDATION The above analysis is verified in the experiment. An alternator test bench has been built in the lab and is shown in Fig. 8. The \u201cvehicle engine\u201d is simulated by an Allen Bradley servo motor. Controlled by the PC, the servo motor drive is able to simulate complex load profiles within a wide speed range. The servo motor is coupled to the alternator through a pulley-serpentine belt assembly. The alternator is electrically loaded with several high-current power resistors and one leadacid battery. The alternator is modified so that four taps are attached to the four adjacent turns in the stator winding, as shown in Fig. 9" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003917_0954406219892301-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003917_0954406219892301-Figure1-1.png", + "caption": "Figure 1. Specimen dimensions in accordance to ISO 1143 standard.", + "texts": [ + " The experimental work was conducted on the EOSINT M280 SLM device, equipped with 200W continuous wave Ytterbium laser, emitting 0.2032mm thick and 1064nm infra-red beam, with a scan speed of 7000mm/s in nitrogen environment. Device working space was 250 250mm2 with a height of 325mm. The material was supplied by EOS: maraging steel MS1 (1.2709, X3NiCoMoTi18-9-5), with nominal chemical composition given in Table 1. Specimens obtained by SLM process, used in this study were in accordance to ISO 1143 standard (Figure 1). Specimens have been built in vertical stacking direction with respect to the horizontal base plate. Specimens were detached from the base plate by wire-cut electrical discharge machining (EDM). Afterwards, specimens underwent surface cleaning by microshot-peening by 400 mm stainless steel spheres. Then, three specimens were left nitrated (N), while three were heat treated by aging up to 490 C for 6 h, as recommended by the material manufacturer. Specimens were sectioned in the longitudinal plane and the cross plane to reveal the laser fused material in two sections" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000338_6.2009-3942-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000338_6.2009-3942-Figure1-1.png", + "caption": "Figure 1. CCW/USB and CCW/OTW pneumatic powered-lift concepts (from Ref. 2)", + "texts": [ + " Based on GTRI personnel\u2019s many years of experience with pneumatic and powered-lift evaluations and development2, 3, 4, it is believed that the combination of Circulation Control Wing (CCW) with either Upper Surface Blowing (USB) or Over the Wing (OTW) propulsion systems can provide an excellent powered-lift configuration capable of very high lift generation at low blowing requirements. These configurations also have the capability of interchanging engine thrust recovery needed for takeoff and climbout with trust offset needed for powered-lift equilibrium approach and landing down steep glide slopes. These concepts are shown schematically in Figure 1. Unlike more-conventional powered-lift concepts (for example, see Ref. 1) which frequently use complex and heavy mechanical flaps, the pneumatic CCW concept with either a small simple curved flap or rounded trailing edge (TE) is used here. The flow entrainment of the CCW curved/rounded TE highly augments the wing lift by itself, producing lift augmentations as high as 80 times the input jet momentum3,4. When combined with engines located above the wing (USB or OTW), the curving CCW flow entrains and deflects engine thrust for even greater lift", + " The thrust contribution to lift here comes not so much from the deflected thrust coefficient component, CT sin ( + jet), but more from the engine exhaust velocity entrained onto the curved upper wing surface and the increased negative pressure yielding increased high lift. Associated with this is either thrust turning and offset for greater drag on approach, or thrust recovery for takeoff and climb. This interchange is accomplished not only by choice of flap angle on a very small CCW flap, but also by variation in the momentum coefficient C\u00b5 and the thrust coefficient CT. An uncertainty here is whether this small-chord highly-deflected CCW flap (Fig. 1, right) will retain its ability to entrain the high velocity engine exhaust and maintain flow attachment to the flap curved upper surface. An important aspect of the performance capability for the CCW/OTW combination will be achieved by optimization of both geometric and aerodynamic performance of this integrated CCW and powered-lift system. This will be accomplished by low-speed experimental evaluations, which are the subject of this Part I paper on aerodynamic/propulsive characteristics of these pneumatic systems" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000274_icara.2000.4803910-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000274_icara.2000.4803910-Figure6-1.png", + "caption": "Fig. 6. Transit Length", + "texts": [ + " The dotted lines from light green points express the transit length. We calculate the transit length as the sum of both transit lengths of the two wheeled when the planning path is centered at the middle of the two wheeled. b) Formula for Transit Length: We show the formula for the transit length for planning with the shortest path method on the grid based map and for planning with the steering set. Here d is the length of the axle. The transit length of the grid-based shortest path search method is computed by the summation of the dotted line in Fig. 6-a) and its formula is shown in Eq. 2. The transit length of the steering set is computed by the summation of the dotted line in Fig. 6-b), here Rn is the radius of the quadrant arc. The formula is shown in Eq. 3. TransitLengthShortestPathSearch = 2 \u00d7 Goal\u2211 n=Start Lengthn + \u2211 n=Discon.Points 2\u03c0 d 2 ( 2\u03b8n 360 ) (2) TransitLengthSteeringSets = 2 \u00d7 \u2211 n=Type StraightLengthn + 2 \u00d7 \u2211 n=Type {CurveCountsn \u00d7 CurveLengthn} (3) In the real world, some parts of the environment will change dynamically during planning. Even though the environment changes every second, it is necessary that the path for wheeled robots is smooth. Grid based shortest path search which needs a recalculation process to smooth by a high-dimensional function, and the trajectory is not guaranteed optimality and completeness" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002795_gt2018-77151-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002795_gt2018-77151-Figure1-1.png", + "caption": "Fig. 1 A geometry of a 4 pads tilting-pad journal bearing and coordinate system.", + "texts": [ + " The current paper details a TEHD analysis for modeling tilting pad bearings operating over a wide range of operating conditions that require introducing models for laminar flow operation, transition regime, and full turbulence flow conditions. The study details a full 3D turbulence flow analysis with pad deformations, thermal and mechanical pressure, to predict performance of a tilting pad bearing. Comparisons to measurements of pads\u2019 temperatures in a dedicated test rig serve to validate the model. Fig. 1 illustrates a typical geometry of a four-pads tiltingpad journal bearing and the coordinate systems for analysis. A pad clearance equals cp and the journal eccentricity has components (ex, ey). The oil films form wedge-like shapes between the pads and journal surfaces. The film thickness is \u210e(\ud835\udf03, \ud835\udc67) = \ud835\udc50\ud835\udc5d + \ud835\udc52\ud835\udc65cos\ud835\udf03 + \ud835\udc52\ud835\udc66sin\ud835\udf03 \u2212 \ud835\udc50\ud835\udc5d\ud835\udc5a\ud835\udc5dcos(\ud835\udf03 \u2212 \u0398\ud835\udc5d) \u2212 (\ud835\udc45\ud835\udc5d + \ud835\udc61\ud835\udc5d)\ud835\udeff\ud835\udc5dsin(\ud835\udf03 \u2212 \u0398\ud835\udc5d) + \ud835\udeff\ud835\udc51(\ud835\udf03, \ud835\udc67) 1(1) where p is the pad tilt angle and d denotes the surface deformation field induced by both pressure and temperature. The pad preload factor mp equals \ud835\udc5a\ud835\udc5d = 1 \u2212 \ud835\udc50\ud835\udc4f \ud835\udc50\ud835\udc5d 2(2) Dowson [24] derived the generalized Reynolds equation that considers oil viscosity changes due to a cross-film temperature variation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003059_iecon.2018.8592894-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003059_iecon.2018.8592894-Figure3-1.png", + "caption": "Fig. 3. IR distance sensor (GP2YOA31SKOF)", + "texts": [], + "surrounding_texts": [ + "Different configurations of infrafed (IR) distance sensors, GP2YOA31SKOF (SHARP(R), Japan), were used to measure foot clearance (FC) in the range of 4 to 30 cm. These IR sensors function based on the reception angle of the reflected IR beam to the IR detectors. The further the distance, the smaller the angle will be. The sensor can measure the distance without the influence of environmental temperature and the operating duration. Output of the sensor is the voltage corresponding to the detection distance. In addition, an inertia measurement unit (IMU), TSND151 (ATR-Promotions(R), Japan), is utilized to estimation foot orientation. The IMU sensor includes an accelerometer, a gyroscope, a magnetometer, and a pressure sensor. Developing system, the IMU, and the IR sensor are shown as Fig 1 - 3. The IMU is attached on the x-z plane of the world coordinate and IR sensors are attached to each side of the toe and the heel. The data of the IR sensors is gathered to STMicroelectronics (Nucleo-F446RE) once. After that, it is transmitted to Lenovo X230 (Intel(R) Core(TM) i5-3320M CPU@ 2.60 Hz). The data of the IMU is also transmitted by wire communication to the laptop. Maximum sampling frequency of the IR sensor is 20 Hz. Maximum sampling frequency of the IMU sensor is 1000 Hz for the accelerometer and the gyro sensor, 100 Hz for the magnetometer, and 25 Hz for the pressure sensor." + ] + }, + { + "image_filename": "designv11_33_0000001_s1023193510110108-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000001_s1023193510110108-Figure4-1.png", + "caption": "Fig. 4. CV curves recorded on Pt (1) and on Pt/PIn5COOH (2) electrodes in buffer solution (pH 6.5) containing 5 and 6 mM of catechol respectively. For com parison CV curve recorded on Pt/PIn5COOH (3) elec trode in pure buffer solution is shown. Sweep rate 40 mV/s.", + "texts": [ + " It indicates that the reaction rate is controlled by CT diffusion. By using Randies Sevcik equation [30] and assuming a two electron CT oxidation we determined the diffusion coefficient of CT to be 3.6 \u00d7 10\u20136 cm2 s\u20131 (from data for Pt/PIn5COOH electrode) and 3.3 \u00d7 10\u20136 cm2 s\u20131 (from data for Pt electrode). These values are two times lower than that determined by others authors [11, 31]. The comparison of CV curves recorded in the presence of CT on bare Pt with that one obtained in supporting electrolyte (buffer solution) (Fig. 4) reveals that CT is oxidized only after the oxidation of polymeric film is initiated. Presence of oxidized polymer on Pt surface enhanced oxidation current of CT values. It suggests that polymer film is permeable for CT. It means that CT penetrates the film and is oxidized not only at polymer surface but also in the bulk of polymer, so that real surface of reaction is much higher than in case of bare Pt electrode. CV curves were also recorded at Pt and Pt/PIn5COOH electrodes for different CT concentra tions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003019_ecce.2018.8558059-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003019_ecce.2018.8558059-Figure2-1.png", + "caption": "Fig. 2. Ideal Hall sensor setup for rotor position sensing in synchronous Machine", + "texts": [ + "1 shows the general block diagram of a current controlled PMSM drive. A PI control mechanism is used for controlling the current. The output of the PI is then fed to a Space Vector PWM (SVPWM) block which convert the dq voltages to phase voltages and applied to the motor. The main two block which uses the position in the current controlled IPMSM drive are a) park transformation from abc to dq block and b)PWM block. The position information may have errors which may cause non-optimal control of the PMSM drive. Fig. 2 shows the ideal placement of the Hall Effect sensor with respect to the rotating magnet. There is always a misalignment between the magnet and the sensor in practical applications. In addition, there is part-to-part variation between the sensor devices resulting in various type of errors as discussed in [1]. Other types of position sensors also will have errors due to misalignment and other manufacturing tolerances. The effect of position error on the performance of the PMSM machine due to these nonlinearities are studied clearly in [2]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003252_robio.2018.8665152-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003252_robio.2018.8665152-Figure7-1.png", + "caption": "Figure 7 Top view of the motion process(target position is a sphere with a radius of 0.02m).", + "texts": [ + " Algorithm 1 Guided Policy Search with BADMM Initialization Initialize controller t tp u x policy t tu x and t ; Build sample set S ; for iteration k=1 to K do for n=1 to N do Sample trajectory k n from t tp u x and estimate cost; Append k n to S ; Optimize t tp u x with respect to (11) ; Optimize with respect to (13) ; Update dual variables t using (6) ; end for Update dynamic model 1 ,t t tp x x u using S ; Clear S end for In this section, we verify the effectiveness of the proposed algorithm for precise position control of the PCC continuum manipulator system. The two-segment simulation robot is built in the simulation software V-REP (see Fig. 7). We have added restrictions to approximate PCC character (the second simplification method described in Section 2), which means that we can only apply control commands to 4 joints. The state of this PCC continuum robot contains the angles and angular velocities of the 4 controlled joints, and the three-dimensional position and velocity of the end-effector. We assume that the observations of the system are the full state of the system during the training and testing process, which means t to s at any time, in order to verify whether the guided policy search method can be used to control the PCC continuum robot and whether the linear time-varying Gaussian model can describe the dynamic of the PCC continuum robot", + " After convergence, the control policy will produce the same control effect as the linear Gaussian controller because the control policy has learned the control performance of the Gaussian controller. Thus, the policy can accomplish the motion planning task, controlling the continuum manipulator to the target position. Fig. 6 shows the comparison between end-effector trajectories generated by the linear Gaussian controller and by the control policy after the 20th iteration. After 20 iterations, the policy provides an effective solution for guiding the two-segment PCC continuum manipulator to the target position (see Fig.7). In the two-segment experiment, among the 24 joints in the 12 cross shafts, only the information of 4 joints is used to make up the state of the continuum manipulator (the information of the remaining 20 joints was not used, although the other 20 joints were related to those 4 joints). It is challenging to implement motion planning with partial information. In order to verify whether the control of the two-segment continuum manipulator is a coincidence, we build a three-segment PCC continuum manipulator, where the nonlinearity and coupling of the dynamics are stronger" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003872_j.matpr.2019.07.171-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003872_j.matpr.2019.07.171-Figure2-1.png", + "caption": "Fig. 2. Tool path simulation of profile and root", + "texts": [ + " Full G-code machine simulation accounts for all the motion, eliminating surprises that can appear during post processing. Machine tool simulation verifies NC programs in the context of a full motion simulation of the machine tool. This can be especially useful for complex, multi-axis, multi-headed machines such as mill/turn machining centers. The advanced version of this capability includes a machine tool builder that enables users to create new machine simulation models, complete with kinematics, using NX geometric models. Figure 2 represents the tool path simulations for profile and root of the blade. To help optimize the toolpath and check for possible errors, NX includes a CAM visualization capability for integrated material removal and toolpath simulation, with checking for collisions and gouges. Integrated verification is immediately accessible during the tool path creation process rather than output for replay at the end of the process. 1.2.3 Post Processing 5-Axis machining requires machine-control (NC) post processing to convert the cutter-location (CL) data that define the tool path data with a CAM system into the NC data that the machine can read" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003745_systol.2019.8864748-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003745_systol.2019.8864748-Figure1-1.png", + "caption": "Fig. 1: Hexarotor I PPNNPN. The red and blue arrows determine the direction of the reactive moment of each rotor", + "texts": [ + " The multirotor model is obtained by expanding (2), doing the transformations from body to inertial frame (1), and applying the previous assumptions: x\u0307 = vx y\u0307 = vy z\u0307 = vz v\u0307x = 1 m [cos(\u03c6) sin(\u03b8) cos(\u03c8) + sin(\u03c6) sin(\u03c8)]fz v\u0307y = 1 m [cos(\u03c6) sin(\u03b8) sin(\u03c8)\u2212 sin(\u03c6) cos(\u03c8)]fz v\u0307z = 1 m [cos(\u03c6) cos(\u03b8)]fz \u2212 g \u03c6\u0307 = p+ sin(\u03c6) tan(\u03b8)q + cos(\u03c6) tan(\u03b8)r \u03b8\u0307 = cos(\u03c6)q \u2212 sin(\u03c6)r \u03c8\u0307 = sin(\u03c6) cos(\u03b8) q + cos(\u03c6) cos(\u03b8) r p\u0307 = 1 Jxx [\u2212(Jzz \u2212 Jyy)qr \u2212 Jpq\u2126p + \u03c4x] q\u0307 = 1 Jyy [(Jzz \u2212 Jxx)pr + Jpp\u2126p + \u03c4y] r\u0307 = 1 Jzz [\u2212(Jyy \u2212 Jxx)pq + \u03c4z] (4) where Jp is the inertia moment of the rotor (rotating parts) and the propeller around Z axis, and \u2126p = num\u2211 i=1 \u2126i (5) where num denotes the number of rotors of multirotor and \u2126i is the angular velocity of the rotor i. In (4), the force and the moments applied to the multirotor are the inputs, as a consequence of the lift generated by the actuators. These inputs can be defined as virtual actions because the system does not have actuators that can generate these responses. For that reason, it is necessary to find the relation (Bstr) between the virtual inputs (uv) and the forces produced by the rotors (um). For example, the Bstr for a hexarotor I PPNNPN(M), represented in Fig.1, is: uv = Bstrum = B6IPNPNPN um (6) fz \u03c4x \u03c4y \u03c4z = 1 1 1 1 0 \u2212L \u00b7 sin(60) \u2212L \u00b7 sin(60) 0 \u2212L \u2212L \u00b7 cos(60) L \u00b7 cos(60) L +k\u03c4 +k\u03c4 \u2212k\u03c4 \u2212k\u03c4 1 1 L \u00b7 sin(60) L \u00b7 sin(60) L \u00b7 cos(60) \u2212L \u00b7 cos(60) +k\u03c4 \u2212k\u03c4 u1 u2 u3 u4 u5 u6 (7) where L is the distance between the center of mass and the center of the rotor, ui is the lift generated by actuator i and k\u03c4 is a constant that relates the force produced by the rotor with respect to its reactive moment. Finally, the model of a specific multirotor is obtained by combining (4) with (7)", + "v T 8 ]T v1 := {mg + fzmin , 0, 0, 0} v5 := {mg, 0,+\u03c4ymin , 0} v2 := {mg \u2212 fzmin , 0, 0, 0} v6 := {mg, 0,\u2212\u03c4ymin , 0} v3 := {mg,+\u03c4xmin , 0, 0} v7 := {mg, 0, 0,+\u03c4zmin } v4 := {mg,\u2212\u03c4xmin , 0, 0} v8 := {mg, 0, 0,\u2212\u03c4zmin } (22) where fzmin , \u03c4xmin , \u03c4ymin and \u03c4zmin are the minimal virtual inputs and mg allows to guarantee the hover position. The size of Pm will depend on the closed-loop dynamics of the multirotor and the time that the fault diagnosis module needs to estimate the fault (position and orientation when a fault occurs). Therefore, complete information on the multirotor will be required to study the controllability of the system when ACS technique is used. The reconfigurability of the hexarotor I PPNNPN(M) (Fig.1) has been studied using ASRM and ACS. The considered assumption is that two rotors can be blocked (total fault) at the same time, thus 21 fault configurations can be considered. The characteristics of the hexarotor are detailed in Table I and Bstr is defined in (7). Firstly, ASRM has been applied. The parameters are: gmin = 0.3 and gmax is not bounded. The gains gfi of each rotor as a function of the fault are presented in Table II. The total faults on rotors 1 / 2 / 3 / 4 / 1-3 / 1-4 and 2-4 are reconfigurable because they are satisfying the imposed conditions explained in section IV-A" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001616_ut.2011.5774088-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001616_ut.2011.5774088-Figure2-1.png", + "caption": "Figure 2 Separation of the vehicle; cargo an crawler units", + "texts": [ + " As shown in Fig. 1, we propose a autonomous amphibious vehicle which covers sea-land-air;3-in-1 amphibious vehicle. In terms of the vehicle technology, the fusion of sea-land-air vehicles\u2019 functions could cover the most kind of the vehicle fusion. It has a dry cargo unit to carry UAVs and AGVs. The lateral thrusters are mounted to move as a hovering type AUV . In the bottom, a crawler unit is installed and that has two tracks with a camera dome. A funnel shape communication tower is mounted. As shown in Fig. 2, the vehicle could separate into two parts in underwater. They are connected with a tether. The bottom crawler unit has an anchor to hook in seabed and a winch to wind the tether. This function enables the cargo unit hold a certain position for a task under strong current status. The vehicle has the following functions. 978-1-4577-0164-1/11/$26.00 \u00a92011 IEEE As shown in Fig. 3, it crawls in seabed(UCV:Underwater Crawler Unit), swims in underwater(AUV) and floats on water surface (ASV). B. For Land As shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002911_s1064230718050118-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002911_s1064230718050118-Figure1-1.png", + "caption": "Fig. 1. Comparative analysis of solutions of time-optimal problem with one restriction.", + "texts": [ + "7) time-optimal problem, which can be reduced to the (4.9) and (4.14) semiinfinite optimization problem, be considered. At the first stage (paragraph 1 of the computational algorithm), only one of the restrictions in (4.11) and (4.12) should be taken into account in turn instead of (4.14) according to the general scheme of its solution by the alternance method, which was described in Section 3. A method of calculating of the form in problem (4.9) and (4.11) with significant use of information about the shape of the curve (Fig. 1a) on the interval, which is determined by the regularities of the processes of nonstationary heat conductivity under induction heating, was proposed in [2\u20134, 10\u201312]. The calculation systems of the equations of the alternance method, which are directly constructed based on Eqs. (3.4) according to the form of this curve taking into account the basic rules (2.7) and (3.2), are reduced to a system of three equations at and : (4.17) which should be solved by the standard numerical methods at the given value in three unknowns, which include the coordinate of the extremum point of the curve along with ", + "12) for the given value is satisfied on the found solution , then is obviously the desired solution of the original time-optimal problem (4.9), (4.11), and (4.12). Otherwise, it is necessary to proceed to the calculation of of the type in problem (4.9) and (4.12). It can be shown that the shape of the curve of the mean value of the difference in (4.6) and (4.12), which is uniquely determined in the conditions of the limit ratio (4.18) by integrating the dependency, does not change its configuration on the interval compared to (Fig. 1b). \u03b5 = \u03b5 \u2265 \u03b51 inf' \u03b5inf =0 2N \u0394 = \u0394 \u0394 \u03b5 > \u03b5 \u03b5 > \u03b50 0 0( ) ( ) ( ) (2) (2) 1 2 1 1min 2 2 min( , ); ;N N N \u0394[1] \u0394(2) \u0394 \u2212[1]( , ) *Q x Q [0,1] x{ = =0 2, * 2N R \u03b5 > \u03b5(2) 1 1min \u0394 \u2212 = \u2212\u03b5 \u0394 \u2212 = \u03b5 < < \u2202 \u0394 = \u2202 [1] 1 [1] 12 1 12 [1] 12 (0, ) ** ; ( , ) ** ; 0 1; ( , ) 0, Q Q Q x Q x Q x x \u03b51 12x \u0394 \u2212[1]( , ) **Q x Q \u0394 = \u0394 \u0394[1] [1] [1] 1 2( , ) \u03a6 \u0394 \u2264 \u03b5[1] 2 2( ) \u03b52 \u0394[1] \u0394[1] \u0394[2] \u0394(2) ( , *)cQ x t \u2212( , *) *Q x t Q \u2192 \u03b7 \u0394 \u2212 \u03b7 \u03b7 = \u0394 \u2212\u222b 0 0( ) ( ) 20 0 2lim ( ( , ) **) (0, ) ** x N N x Q Q d Q Q x \u0394 \u2212[2]( , ) **Q x Q [0,1] x{ \u0394 \u2212[1]( , ) **Q x Q JOURNAL OF COMPUTER AND SYSTEMS SCIENCES INTERNATIONAL Vol", + "13) by the proposed method by the criteria of the speed and power consumption of the process of induction heating cylindrical ingots of titanium alloys with a diameter of 0.54 m at an industrial frequency of 50 Hz, when the maximum surface heating density is 106 kVt/m2, before pressing are presented in Figs. 1\u20133.1 The resulting temperature distribution (1 indicates , 2 indicates ) at the end of the time-optimal control process that takes into account only one of the (4.11) and (4.12) restrictions at the specified values \u03b51 = 60\u00b0 and \u03b52 = 25\u00b0 are shown in Fig. 1a ( and Fig. 1b ( ). As follows from Fig. 1, here and , which indicates the need for the transition to the solution of the time-optimal problem taking into account both restrictions according to the proposed computational algorithm. 1 The calculations were performed by Ya.V. Kaznacheeva and Yu.A. Sadikov. \u0394 \u2212[1]( , ) **Q x Q \u0394[1]( , )cQ x [0,1] x{ [0,1] x{ \u0394 \u2212 = \u2212\u03b5 \u0394 \u2212 = \u2212\u03b5[1] [1] 1 1(0, ) ** ; (1, ) **Q Q Q Q \u0394 \u2212 = \u2212\u03b5 \u0394 = \u2212\u03b5[2] [2] 2 2(0, ) ** ; (1, ) .cQ Q Q \u03b5 > \u03b5(2) 1 min \u03b5 > \u03b5(2) 2 min \u0394[1] \u0394[2] \u0394 0( )N \u03a6 \u0394 \u2264 \u03b5 \u03a6 \u0394 \u2264 \u03b5[1] [2] 2 2 1 1( ) or ( ) \u03b51 \u03b52 \u0394 = \u03940( ) [2]N \u0394 \u2212[ ]( , ) **kQ x Q \u0394[ ]( , )k cQ x = \u0394 = \u0394 =[1] [1] 1 21, 5882 ; 605k c c = \u0394 = \u0394 =[2] [2] 1 22, 5921 ; 883k c c \u0394 > \u03b5[1] 2max | ( , )|cx Q x \u0394 \u2212 > \u03b5[2] 1max | ( , ) **| x Q x Q JOURNAL OF COMPUTER AND SYSTEMS SCIENCES INTERNATIONAL Vol" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002123_978-3-319-65298-6_65-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002123_978-3-319-65298-6_65-Figure2-1.png", + "caption": "Fig. 2. The definition of SEW plane and arm angle w", + "texts": [ + " The damped-least-squares (DLS) method of the configuration control scheme is expressed as follows: _q \u00bc W 1 v JT \u00bdJW 1 v JT \u00fe k2W 1 1\u00f0 _Xd \u00feKEe\u00de \u00f02\u00de where W\u2014\u2014symmetric positive-definite weighting matrix, W = diag[We, Wc], We and Wc are symmetric positive-definite weighting matrices for the basis of the basic task and additional task, respectively; Wv\u2014\u2014symmetric positive-definite weighting matrix; k\u2014\u2014positive scalar constant; Xd\u2014\u2014the desired behavior of the robot; Ee\u2014\u2014error, Ee = Xd\u2013X; K\u2014\u2014symmetric positive-definite feedback gain (constant); J\u2014\u2014Jacobian matrix, J = (Jee, Jw)T. where Jee denotes the end-effector Jacobian matrix, and w denotes the arm angle [12]. Through adjusting and test we can obtain the corresponding values ofWe,Wc,Wv, k, Xd and K. Via integrating Eq. (2) we can obtain the position solution of inverse kinematics. The arm angle is an important parameter which is often used to solve the inverse kinematics of redundant robot with offset configuration. It is defined as the angle between the reference plane and the current plane, as shown in Fig. 2. In Fig. 2, we define the intersection between the axes of joint 2 and joint 3 as the point S, the coordinate origin of the joint 4 as the point E, and the intersection between the axes of joint 6 and joint 7 as the point W. The plane formed by points S, E and W is called the current plane. And the position vectors from the reference coordinate origin S to points E and W are e and w, respectively. The unit vector V (V = [0, 0, 1]T) is parallel with the axis of the joint 1. The plane formed by vector V and point W is defined as the reference plane" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000988_s00542-010-1189-3-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000988_s00542-010-1189-3-Figure2-1.png", + "caption": "Figure 2 shows the coordinate system of the coupled journal and thrust bearings. This paper follows Jang\u2019s method to determine the dynamic coefficients of the coupled journal and thrust bearings by including the rotating degrees of freedom in the perturbed method (Jang and Lee 2006, Jang et al. 2007b). The perturbation equations were", + "texts": [], + "surrounding_texts": [ + "2.1 Stability analysis derived by substituting a first-order expansion of film thickness and pressure, with respect to small displacements and velocities, into the Reynolds equation. The dynamic coefficients of the journal and thrust bearings can be calculated by integrating the pressure change across the fluid film, which can be determined by using the global matrix equation of the finite element equation of the Reynolds equation and each perturbation equation in quasi-equilibrium. The total stiffness and total damping coefficients of the coupled journal and thrust bearings are the summation of all components of the stiffness and damping coefficients, respectively, of the journal and thrust bearings. The equation of motion of the rigid rotor of the HDD spindle system supported by the coupled journal and thrust bearings, including both the translation and tilting motions, can be derived as follows: M\u20acx\u00fe \u00f0C\u00feG\u00de _x\u00feKx \u00bc 0 M \u00bc ma 0 0 0 0 0 ma 0 0 0 0 0 ma 0 0 0 0 0 Ix 0 0 0 0 0 Iy 2 6666664 3 7777775 C \u00bc cxx cxy cxz cxhx cxhy cyx cyy cyz cyhx cyhy czx czy czz czhx czhy chxx chxy chxz chxhx chxhy chyx chyy chyz chyhx chyhy 2 6666664 3 7777775 G \u00bc 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 _hzIz 0 0 0 _hzIz 0 2 6666664 3 7777775 K \u00bc kxx kxy kxz kxhx kxhy kyx kyy kyz kyhx kyhy kzx kzy kzz kzhx kzhy khxx khxy khxz khxhx khxhy khyx khyy khyz khyhx khyhy 2 6666664 3 7777775 x \u00bc Dx Dy Dz Dhx Dhy 8>>>< >>>>: 9>>>= >>>>; \u00f01\u00de where ma, Ix, Iy and Iz are the mass and the mass moments of inertia of the rotor, respectively. The radius of gyration is introduced to express the mass moment of inertia with respect to the mass, Il \u00bc K2 l m \u00f02\u00de where Kl is the radius of gyration and the subscripts of l denote the x, y and z axes. Therefore, Eq. (2) can be represented as a single variable, i.e., the mass of the rotor. This makes it possible to define the stability problem of the HDD spindle system with respect to the five degrees of freedom in order to determine the critical mass of the rotor that is the threshold between stability and instability. The homogeneous solution of Eq. (1) can be assumed to be an exponential function, as follows: x \u00bc xh exp Xt\u00f0 \u00de; xh \u00bc xh yh zh hxh hyh 8>>>< >>: 9>>>= >>; \u00f03\u00de By substituting Eqs. (2) and (3) into Eq. (1), the following equation can be obtained. X2M\u00fe X\u00f0C\u00feG\u00de \u00feK xh exp\u00f0Xt\u00de \u00bc 0 \u00f04\u00de The eigenvalue of Eq. (4) can be generally expressed as X = - Xreal ? iXimg. The motion of the rotor is stable if Xreal is larger than zero, and it is unstable if Xreal is less than zero. Therefore, the critical condition can be determined when Xreal is equal to zero, such that the following two equations are obtained from the characteristic determinant of Eq. (4). A1\u00f0ma\u00de5c\u00f0X2 img\u00de 5 c \u00feA2\u00f0ma\u00de5c\u00f0X2 img\u00de 4 c \u00feA3\u00f0ma\u00de4c\u00f0X2 img\u00de 4 c \u00feA4\u00f0ma\u00de4c\u00f0X 2 img\u00de 3 c \u00feA5\u00f0ma\u00de3c\u00f0X 2 img\u00de 4 c \u00feA6\u00f0ma\u00de3c\u00f0X 2 img\u00de 3 c \u00feA7\u00f0ma\u00de3c\u00f0X2 img\u00de 2 c \u00feA8\u00f0ma\u00de2c\u00f0X2 img\u00de 3 c \u00feA9\u00f0ma\u00de2c\u00f0X2 img\u00de 2 c \u00feA10\u00f0ma\u00de2c\u00f0X2 img\u00dec\u00feA11\u00f0ma\u00dec\u00f0X2 img\u00de 3 c \u00feA12\u00f0ma\u00dec\u00f0X2 img\u00de 2 c \u00feA13\u00f0ma\u00dec\u00f0X2 img\u00dec\u00feA14\u00f0X2 img\u00de 2 c \u00feA15\u00f0X2 img\u00dec\u00feA16 \u00bc 0 \u00f05\u00de B1\u00f0ma\u00de4c\u00f0X2 img\u00de 4 c \u00fe B2\u00f0ma\u00de4c\u00f0X2 img\u00de 3 c \u00fe B3\u00f0ma\u00de3c\u00f0X2 img\u00de 3 c \u00fe B4\u00f0ma\u00de3c\u00f0X 2 img\u00de 2 c \u00fe B5\u00f0ma\u00de2c\u00f0X 2 img\u00de 3 c \u00fe B6\u00f0ma\u00de2c\u00f0X 2 img\u00de 2 c \u00fe B7\u00f0ma\u00de2c\u00f0X2 img\u00dec \u00fe B8\u00f0ma\u00dec\u00f0X2 img\u00de 2 c \u00fe B9\u00f0ma\u00dec\u00f0X2 img\u00dec \u00fe B10\u00f0ma\u00dec \u00fe B11\u00f0X2 img\u00de 2 c \u00fe B12\u00f0X2 img\u00dec \u00fe B13 \u00bc 0 \u00f06\u00de where Ai(i = 1, 2,\u2026,16) and Bj(j = 1, 2,\u2026,13) are functions of the dynamic coefficients of the FDBs, the radius of gyration of the rotor, and the rotating speed of the HDD spindle system. The solutions to Eqs. (5) and (6) are the critical mass (ma)c, and the corresponding frequency, Ximg. The behavior of the disk-spindle system supported by the coupled journal and thrust bearings is stable if the mass, ma, is smaller than the critical mass, (ma)c, and it is unstable if ma is greater than (ma)c. 2.2 Nonlinear equations of motion The proposed stability analysis assumes that the rotating disk-spindle system supported by the coupled journal and thrust bearings is linear and that the dynamic coefficients of the coupled journal and thrust bearings are time-invariant, i.e., constant. However, as the disk-spindle system rotates, the dynamic coefficients change depending on the relative positions of the rotating shaft and the sleeve. This paper derives the nonlinear equations of a rotating disk-spindle system supported by coupled journal and thrust bearings to show that the proposed stability analysis method is also effective for a nonlinear rotating disk-spindle system supported by coupled journal and thrust bearings. Figure 3 shows the free body diagram of the HDD spindle system, including the bearing reaction force, rotor weight, and centrifugal force due to the mass unbalance. The motion of the rigid rotor of the HDD spindle system can be described in terms of the five degrees of freedom, i.e., the three translational displacements in the x, y and z directions and two tilting displacements in the hx and hy directions, so that the five nonlinear differential equations can be derived, as follows: ma\u20aceX \u00bc FxJ1 \u00fe FxJ2 mueu _h2 z cos / c2 \u00fe FzT1 \u00fe FzT2\u00f0 \u00des2 \u00feWX \u00f07\u00de ma\u20aceY \u00bc FxJ1 \u00fe FxJ2 mueu _h2 z cos / s1s2 \u00fe FyJ1 \u00fe FyJ2 mueu _h2 z sin / c1 FzT1 \u00fe FzT2\u00f0 \u00des1c2 \u00feWY \u00f08\u00de ma\u20aceZ \u00bc FxJ1 \u00fe FxJ2\u00f0 \u00dec1s2 \u00fe FyJ1 \u00fe FyJ2 s1 \u00fe FzT1 \u00fe FzT2\u00f0 \u00dec1c2 \u00fe mag\u00fe mug\u00f0 \u00de mueu _h2 z sin /s1 cos /c1s2\u00f0 \u00de \u00feWZ \u00f09\u00de Ixc2 \u20achx \u00bc Ix \u00fe Iy Iz s2 _hx _hy Iz _hy _hz \u00feMxJ1 \u00feMxJ2 \u00feMxT1 \u00feMxT2 zu mugs1 mueu _h2 z sin / mugc1c2eu sin / WY Lf \u00f010\u00de Iy \u20achy \u00bc Ix Iz\u00f0 \u00dec2s2 _hx _hx \u00fe Izc2 _hz _hx \u00feMyJ1 \u00feMyJ2 \u00feMyT1 \u00feMyT2 \u00fe zu mugc1s2 mueu _h2 z cos / \u00fe mugc1c2eu cos /\u00feWXLf \u00f011\u00de where c1, s1, c2 and s2 are coshx, sinhx, coshy and sinhy, respectively, and mu, eu, zu and / are the mass unbalance of the rotor of the HDD spindle system and the positions of the mass unbalance from the mass center G, respectively. Fx, Fy, Fz, Mx and My are the reaction forces and moments, respectively, generated by the FDBs, which are calculated by integrating the pressure and shear stress along the fluid film. The subscripts of J and T denote the journal and thrust bearings, respectively. WX, WY, WZ and Lf are the external force of each axis and the z distance of the external force from the mass center G, respectively. The nonlinear equations of motion of the HDD spindle system are solved by using the fourth-order Runge\u2013Kutta method to investigate the motion of the HDD spindle system, including the whirling, tilting, and axial motions." + ] + }, + { + "image_filename": "designv11_33_0002063_s00170-017-0656-8-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002063_s00170-017-0656-8-Figure2-1.png", + "caption": "Fig. 2 The kinematic relationship", + "texts": [ + "1 and the tooth surface forming principle will be deduced in Sect. 2.2. To keep the cutting tool and the workpiece in the correct position during machining hypoid gears using forming method, the workpiece mounting angle and the cutting tool position must be adjusted in the first place. There are mainly two movements in the processing: one is the cutting movement formed by the cutting tool rotation, and the other is the feed movement formed by the cutting tool moving in the axial direction, and the movement style is shown in Fig. 2. Under the domination of the two movements, all the trajectories of the inner cutting blades, outer cutting blades, and middle cutting blades are spiral lines, and these spiral lines will be extended until the cutting tool reaches the root; a complete tooth slot will be formed after the cutting tool goes back. Right now, the workpiece will be rotated a particular angle around the axis to process the next tooth, the degree of which is decided by the tooth number, and this process will be repeated until all the teeth are processed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003346_012081-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003346_012081-Figure1-1.png", + "caption": "Figure 1. The forces on the hollow cylinder move in the incline.", + "texts": [ + " UNNES Physics International Symposium 2018 (UPIS2018) IOP Conf. Series: Journal of Physics: Conf. Series 1170 (2019) 012081 IOP Publishing doi:10.1088/1742-6596/1170/1/012081 Suppose a hollow cylinder is released downhill from a stationary state on an inclined plane. It hollow rolls without slippage under the influence of two forces. They are a gravitational force and frictional force f between the spherical edge and the inclined planes [13]. The description of the forces on the hollow cylinder that moves on the inclined plane is shown in Figure 1. The description of hollow cylinder motion can be done mathematically by using two basic equations. The first equation is the translational motion equation, assuming that all external forces are working in the center of hollow cylinder mass. The Newton's law II can be written as [5,6,8]: (1) The second equation is the equation of rotational motion against the center of the mass of the hollow cylinder torque whose direction is perpendicular to the axis is rotation [5,7,12]: (2) with is torque, force F and R are arm styles respectively. . In Fig. 1 it appears that the force which is directed perpendicular to the rotation axis is the frictional force f, and the hollow cylinder force arm rotating with the rotary axis of one of its diameters is the radius of the hollow cylinder r, so Equation (2) becomes [5,6,11]: (3) or the relationship between torque \u03c4, the moment of inertia I and the angular accent \u03b1 of the rotating object are [5,6,7,11,14, 18]: (4) and [8,9,11] (5) where in the general equation of the moment of inertia I am theoretically (IT) a hollow cylinder of mass m and the radius r of the center axis is clarified [15,16]: (6) We can use 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003710_tro.2019.2938348-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003710_tro.2019.2938348-Figure3-1.png", + "caption": "Fig. 3. (a) KG for SM4 (R(z)) by adding a rotational DoF of R(z) to the middle of the original subchain Mj 4 in a M4 KG. (b) KG for SM3A (R(y)).", + "texts": [ + " Example 6 (KG for SM4 (R(z))): It is easy to see that the maximal inscribing symmetric subspace Mmax of SM4 (R(z)) is M4. For M4, we have [m4,m4] = {e3, e6}. We choose Q2 = R(z) = {ez\u0302\u03b8 | \u03b8 \u2208 R}, which satisfies the condition in Proposition 5. Now we add this additional rotational DoF R(z) to the middle of the original subchain M j 4 of M4. The new subchain is denoted as N j 4 . Assembling 3 N j 4 together and interconnecting them with a prismatic joint (instead of the cylindrical pair in the original M4 KG) yields a KG for SM4 (R(z)), as shown in Fig. 3(a). The KG for SM4 (Hp(z)) can be synthesized in the same way. Some reflective-type adjoint-invariant submanifolds are contained in one or multiple reflective-type submanifolds (called covering reflectivetype submanifolds). The KG for these covering reflective-type submanifolds can be used as the primitive subchains. Example 7 (KG for SM3A (R(y))): Notice that SM3A (R(y)) \u2282 SM\u03024 (R(y)), where M\u03024 = Iex\u0302\u03c0/2(M4) is a 4-D symmetric subspace satisfyingM3A \u2282 M\u03024. SM\u03024 (R(y)) is equivalent to Iex\u0302\u03c0/2(SM4 (R(z)))", + " On the other hand SM3A (R(y)) = {ee\u03021\u03b81ee\u03023\u03b82+e\u03024\u03b83ey\u0302\u03b84ee\u03021\u03b81ee\u03023\u03b82+e\u03024\u03b83}based on the facts thatM3A = T (x) \u00b7M2A. Since ee\u03021\u03b81ee\u03023\u03b82+e\u03024\u03b83 = ee\u03023\u03b82+e\u03024\u03b83ee\u03021\u03b81 by direct computation, we have SM3A (R(y)) = ST (x)(SM2A (R(y))) \u2282 ST (x)(M3B) \u2282 T \u2212(x) \u00b7M3B \u00b7 T +(x) where T \u2212(x) \u00b7M3B \u00b7 T +(x) can be generated by cascading a pair of symmetric translational pair (T \u2212(x), T +(x)) with a KG (e.g.,[10, 4]) for M3B in between. Finally, assembling the KG for SM\u03024 (R(y)) and that for T \u2212(x) \u00b7M3B \u00b7 T +(x) yields a KG for SM3A (R(y)), as illustrated in Fig. 3(b). This can be proved by recalling that at home configuration e, the constraint force of the former subchain is {e2}, while that of the latter subchain is {e6} and, therefore,TeSM\u03024 (R(y)) \u2229 Te(T \u2212(x) \u00b7M3B \u00b7 T +(x)) = TeSM3A (R(y)). In this article, we proposed a class of submanifolds of SE(3), the adjoint-invariant submanifolds, which extended the theory of Lie subgroups and symmetric subspaces by relaxing the symmetry requirements in these objects. We studied global geometric properties as well as existence and uniqueness of adjoint-invariant submanifolds based on the theories of distributions on manifolds and their integrability" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000661_raad.2010.5524601-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000661_raad.2010.5524601-Figure4-1.png", + "caption": "Figure 4. Cutting parameters", + "texts": [ + " The adjustment of strategy parameters is constrained by compliance to the cutting quality (no bone splinters or cartilage) (Fig. 3). Many meat-cutting models exist. The present work was mainly interested in the reduction of cutting forces for operations carried out manually. This work focuses on the cutting angle [7] [8] or the shape of the blade [9]. It shows, for example, that when the cutting angle is less than 30\u00b0 efforts are reduced by 30%. As part of the proposed cutting strategy, we introduced a set of additional parameters (Fig. 4) concerning the lateral contact force to be maintained and the orientation of the blade. The strategy is defined by the following parameters: Vf, the feed rate, Fl, the lateral support force of the blade on the bone, , the angle between the cutting edge of the blade and the feed direction, , clamp angle between the plane tangent to the surface of the bone and the ground plane of the blade, which ensures a path as near as possible to the bone. A path is defined by a set of poses (position and orientation) with n the normal to the local tangent plane, u the tangent of the blade in the plane tangent to the surface of the bone and h the normal to the feed rate contained in the tangent plane" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003466_j.mechmachtheory.2019.06.017-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003466_j.mechmachtheory.2019.06.017-Figure4-1.png", + "caption": "Fig. 4. Generation of tooth surfaces p and s of the tapered involute pinion and shaper.", + "texts": [ + " \u03b2 is the helix angle; \u03b2 > 0 and \u03b2 = 0 correspond to a right-hand pinion and a spur pinion, respectively. \u03b1n is the pressure angle, which is measured on the normal plane of the rack-cutter. Note that L d in Eq. (3) is calculated by the following equation: L d = \u03bb\u03c0m n 2 ( 1 + \u03bb) cos \u03b1n (4) where \u03bb= s 0 / w 0 , m n is the module of the rack-cutter. Parameter \u03bb represents the ratio between the space width and tooth thickness measured on the pitch line. Typically, \u03bb= 1. The coordinate systems used for generation of the tapered involute pinion and the tapered involute shaper are shown in Fig. 4 . Both the pinion and the shaper are generated on a conical blank by a rack-cutter. The pinion surfaces p and shaper surfaces s are conjugated as envelope sets of the rack-cutter surfaces ri . The coordinate system S e is fixed in the spatial framework. The coordinate system S i is attached to the tapered involute pinion or the shaper. The coordinate system S r is attached to the rack. S d is an auxiliary coordinate system. The tooth surfaces p and s of the tapered involute pinion and shaper can be expressed as follows: { r i ( u i , l i , \u03c8 i ) = M ie ( \u03c8 i ) M ed M dr r ri ( u i , l i ) f ri ( u i , l i , \u03c8 i ) = n ri ( u i ) \u00b7 v ( ri ) r ( u i , l i , \u03c8 i ) = 0 (5) M ie ( \u03c8 i ) = \u23a1 \u23a2 \u23a3 cos \u03c8 i sin \u03c8 i 0 0 \u2212 sin \u03c8 i cos \u03c8 i 0 0 0 0 1 0 0 0 0 1 \u23a4 \u23a5 \u23a6 (6) M ed = \u23a1 \u23a2 \u23a3 1 0 0 \u03c8 i r pi 0 1 0 \u2212r pi 0 0 1 0 0 0 0 1 \u23a4 \u23a5 \u23a6 (7) M dr = \u23a1 \u23a2 \u23a3 1 0 0 0 0 cos \u03b4 sin \u03b4 0 0 \u2212 sin \u03b4 cos \u03b4 0 0 0 0 1 \u23a4 \u23a5 \u23a6 (8) where r i ( i = p, s ) is the position vector of tooth surfaces p of the pinion or tooth surfaces s of the shaper" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000856_2011-01-2117-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000856_2011-01-2117-Figure6-1.png", + "caption": "Fig. 6 Schematic illustration of thin-film electrode configuration", + "texts": [], + "surrounding_texts": [ + "Oil film thickness in engine main bearings was determined from the capacitance between the crankshaft and a thin-film electrode formed on the surface of the bearing. Since the thin-film electrode completely follows the bearing deformation, this technique is effective for measuring oil film thickness in engine bearings whose elastic deformation is not negligible. The new oil film thickness measurement technique is hereinafter called the \u201cthin-film method.\u201d" + ] + }, + { + "image_filename": "designv11_33_0000758_s11465-012-0317-4-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000758_s11465-012-0317-4-Figure7-1.png", + "caption": "Fig. 7 SNU 3-UPU parallel mechanism. (a) SNU mechanism; (b) limb constraint system", + "texts": [ + " As we known that the SNU 3-UPU parallel mechanism prototype will always exhibit the parasitic rotations in practice [17]. Now we will use the value of norm\u00f0\u03c9\u00de to find the bound of mechanism instability, which indicates there, will appear the parasitic rotation when the number exceeds the bound. For SNU 3-UPU parallel mechanism, the first revolute axes of universal joint attached to the fixed base intersecting at one common point. And the same is true for the last revolute axes of universal joint attached to the moving platform, as shown in Fig. 7(a). Take one limb as an example, the direction of the constraint couple applied on the moving platform is sketched in Fig. 7(b). It\u2019s easy to find that the angle \u03b2i from the direction of constraint couple to the moving platform keeps in a relative stable area. It is around \u03c0=2, and we express the angle \u03b2i as \u03b2i \u00bc \u00bd\u03c0=2 . Of course, the mechanism will be singularity when all the three constraint couples are parallel with each other. However, since the existing clearances, these three constraint couples will never truly be parallel with each other in practice. Since the value \u03b2i \u00bc \u00bd\u03c0=2 , we choose a values set of \u03b23, (80\u03c0=180, 85\u03c0=180, 87\u03c0=180, 89\u03c0=180), to get the contour atlas of norm\u00f0\u03c9\u00de, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003813_s12206-019-1039-x-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003813_s12206-019-1039-x-Figure8-1.png", + "caption": "Fig. 8. Resultant force acting on the Segway (in case of acceleration).", + "texts": [ + " 6 in order to improve the driving stability of driver. If the Segway rapidly accelerates or decelerates in straight line motion, an inertial force acts in the opposite direction to the acceleration of the Segway. If the inertial force acting on the driver and the force acting on the Segway are unbalanced, the risk of rollover is increased. Therefore, if the footplate of Segway is lifted to oppose the force acting on the Segway, the stability of the driver can be increased during rapid acceleration or deceleration as shown in Fig. 8. In Fig. 8, the inertial force Fi acting on the driver is the same magnitude as the force F acting on the Segway, and acts in the opposite direction. It can be calculated using the weight and acceleration of the Segway as follows: ( )iF F m a= - = - r . (10) In other words, the value of how much footplate of the Segway has to be lifted can be calculated using the inertial force Fi. The angle, Pq , that the sum of Fi and F become zero can also be calculated. When the Segway is moving at a fixed speed in a straight line, if it suddenly and rapidly decelerates by a factor of a r , the force acting on the Segway is as shown in the Fig. 8. As illustrated in the Fig. 8, a normal force n r can be decomposed into vertical and horizontal vectors. The vertical vector of the normal force is equal and acts the opposite direction to the force of gravity. So, the vertical vector of normal force can be expressed as follows: cos 0Pn mgq - = r r . (11) Also shown in Fig. 8, a horizontal vector of the normal force n r acts opposite direction and is equal to the inertial force acting on the driver, which can be expressed as follows: sini PF n q- = r . (12) When the Segway rapidly decelerates or accelerates by the a r , the control angle Pq which how much the footplate of Segway has to be lifted can be calculated: ( )arctan( )P a g q - = r r , (13) which can be derived by putting Eqs. (10) and (11) into Eq. (12). When the Segway runs in a curved trajectory, centripetal force acts on the Segway and centrifugal force acts on the driver" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000777_s0025654411040017-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000777_s0025654411040017-Figure4-1.png", + "caption": "Fig. 4. Fig. 5.", + "texts": [], + "surrounding_texts": [ + "By setting \u03b4 = 0 in Eqs. (3.6) and (3.7) and by introducing new variables z1 and z2 by formulas z1 = z + a1\u03d5 and z2 = z \u2212 a2\u03d5, we obtain the equations z\u2022\u20221 + a11(c1z1 + \u03b31z \u2022 1) + a12(c2z2 + \u03b32z \u2022 2) = 0, z\u2022\u20222 + a21(c1z1 + \u03b31z \u2022 1) + a22(c2z2 + \u03b32z \u2022 2) = 0, (4.1) MECHANICS OF SOLIDS Vol. 46 No. 4 2011 a11 = 1 + a2 1 \u2212 a1f1, a12 = 1 \u2212 a1a2 \u2212 a1f2, a21 = 1 \u2212 a1a2 + a2f1, a22 = 1 + a2 2 + a2f2. (4.2) The dimensionless variables ak and fk (k = 1, 2) in formulas (4.2) are the ratios of the corresponding variables introduced in the preceding sections to the radius of inertia \u03c1; i.e., we assume that the radius of inertia \u03c1 is equal to unity. The variables z1 and z2 have the following physical meaning. These are small deviations of the points O1 and O2 from their equilibrium positions described in Sec. 3. Equations (4.1) and (4.2) and relations (1.1) with the use of the Lyapunov theorem permit solving the above stability and instability problem for any fixed v. Let us write out the characteristic polynomial of the linear system (4.1) as p4 + p3(a11\u03b31 + a22\u03b32) + p2(a11c1 + a22c2 + \u03b31\u03b32\u0394) + p(\u03b31c2 + \u03b32c1)\u0394 + c1c2\u0394 = 0, (4.3) where \u0394 = a11a22 \u2212 a12a21. Applying the Routh\u2013Hurwitz theorem to the polynomial (4.3), we arrive at the following result. Proposition 1. For the asymptotic stability of the equilibrium in the case of rectilinear and uniform motion of the vehicle, it is necessary and sufficient that the rolling friction coefficients f1 and f2 (and hence the velocity v of the vehicle translational motion determined by (1.1)) satisfy the inequalities a11\u03b31 + a22\u03b32 > 0, a11c1 + a22c2 + \u03b31\u03b32\u0394 > 0, \u0394 > 0, (4.4) (a11\u03b31 + a22\u03b32)[a11\u03b32c 2 1 + a22\u03b31c 2 2 + \u03b31\u03b32\u0394(\u03b31c2 + \u03b32c1)] \u2212 (\u03b31c2 + \u03b32c1)2\u0394 > 0, \u0394 = a11a22 \u2212 a12a21 = a(a + f2 \u2212 f1), a = a1 + a2. (4.5) The dimensionless parameters aij (i, j = 1, 2) are determined by formulas (4.2). Consider several particular cases where Proposition 1 applies. Proposition 2. If the system is completely symmetric, i.e., if \u03b3k = \u03b3 = 0, ck = c, fk = f, k = 1, 2, a1 = a2 = a0, (4.6) then inequalities (4.4) hold for each f ; i.e., the asymptotic stability of the vehicle equilibrium considered above is preserved for any wheel rolling friction coefficients and hence for any velocities of its translational motion. The proof is by a straightforward verification of inequalities (4.4) under conditions (4.6). In this case, it follows from (4.5) that \u0394 = a2 > 0 and the first two inequalities in (4.4) are satisfied, because they are reduced to the following ones: 2\u03b3(1 + a2 0) > 0, 2c(1 + a2 0) + \u03b32a2 > 0. The last inequality in (4.4) has the form c2(1 \u2212 a2 0) 2 + 4\u03b32c2a2 0(1 + a2 0) > 0. Thus, the stability conditions (4.4) are satisfied for any f \u2265 0, and the proof of Proposition 2 is complete. Proposition 3. Assume that all conditions (4.6) except for the last are satisfied; i.e., a1 =a2 (the case of incomplete symmetry; i.e., the car body center of mass is displaced). Then, for a1 < a2, the stability conditions (4.4) are satisfied for every f \u2265 0, and for a1 > a2, the stability is violated for f > f0 = 1 a1 \u2212 a2 [ 2 + a2 1 + a2 2 + \u03b5(a1 + a2)2 \u2212 \u221a \u03b52(a1 + a2)4 + 4(a1 + a2)2 ] > 0, \u03b5 = \u03b32 c . (4.7) Remark. The quantity f0 in inequality (4.7) can be arbitrarily small for some values of the system parameters. For example, as \u03b5 \u2192 0 (small damping or large rigidity of the wheel) and for a1a2 = 1 and a1 > 0, we have f0 = (a1 \u2212 1)(1 + a2 1)/(a1 + a2 1), which can be arbitrarily small as a1 \u2192 1. MECHANICS OF SOLIDS Vol. 46 No. 4 2011 The proof is by a straightforward verification. Under the conditions given in Proposition 3, the instability inequalities (4.4) become \u03b3[2 + a2 1 + a2 2 + (a2 \u2212 a1)f ] > 0, c[2 + a2 1 + a2 2 + (a2 \u2212 a1)f ] + \u03b32(a1 + a2)2 > 0, \u0394 = (a1 + a2)2 > 0, x2 + 2x\u03b5(a1 + a2)2 \u2212 4(a1 + a2)2 > 0, x = 2 + a2 1 + a2 2 + (a2 \u2212 a1)f. (4.8) For a2 > a1, these inequalities are necessarily satisfied for any f > 0. (For the first three inequalities in (4.8), this is obvious, and for the last, this follows from the inequality x > 2 + a2 1 + a2 2 > 2(a1 + a2).) But if a2 < a1, then the first inequality in (4.8) is violated for f > f1 = (2 + a2 1 + a2 2)/(a1 \u2212 a2), the second is violated for f > f2 > f1, the third is always satisfied, and the last is violated for f \u2208 [f0, f00], where f0 is given by the formula in (4.7) and f00 is given by the same formula with the plus sign at the radical. It is easily seen that 0 < f0 < f1 < f00, which implies that all inequalities in (4.8) are violated for f > f0. The proof of Proposition 3 is complete. Proposition 4. Let f2 = 0, \u03b30 = 0, and f1 = f > 0; i.e., assume that the fore (driving) wheel does not experience any rolling resistance and there is no dissipation in it. In this case, if a1a2 < 1, then the instability arises for f > a1, and if a1a2 > 1, then the instability arises for f > (a1a2 \u2212 1)/a2. Proposition 5. Let f1 = 0, \u03b31 = 0, and f2 = f > 0; i.e., assume that the rear (driven) wheel does not experience any rolling resistance and there is no dissipation in it. In this case, if a1a2 < 1, then the instability arises for f > (1 \u2212 a1a2)/a1, and if a1a2 > 1, then the system is stable for any f > 0. The proof of Propositions 4 and 5 is by a straightforward verification of inequalities (4.4) in Proposition 1 under the assumptions stated above. Let \u03b31 > 0, \u03b32 = 0, and \u03b32 > 0. Then the following assertion holds. Proposition 6. Under the above conditions, the asymptotic stability domain D1 in the plane of the parameters (f1, f2) is given by the following inequalities. If 0 < a1a2 < 1, then 0 < f2 < 1 \u2212 a1a2 a1 , 0 < f1, c1a1f1 \u2212 c2a2f1 < c0, f1 \u2212 f2 < a1 + a2. (4.9) If 1 < a1a2 < \u221e, then 0 < f1 < a1a2 \u2212 1 a2 , 0 < f2, c1a1f1 \u2212 c2a2f2 < c0, (4.10) c0 = c1(1 + a2 1) + c2(1 + a2 2). (4.11) The proof of Proposition 6 is by a straightforward verification of inequalities (4.4) under the restrictions given in the statement. Note that if one of the inequalities in conditions (4.10), (4.11) is not satisfied, then the vibrations under study are unstable. Now let \u03b31 = \u03b32 = 0; i.e., assume that there is no dissipation in the wheels. In this case, Eq. (4.3) is biquadratic, the stability of system (4.1) can be only nonasymptotic, and the conditions for such stability are equivalent to the conditions that this biquadratic equation has only pure imaginary roots. These considerations lead to the following result. Proposition 7. For \u03b31 = \u03b32 = 0, the stability domain D0 in the plane of parameters (f1, f2) is given by the following inequalities: f1 > 0, f2 > 0, c0 > c1a1f1 \u2212 c2a2f2, a1 + a2 > f1 \u2212 f2, [c0 \u2212 (c1a1f1 \u2212 c2a2f2)]2 > 4c1c2(a1 + a2)(a1 + a2 \u2212 f1 + f2). (4.12) The proof of Proposition 7 follows from well-known necessary and sufficient conditions for the existence of pure imaginary roots of a biquadratic equation. Proposition 8. The inclusion D1 \u2282 D0 holds. MECHANICS OF SOLIDS Vol. 46 No. 4 2011 The proof of Proposition 8 is by a straightforward substitution of the extreme values in the linear inequalities (4.9) or (4.10) into inequalities (4.12) and verification that the latter are satisfied. Remark. Proposition 8 means that this nonsymmetric addition of dissipation narrows down the stability domain in the plane of the parameters (f1, f2). Graphically (schematically), the domains D1 and D0 in various possible cases are shown in Figs. 2\u201315, where digit 1 corresponds to the line f1 \u2212 f2 = a1 + a2, digit 2 corresponds to the line c1a1f1 \u2212 c2a2f2 = c0 = c1(1 + a2 1) + c2(1 + a2 2), digit 3 corresponds to the line f2 = |1 \u2212 a1a2|/a1, digit 4 corresponds to the line f1 = |1 \u2212 a1a2|/a2, digit 5 corresponds to the line f1 \u2212 f2 = \u2212|1 \u2212 a1a2|(a1 + a2)/(a1a2), and digit 6 corresponds to the parabola (c0 \u2212 c1a1f1 + c2a2f2)2 \u2212 4c1c2(a1 + a2)(a1 + a2 \u2212 f1 + f2) = 0. Figures 2, 4, 6, 8, 10, 12, and 14 correspond to the case of \u03b31 = \u03b32 = 0, while Figs. 3, 5, 7, 9, 11, 13, and 15 correspond to the case of \u03b31 = 0, \u03b32 = 0 or \u03b31 = 0, \u03b32 = 0 for all possible distinct values of the parameters a1, a2, c1, and c2. Now consider the case of weighted wheels, m1 = 0 and m2 = 0. Since we assume that the car body mass is m = 1, the quantities m1 and m2 are the respective dimensionless ratios of the wheel masses to the car body mass. Simple calculations show that the form of the vibration equations (4.1) remains the same and the coefficients aij (i, j = 1, 2) are determined by the functions a11 = \u03c12 \u2212 (a1 + am2)(\u2212a1 + f1) \u03c12 0 , a12 = \u03c12 \u2212 (a1 + am2)(a2 + f2) \u03c12 0 , a21 = \u03c12 + (a2 + am1)(\u2212a1 + f1) \u03c12 0 , a22 = \u03c12 + (a2 + am1)(a2 + f2) \u03c12 0 , where \u03c12 0 = m0\u03c1 2 and m0 = 1 + m1 + m2. Then one can readily compute that \u0394 = a11a22 \u2212 a12a21 = a(a + f2 \u2212 f1)/\u03c12 0. Proposition 1 remains valid under the above variations in aij , i, j = 1, 2, and \u0394. In a similar way, one can restate Propositions 2\u20138. Equations (4.1) and (4.2) describing the vehicle vibrations under the action of the rolling friction forces admit the following analogies (interpretations) to well-known systems in mechanics. Each of the coefficient matrices multiplying zk and z\u2022k (k = 1, 2) in (4.1) with (4.2) taken into account consists of MECHANICS OF SOLIDS Vol. 46 No. 4 2011 two parts. The first part (it is said to be unperturbed) does not contain the friction coefficients f1 and f2. The second (perturbed) part depends on these coefficients linearly. As follows from the above derivation of the equations of motion, the origin of the perturbed part significantly depends on the following two external causes. One of them is the interaction between the deformable wheels and the rough road, which causes the rolling resistance torques Mk1 and Mk2; the second cause is the torque M1 from the engine (external energy source) to the driving wheels ensuring their (and the driven wheels) permanent rolling if there are resistance torques mentioned above. If these causes are absent, then the system motion occurs for f1 = f2 = 0, and Eqs. (4.1) contain only the first parts of the above-mentioned matrices for which the solutions have the property of (asymptotic) stability (a conservative system MECHANICS OF SOLIDS Vol. 46 No. 4 2011 under the action of some forces with complete or incomplete dissipation). But if f2 1 + f2 2 = 0, then these causes result in the appearance of additional terms in Eqs. (4.1), which, according to [6], can be interpreted as positional (circular) forces (for the corresponding terms with z1 and z2); gyroscopic and nonconservative (depending on the velocities) forces (for the corresponding terms with z\u20221 and z\u20222); conservative positional forces (for the corresponding terms with z1 and z2); and dissipative (Rayleigh) MECHANICS OF SOLIDS Vol. 46 No. 4 2011 MECHANICS OF SOLIDS Vol. 46 No. 4 2011 forces (for the corresponding terms with z\u20221 and z\u20222). In the general case, such a decomposition of forces was performed in [7]. Thus, the existence of the above-mentioned external causes leads to the appearance of additional forces of the described structure, whose destructive character is well known in mechanics of gyroscopic systems, structural mechanics (in particular, in mechanics of bridges), and other adjacent fields; it is described in the literature [8\u201310]. Even if there forces do not actually lead to the vehicle instability, they still significantly change the frequencies of its vibrations in the vertical plane, and this fact cannot be ignored when solving the corresponding problems. These facts were already considered earlier by the authors in the paper [11]." + ] + }, + { + "image_filename": "designv11_33_0000143_1.3514074-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000143_1.3514074-Figure1-1.png", + "caption": "Figure 1. (a) Device schematic of a contacted gold nanowire. Contact electrodes were passivated using a photoresist insulating passivation layer. A trench was opened in the photoresist directly over the nanowire electrode to allow contact between the nanowire and the external environment. (b) Optical micrograph of a fully fabricated nanowire device. Scale bar, 50 \u03bcm.", + "texts": [ + " Although individual nanoelectrodes have been used to probe nanoscale electrochemical phenomena, major disadvantages associated with these electrodes include: the difficulty and length of time for fabrication, leaking of electrolyte though the seals, and the extremely small currents (10-100 pA) that may be achieved with them. Such small currents pose problems when the measured current is to be used as the signal in a sensor electroanalytical application. Experimental: Nanowire Fabrication and Integration Herein we explore individual gold nanowire (NW) devices for use as nanoelectrodes in electrochemical studies fabricated using a low cost robust fabrication approach capable of achieving comparatively high electrochemical currents (nA); see Figure 1a. Nanowires were fabricated using the nanoskiving technique recently pioneered by the Whitesides Group (28). Briefly, a thin film of gold was evaporated onto an epoxy substrate which was further encapsulated in epoxy to form a block. Following curing, sections were sliced from the block using ultramicrotomy to yield nanomembranes each containing one gold nanowire. The nanomembranes were then deposited onto an oxidized silicon chip substrate (90 nm, thermal SiO2) bearing arrays of micron-scale binary alignment marks, which were used for subsequent overlay of top-contact electrodes using optical lithography", + " Nanowires with lengths of several hundred microns and widths and heights of ~ 207 \u00b1 1 nm and 254 \u00b1 11 nm, respectively, were routinely fabricated. Using the binary mark identifiers for alignment and registration, contact electrodes were overlaid using optical lithography, metal evaporation (Ti 10 nm, Au 200 nm) and liftoff. Finally, a layer of photoresist (1 \u03bcm) was spin coated on to the chip to serve as an insulating layer preventing unwanted electrochemical reactions occurring at the interconnection tracks. Optical lithography was then employed to open a trench over the nanowire; see Figure 1b. Finally, chips bearing electricallycontacted nanowires were assembled onto printed circuit boards, electrically contacted using wedge wire bonding (25 \u00b5m aluminum wire) and the bonds protected by epoxy to complete device packaging. Control devices (electrodes overlaid onto substrates without a nanowire present) were fabricated in a similar manner. Results: Following fabrication, nanowires were characterized using a combination of currentvoltage (I-V) measurements and cyclic voltammetry, (CV). Two point electrical characterization of packaged devices demonstrated fully functional nanowires that exhibited very low resistance ~ 80 \u2126 (68 -93 \u2126) for ~ 50 \u03bcm long wires; see Figure 2a" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002349_s12289-017-1388-x-Figure13-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002349_s12289-017-1388-x-Figure13-1.png", + "caption": "Fig. 13 Measurement of the projection of the contact arc between ring and tools", + "texts": [ + " Finally, the radial forming force has been analytically calculated by utilizing the three different average projections of the projection of the contact arc based on 18, 36 and 72 slices subdivision strategies. The results of the comparison between analytical results, for the three slices subdivision strategies, and the numerical simulation results, are reported in Table 6. In Table 6, the value of the projection of the contact arc between ring and tools in the numerical simulation have been estimated following the procedure shown in Fig. 13. For each round of the process, four different values have beenmeasured on four different sections of the ring located at 0\u00b0, 90\u00b0, 180\u00b0 and 270\u00b0 with respect to the central axis of mandrel and main roll. In all the cases, the projection of the contact arc in the numerical simulations has been estimated at the center of the ring, along the vertical direction, as shown in Fig. 13. Based on the results comparison reported in Table 6, the subdivision of the ring into 72 slices results in a considerable increase of the computational time tcomp,CAD with a light enhance in the estimation of both projection of the contact arc Lc,CAD and radial forming force Fauthors. For this reason, authors have chosen to utilize the subdivision of the ring into 36 slices for the estimation of the geometries of the ring to be utilized as input for the CAD-analytical model presented in this paper" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003262_tmag.2019.2900447-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003262_tmag.2019.2900447-Figure5-1.png", + "caption": "Fig. 5. Model of the rack-and-pinion gear for production: (a) pinion and (b) rack.", + "texts": [ + " 4 shows the states; the pinion gears are fixed, and the change in the magnetic flux density can be predicted even when only the rack gear is moved. This shows that the change in the magnetic flux density can be determined as the gear moves over time. Therefore, the torque and force values can be predicted using the magnetic flux density at each point. The results, shown in Figs. 3 and 4, confirm that the proposed analytical method provides fast and accurate results for the analysis of magnetic rack\u2013pinion gears. Table II lists the specifications of the manufactured rack\u2013 pinion gear. Fig. 5 shows the 3-D model of the magnetic rack-and-pinion gear intended for production. Fig. 6 shows the configured set of experiments for the manufactured rack\u2013pinion gear. The testing apparatus comprises a handle to measure the maximum torque values corresponding to each air-gap width of the rack\u2013 pinion gear. One part of the rack\u2013pinion gear is connected to the shaft of the torque sensor and the other one is connected to the load cell. A torque is generated if the part connected to the pinion gear shaft is rotated and the remaining portion is fixed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000023_13552540910925036-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000023_13552540910925036-Figure3-1.png", + "caption": "Figure 3 Schematic of sintering direction", + "texts": [ + " The laser machine used in this experiment was HJ-4 5KW cross-flow CO2 laser machine, supplied by Shanghai Institute of Optics and Fine Mechanics. DPSF-3 powder feed machine was utilized to feed the powders along the laser beam coaxially. The thick-wall parts will deposit on a 100mm \u00a3 50mm \u00a3 50mm stainless steel substrate, which were placed on TK13400B numerical controlled working table. The substrate was preheated before laser shaping so that very good metallurgy bonding between substrate and the primary metal layers could be obtained. Also the thermal stress occurred at the interface can be partly suppressed. Figure 3 shows the schematic of the shaping/building direction. The motion of the working table was controlled by the computer. Argon gas was used to deliver the metal powders to prevent the melt pool from oxidizing at a hightemperature during processing. Several metal thick-wall parts with size of 60mm \u00a3 20mm \u00a3 10mm were prepared in our experiments. The laser shaping parameters are shown in Table I. Analysis of all samples employed standard metallographic techniques for sectioning and polishing. XY, YZ, andXZ section specimens were photographed using a NEOPHOTO21 light optical microscope" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002933_icelmach.2018.8507125-Figure11-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002933_icelmach.2018.8507125-Figure11-1.png", + "caption": "Fig. 11. Acceleration distribution at 2f = 533.3Hz (a) Radial acceleration; (b) Mode shape at 529Hz.", + "texts": [ + " After taking the MMF phase compensation strategy, the amplitudes of the resonance peaks hardly decrease but see a slight rise at some frequencies including 6f. This surprising result indicates that the MMF compensation measure is not able to suppress the additional vibration and resonance that is caused by OC fault. This can also be proved by the acceleration waveform Fig. 10(a). Take 2f (533.3Hz) and peak 6f (1600Hz) for example, the acceleration distributions before and after OC fault are illustrated in Fig. 11 and Fig. 12. They are compared with the corresponding mode shapes in the vicinity of the frequencies. The region between position 0\u00b0 and 90\u00b0 is where the faulty unit 1 is located. Mode shape at 533.3Hz combines the circumferential and axial parts, so the acceleration distribution at this frequency tends to be irregular. It is obvious that the distribution is basically the same as the corresponding mode shape at 1600Hz. There is a big hump between position 0\u00b0 and 90\u00b0 (Unit 1) at 1600Hz under phase A1 failure because the acceleration amplitude at this frequency is largely dependent on the rotating speed and output torque" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001626_s0219878910002117-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001626_s0219878910002117-Figure3-1.png", + "caption": "Fig. 3. 2D orthogonal spring model.", + "texts": [ + " (A3) Contact position, direction of contact nor- mal, and local curvature at the contact point are known. (A4) Initial grasp configuration is in wrench equilibrium. (A5) Infinitesimal configuration displacement of objects occurs due to external disturbance. (A6) The relationship between displacement of fingertip position and reaction force can be replaced with a two-dimensional spring model. In Assumption (A6), the spring stiffness is denoted by K = diag[kx, ky] and is fixed along the axes of the fingertip coordinate frame (Fig. 3). In Assumption (A4), the spring is compressed at initial configuration and generates initial fingertip force f = [fx, fy]T . The relation of compression xf0 and the force is represented as f = Kxf0. Assumption (A6) means that rotational effect of the finger is much smaller than the translational effect and rotational displacement of the finger is much smaller than the translational displacement. These conditions are satisfied if we use one of the following hands: (i) each finger joint is constructed with prismatic joints, (ii) each finger has redundant angular joints and is totally controlled like prismatic finger" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003893_0954406219888240-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003893_0954406219888240-Figure5-1.png", + "caption": "Figure 5. The o-UV projection of the universal joint.", + "texts": [ + " Range III (04 4 p/2 and -p/2< 4 0): in order to avoid interference, the line B1B2 cannot intersect with Side 1 and Side 2. When the point P is on the projection of Side B (line B1B2) in U-V plane, as shown in Figure 8, the upper hinge has interference with the bottom hinge. Therefore, it is necessary to discuss the critical conditions. Based on equation (4), the function of line B1B2 in the o-UV plane is V d 2 cos sin \u00bc U c 2 cos d 2 sin sin d sin cos \u00f09\u00de In turn sin cos V\u00fe sin U c 2 cos sin d 2 sin \u00bc 0 \u00f010\u00de In Figure 5, the value of point P is P \u00bc d 2 , c 2 \u00f011\u00de When the point P is on the projection of line B1B2, this situation is called the critical position. Taking the value of point P into equation (10), its expression is rewritten as c 2 sin cos \u00fe d 2 sin c 2 cos sin d 2 sin \u00bc 0 \u00f012\u00de Based on equation (12) and the range of two Euler angles in this part, the critical line can be drawn in ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2q 0 1 If there is no intersection between line B1B2 with two sides and referencing the critical positions in Figure 13, the values of two Euler angles should locate up the critical line as shown in Figure 12" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001319_icelmach.2012.6350265-Figure10-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001319_icelmach.2012.6350265-Figure10-1.png", + "caption": "Fig. 10. Description of the mesh in the air gap.", + "texts": [ + " IRV c (4) where \u03a6 is the flux vector computed by the finite element software in every coil, V and I are the voltage and current vectors of the circuit branches, cR is the resistance matrix of the circuit. The speed \u2126 is assumed to be constant. The angular step \u0394\u03b8 has to be related to the size of the mesh in the rotating air gap. Authors have already study movement for field computations [6]. This step must be connected to the size of an element in the air gap in order to limit potential numerical instabilities (in particular for the calculation of electromagnetic torque) due to mesh noise. The mesh description of the entire machine is shown in Fig. 9 and in Fig. 10 for the air gap. Mesh in the stator must be particularly refined to take into account eddy currents in the solid iron stator. C. 12BImplementation of the brush/segment contact resistance and the arc voltage in the coupled circuit As mentioned in the previous section, the knowledge of the contact brushes/segment is essential for the study of commutation. The brush-segment contacts are modeled in the circuit by Nb \u00d7 Ns switches as shown on Fig. 8 where Nb and Ns are respectively the number of brushes and segments" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002636_j.mechmachtheory.2018.05.003-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002636_j.mechmachtheory.2018.05.003-Figure2-1.png", + "caption": "Fig. 2. Radius of the lower rotor lobe R lc of the circolima\u00e7on and lima\u00e7on-to-circular machines.", + "texts": [ + " As such, the rotor apices can be equipped with grooves so that apex seals can be mounted to prevent leakage and improve the machine\u2019s efficiency. Whilst the radial distance from the rotor centre point to the lobe of the lima\u00e7on-to-lima\u00e7on machine varies as per Eq. (2) , for the circolima\u00e7on and the lima\u00e7on-to-circular machines, the centre point, O r , of each rotor circular lobe is placed on the Y r -axis at a distance k , from the pole m , set by the designer. As such, the radius, R lc , of the rotor lobe is calculated, as suggested by Fig. 2 , as follows R lc = \u221a ( L \u2212 C ) 2 + ( k ) 2 (5) In the following sections, the forces and pressures, as well as the motion of the seal will be investigated. 3. Apex seal kinematics model formulation The apex seal is modelled as a simple spring-mass-damper system in which the seal is positioned inside a groove measured W g in width. A spring is attached to the inner end of the groove and one end of the seal; the other end of the seal pushes against the housing. The cross-sectional area of the seal is almost rectangular with one rounded end, dimensions and geometry of which is depicted in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003473_ab2dbb-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003473_ab2dbb-Figure2-1.png", + "caption": "Figure 2.The schematic diagramof the laser-MIGhybridweld.", + "texts": [ + " The basemetal is cut into specimens with the dimensions of 240\u00d7120\u00d720 mm. JM-68with 1.2 mm diameter is selected as fillingwire, whose chemical compositions are similar to that of basemetal, as shown in table 2. It contains appropriate alloying elements such asCr andTi, which can reduce splatter, refine grain and increase the impact toughness of weld bead. As shown infigure 1, thewelding system consists of KUKARobotKR 30HA, TruDisk-12 000 disk Laser, and Fronius TPS-5000weldingmachine is employed in this experiment with comprehensive performance. Figure 2 shows the schematic diagramof laser-MIGhybridwelding andmolten pool, which is influenced by both laser and arc. After thewelding process, theweldment ismachined along the cross-section of theweld bead to investigate themicrostructure and poremorphology using LeicaDMILMmetallurgicalmicroscope. 2.3.Weld process The diagramof thewelding groove is displayed infigure 3, which has 40\u00b0with a root face of 5 mm. In addition, the gap of the blunted edge is 0.8 mm. 99.99%pureArgon is used as shielding gas, whose flow rate is 15 L min\u22121, which is appropriate for the experiment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001876_tia.2017.2669326-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001876_tia.2017.2669326-Figure9-1.png", + "caption": "Fig. 9. Space vectors generated by the conventional 2.", + "texts": [ + " Note that the THD of the currents for configurations 1 and 2 was approximately 13 times lower than conventional 1, because the phase currents of both proposed configurations are sinusoidal, while conventional 1 generates a waveform of asymmetric phase current, as illustrated in Fig. 8. Compared to conventional 2, configurations 1 and 2 had a THD reduction of approximately 64%. This decrease occurred because the proposed topologies synthesize the reference voltage V \u2217 using the closest voltage vectors, unlike the conventional 2, since this structure generates only four voltage vectors that are phase shifted to each other by 90\u25e6, as illustrated in Fig. 9. Consequently, there is a reduction in the ripple of the input current in the proposed topologies, which results in improved THD levels. Regarding conventional 3, configurations 1 and 2 had a small advantage in terms of THD, since they generate a pulsating ripple pulsating ripple current with lower di/dt (current slew rate) during switching transition compared to conventional 3. Unlike configuration 2, in scenario 1, only configuration 1 presented a competitive efficiency in terms of total losses when compared to conventional 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002062_978-3-319-61431-1_2-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002062_978-3-319-61431-1_2-Figure2-1.png", + "caption": "Fig. 2. Schematics of a planar cable robot with 3 cables.", + "texts": [ + " 2, the model of a single cable and its winder, undergoing transverse deformation in a plane is considered. Based on Lagrange approach, a dynamic model is derived. In Sect. 3, the model of a planar robot with three or more cables is considered. The DAE model is developed and then reduced. In Sect. 4, some simulation results are presented and discussed. The model derived with Maple and the simulation with Matlab-Simulink are available online1. In this section, we focus on the an elementary constitutive element of the planar robot depicted in Fig. 2, namely, one single cable winded at one side and submitted at the other side to an external force. Up to four deformation fields can be considered when modeling a deformable beam under Euler-Bernoulli assumption [12]. Herein, the cable subjected to sagging is considered as a perfectly flexible and inextensible 1-dimensional body. In the current study, the only deformation field of interest is the transverse deformation in the plane of motion. The final geometry of the cable will be given as the composition of three steps: unwinding, shaping and rotation", + " The entries of the 1\u00d7 (N +2) line matrix \u0393k Qk correspond to the generalized forces acting on the cable. Denoting pk = \u2202Tk \u2202q\u0307k = q\u0307Tk Mk, the line matrix of generalized momentum, the model can be rewritten under the following state-space representation: p\u0307k = Ck + \u0393k Qk \u2212 Gk (9) q\u0307k = M\u22121 k pTk (10) where Ck = \u2202Tk \u2202qk = [ 1 2 q\u0307Tk \u2202Mk \u2202qk1 q\u0307k . . . 1 2 q\u0307Tk \u2202Mk \u2202qkN+2 q\u0307k ] (11) Gk = \u2202Vk \u2202qk = [ \u2202Vk \u2202qk1 . . . \u2202Vk \u2202qkN+2 ] (12) A planar cable robot operated by several cables is now considered as presented in Fig. 2. Its platform is considered as a punctual mass m located at point P of coordinates (xP , yP ) in the global reference frame. The number of cables in this example is three but the presented method is applicable to any number of cables. The generalized coordinate vector q for the system includes the two parameters of the mobile platform and the n sets of parameters relative to each cable. The column vector q can be written symbolically as q = [ xP yP qT1 . . . qTn ]T (13) which corresponds to n(N + 2) + 2 non independent parameters" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002143_1.4038869-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002143_1.4038869-Figure6-1.png", + "caption": "Figure 6: Amplification factor frequency mistuning", + "texts": [ + "4038869 Copyright (c) 2017 by ASME Downloaded From: http://turbomachinery.asmedigitalcollection.asme.org/ on 01/28/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use Insert ASME Journal Title in the Header Here TURBO-17-1127; Gross, Page -14- First, only frequency mistuning is considered. It appears to be common practice [19, 20 and 21] to describe the blade eigenfrequency variation by a normal distribution. In this study, a standard deviation between 0.5% and 5% is specified for the modal mistuning parameters \u03b4b,n. Figure 6 shows the amplification factor without aerocoupling (upper part, a) and with aerocoupling (lower part, b). The maximum value without aerocoupling is 2.4 for engine order 0 and a mistuning level for 2% STD. This value is well below the Whitehead limit. For higher engine orders, the AF drops to values between 1.5 and 2 and rises again towards higher values starting with EO26. With aerocoupling, the maximum amplification factor of 2.24 evolves at an engine order of 4. This EO predominantly excites ND4, where the aerodynamic damping has a sharp increase due to an acoustic resonance, see Figure 4", + " Figure 9 illustrates the global distribution of the AF over the EOs and the frequency mistuning level for large excitation mistuning (FM+EM50) and confirms that EM has a stronger impact on EOs with high wave numbers. Compared to the case with tuned EO Acc ep te d Man us cr ip t N t C op ye di te d Journal of Turbomachinery. Received August 21, 2017; Accepted manuscript posted January 05, 2018. doi:10.1115/1.4038869 Copyright (c) 2017 by ASME Downloaded From: http://turbomachinery.asmedigitalcollection.asme.org/ on 01/28/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use Insert ASME Journal Title in the Header Here TURBO-17-1127; Gross, Page -18- excitation, illustrated in Figure 6, the overall sensitivity to excited EO decreased, i.e., the extreme values at EO0 and EO31 are gone and feature now similar magnitude as in the midfield of EOs. This phenomenon seems plausible, since the excitation force mistuning leads to a distortion of the tuned EO excitation and causes stronger participation of different traveling wave modes close to the excited frequency. In the aeroelastically coupled case, a comparatively high amplification factor (>4, which is much higher than the Whitehead limit) is reached for EO16 (cf", + "4038869 Copyright (c) 2017 by ASME Downloaded From: http://turbomachinery.asmedigitalcollection.asme.org/ on 01/28/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use Insert ASME Journal Title in the Header Here TURBO-17-1127; Gross, Page -25- Figure Captions List Figure 1: Variation of excitation force Rotor 2 Figure 2: Finite-Element model Rotor 2 Figure 3: Upper subfigure: dispersion diagram of Rotor 2; lower subfigure: zoom into Mode F1 Figure 4: Aerodynamic influence coefficients Figure 5: Comparison of parent FE model vs. ROM for FRF @ EO5 Figure 6: Amplification factor frequency mistuning Figure 7: Amplification factor excitation mistuning Figure 8: Comparison of AFs @ STD=10% and 50% excitation force mistuning and varying frequency mistuning Figure 9: Amplification factor frequency mistuning @ 50% excitation mistuning Figure 10: Amplification factor of excitation force mistuning @ 2% STD frequency mistuning Figure 11: Sections @2% frequency mistuning without aerodynamic coupling Figure 12: Mistuning pattern 469 \u2013 maximum Figure 13: Mistuning pattern 421 - minimum Figure 14: Amplification factor excitation mistuning for maximum and minimum pattern Acc ep te d Man us cr ip t N ot C op ye di te d Journal of Turbomachinery" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002949_detc2018-85642-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002949_detc2018-85642-Figure2-1.png", + "caption": "Figure 2 Experiment Devices (a: MakerGear M2e FDM 3D printer; b: Pronterface software; c: HOBO data logger software; d: HOBO UX 120 Plug Load Data Logger)", + "texts": [ + " Three process parameters are included as control variables and AM energy consumption is the response. Each of these parts was then sent to 3D scanner to generate point cloud for geometric accuracy evaluation. Table 1 presents the process parameters and their corresponding levels. The part shown in Figure 1 is used for experiments, and the design is similar with the one used in [27]. Using full factorial design, 27 cases with 2 replications for each case are developed and conducted by MakerGear M2e FDM 3D printer (Fig. 2a). The HOBO UX 120 Plug Load logger (Fig. 2 d) is used to collect the energy consumption data during the AM process. This device was used as a data logger to record the instant energy consumption data with sampling frequency of 1 Hz. The energy consumption data were collected through HOBO data logger software (Fig. 2 c) during the entire printing operation, which excludes energy consumption to preheat the nozzle and the printer bed. Table 1 Experimental Design (Levels are selected based on recommended working ranges) Process Parameter Levels Low Medium High \ud835\udc651: Resolution (mm) 0.1 0.2 0.3 \ud835\udc652: Printing Speed (mm/s) 30 55 80 \ud835\udc653: Extruder Temperature (\u00b0C) 200 215 230 After each part has been fabricated, the Taylor-Hobson Talysurf CLI 2000 Gauge System (Fig. 3) is used to collect the point cloud data to characterize the geometric accuracy of the 3D printed parts" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003921_b978-0-12-803581-8.11722-9-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003921_b978-0-12-803581-8.11722-9-Figure5-1.png", + "caption": "Fig. 5 Hydrodynamic clutch with ER fluid: 1 \u2013 turbine, 2 \u2013 housing, 3 \u2013 pump, 4 \u2013 blade, 5 \u2013 sealing ring, 6 \u2013 slip ring for supplying high voltage.", + "texts": [ + " The width of the gaps between the rings is 1 mm each. The way of action of a cylindrical viscous clutch with ER fluid is the same as in the case of a disk viscous brake with ER fluid. The fluids used in the clutch were LID3354S and ERF#6. Hydrodynamic clutch with flat radial blades and an ER fluid consists of two rotors without an outer ring with a radial-axial flow and the housing connected to the pump\u2019s rotor (Olszak et al., 2019b). The active diameter of the rotors is 116 mm and there are 38 blades in each rotor, Fig. 5. The channels created by the blades and walls of the rotors have a wedge cross-section, so that the smaller distance between blades is 2.5 mm and the largest distance is 8 mm. The pump blades and the turbines alike, are made of 1 mm thick steel plate and are electrodes. The blades were electrically isolated from each other by the rotors\u2019 walls made of nylon and placed in the slots of the rotors\u2019 housing. The blades have edgings which go through notches in the rotors\u2019 housing, designed to connect the blades to the insulated slip rings" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003903_s11370-019-00306-6-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003903_s11370-019-00306-6-Figure4-1.png", + "caption": "Fig. 4 Kinematic and dynamic coordinate system diagram of the microinstrument", + "texts": [ + " The cable tension can be obtained by tension sensors, the joint motion parameters can be obtained by kinematics, but the joint integrated (2) ( 1T1 \u2212 1T2 ) r0 = Dyna1 +Mf1 + JT 1 F1 Dyna1 = h1 ( \ud835\udf031 ) ?\u0308?1 + c1 ( \ud835\udf031, ?\u0307?1 ) ?\u0307?1 + g1 ( \ud835\udf031 ) Mf1 = 1Ff1r0 + 1Ff2r0 + \ud835\udf0ff1 (3)F1 = ( JT 1 )\u22121[(1T1 \u2212 1T2 ) r0 \u2212 Dyna1 \u2212Mf1 ] friction torque cannot be accurately obtained. If the motion parameters and cable tension are known, the contact force can be estimated by a neural network prediction model. The kinematic coordinate system established by the DH method is shown in Fig.\u00a04, and the link parameters are given in Table\u00a03. Since the microinstrument is a series structure, the two yaw joints are independent, and the tool coordinate system is established at the end of portions a and b of the forceps. The homogeneous transformation matrix between adjacent links is i\u22121Ai = Az,dAz, Ax,aAx, = \u23a1 \u23a2 \u23a2 \u23a2 \u23a3 cos i \u2212 sin i \u22c5 cos i sin i \u22c5 sin i ai \u22c5 cos i sin i cos i \u22c5 cos i \u2212 cos i \u22c5 sin i ai \u22c5 sin i 0 sin i cos i di 0 0 0 1 \u23a4 \u23a5 \u23a5 \u23a5 \u23a6 Fig. 3 Dynamic analysis diagram of the cable-driven joint unit reset motor traction motor z0 x0 y0 \u03b8 1 P1 displacement -P1 tension sensors 1T2 1T1 set of pilot wheels 1Ff 2 1Ff 1 1Ft2 1Ft1 \u03c4 f 1\u03c4 1 r0 initial position F1z F1x F1y g 1 3 According to the coordinate system, the transformation matrix between the base coordinate system and the tool coordinate system is obtained, and the forward kinematics are obtained by where si \u225c sin i, ci \u225c cos i , m = a, b", + " Assuming that the target rotation angles of portions a and b of the forceps are \u03b82a and \u03b82b, respectively, when the pitch joint is rotated \u03b81, the actual rotation angles of portions a and b of the forceps are \ufffd 2a = 2a + 1r0\u2215r1 and \ufffd 2b = 2b + 1r0\u2215r1 , respectively. The corresponding motor output displacements are P1 = r0 1 , P2a = r1 \ufffd 2a and P2b = r1 \ufffd 2b . Considering both the kinematics and the decoupled motion, the complete kinematics of the microinstrument are as follows: The dynamic coordinate system is shown in Fig.\u00a04. Let the generalized coordinate be q = [q1 q2 q3]T = [\u03b81 \u03b82a \u03b82b]T; the centroid positions are 0rc1, 1rc2a, 1rc2b, the masses are m1, m2a, m2b, the pseudo inertia matrices are J1, J2a, J2b, the joint driving torque is \u03c4 = [\u03c41 \u03c42a \u03c42b]T, the force Jacobian matrix is J T, the external force is F = [Fmx Fmy Fmz]T, the joint friction torque is Mf, and the homogeneous coordinate of gravity acceleration g\u0304T = [ 0 g 0 0 ] . Since the two yaw joints are independent and have high similarity, a dynamic model was established for one of them; a 2-DOF mechanism composed of a yaw joint (forcep a) and the pitch joint was used as an example" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000917_20110828-6-it-1002.02534-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000917_20110828-6-it-1002.02534-Figure5-1.png", + "caption": "Fig. 5. K(1) gain surface", + "texts": [], + "surrounding_texts": [ + "J. M. Biannic and P. Apkarian. Missile autopilot design via a modified LPV synthesis technique. Aerospace Science and Technology, 3(3):153\u2013160, 1999. ISSN 1270-9638. Stephen Boyd, Laurent El Ghaoui, Eric Feron, and Venkataramanan Balakrishnan. Linear matrix inequalities in system and control theory. SIAM, 1994. M. Chilali and P. Gahinet. H-infinity design with pole placement constraints: an LMI approach. IEEE Transactions on Automatic Control, 41(3):358\u2013367, 1996. A. Fujimori, F. Terui, and P. N. Nikiforuk. Flight control designs using v-gap metric and local multi-objective gain-scheduling. In AIAA Guidance, Navigation and Control Conference, 2003. AIAA 2003-5414. P. Gahinet, A. Nemirovskii, A.J. Laub, and M. Chilali. The lmi control toolbox. In Proceedings of the 33rd IEEE Conference on Decision and Control, pages 2038\u2013 2041, 1994. S. L. Gatley, D. G. Bates, M. J. Hayes, and I. Postlethwaite. Robustness analysis of an integrated flight and propulsion control system using \u00b5 and the \u03bd-gap metric. Control Engineering Practice, 10(3):261\u2013275, 2002. ISSN 0967-0661. Jung-Yub Kim, Sung Kyung Hong, and Sungsu Park. LMI-based robust flight control of an aircraft subject to CG variation. International Journal of Systems Science, 41(5):585\u2013592, 2010. C. H. Lee, T. H. Kim, and M. J. Tahk. Design of missile autopilot for agile turn using nonlinear control. In Proceeding of the KSAS Conference, pages 667\u2013670, 2009. D. J. Leith and W. E. Leithead. Survey of gain-scheduling analysis and design. International Journal of Control, 73(11):1001\u20131025, 2000. Kenneth H. McNichols and M. Sami Fadali. Selecting operating points for discrete-time gain scheduling. Computers & Electrical Engineering, 29(2):289\u2013301, 2003. ISSN 0045-7906. P. K. Menon and E. J. Ohlmeyer. Integrated design of agile missile guidance and autopilot systems. Control Engineering Practice, 9(10):1095 \u2013 1106, 2001. ISSN 0967-0661. Andy Packard. Gain scheduling via linear fractional transformations. Systems & Control Letters, 22(2):79\u2013 92, 1994. ISSN 0167-6911. Wen qiang Li and Zhi qiang Zheng. Robust gainscheduling controller to LPV system using gap metric. In International Conference on Information and Automation, pages 514\u2013518, 2008. Wilson J. Rugh and Jeff S. Shamma. Research on gain scheduling. Automatica, 36(10):1401\u20131425, 2000. ISSN 0005-1098. David Saussie, Lahcen Saydy, and Ouassima Akhrif. Gain scheduling control design for a pitch-axis missile autopilot. In AIAA Guidance, Navigation and Control Conference, 2008. AIAA 2008-7000. C. Scherer, P. Gahinet, and M. Chilali. Multiobjective output-feedback control via LMI optimization. IEEE Transactions on Automatic Control, 42(7):896\u2013 911, 1997. J. S. Shamma and James R. Cloutier. Gain-scheduled missile autopilot design using linear parameter varying transformations. Journal of Guidance, Control, and Dynamics, 16(2):256\u2013263, 1993. J.S. Shamma and M. Athans. Gain scheduling: potential hazards and possible remedies. IEEE Control Systems Magazine, 12(3):101\u2013107, Jun. 1992. Spilios Theodoulis and Gilles Duc. Missile autopilot design: Gain-scheduling and the gap metric. Journal of Guidance, Control, and Dynamics, 32(2):986\u2013996, 2009. Ajay Thukral and Mario Innocenti. A sliding model missile pitch autopilot synthesis for high angle of attack maneuvering. IEEE Transactions on Control Systems Technology, 6(3):359\u2013371, 1998. D.P. White, J.G. Wozniak, and D.A. Lawrence. Missile autopilot design using a gain scheduling technique. In Proceedings of the 26th Southeastern Symposium on System Theory, pages 606\u2013610, Athens, 1994. Kevin A. Wise and David J. Broy. Agile missile dynamics and control. Journal of Guidance, Control and Dynamics, 21(3):441\u2013449, 1998. G. Zames and K. El-Sakkary. Unstable systems and feedback: The gap metric. In Proceedings of the 18th Allerton Conference, pages 380\u2013385, 1980. Paul Zarchan. Tactical and Strategic Missile Guidance. Progress in Astronautics and Aeronautics Series, 2007. K. Zhou and J.C. Doyle. Essentials of Robust Control. Prentice Hall, 1997. P. H. Zipfel. Modeling and Simulation of Aerospace Vehicle Dynamics. AIAA Education Series, 2000." + ] + }, + { + "image_filename": "designv11_33_0002701_23311916.2018.1493672-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002701_23311916.2018.1493672-Figure4-1.png", + "caption": "Figure 4. Mathematical formulation of film thickness using trigonometric relations.", + "texts": [ + " The pad radius (Rp) along with clearance parameter (Cp) facilitates the radial adjustment of pads: LM \u00bc 2 Rp \u00fe Cp sin \u03b4 2 (2) Figure 3. Representation of the mathematical approach applied to develop modified film thickness: (1) profile of multipad bearing, (2) negative radial adjustment, (3) positive radial adjustment, and (4) positive tilt adjustment. Page 5 of 15 Xj \u00bc Yj \u00bc 900 \u03b4 2 (3) The leading edges of four pads are subjected to tilt displacement resulting in a non-uniform bearing configuration. Trigonometric relations are applied to model the tilt orientations of bearing elements as shown in Figure 4. The tilt angle provided is considered with respect to the bearing center BC\u00f0 \u00de and tilted pad center position TC\u00f0 \u00de . The radial distance of the tilted pad about the bearing center varies along the pad curvature and is measured using Equation (4). The nondimensional modified film thickness equation developed by incorporating the radial adjustments and tilt angles in either positive or negative directions is given as follows: ZL \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi Rp \u00fe Cp 2 \u00fe LM\u00f0 \u00de2 2 Rp \u00fe Cp LM cos \u03b1\u00fe Y\u00f0 \u00de q (4) where \u03b1 is the pad angle h \u00bc ZL R\u00f0 \u00de=C\u00f0 \u00de \u00fe \u03b5 cos \u03b8 0 (5) \u03b8 0 \u00bc \u03b8 \u03c8 \u03b1=2\u00f0 \u00de\u00f0 \u00de The generalized Reynolds equation governs the pressure variation in the bearing clearance spaces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003931_s11012-019-01101-4-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003931_s11012-019-01101-4-Figure2-1.png", + "caption": "Fig. 2 Trailling arm model representation, with 3 DOF", + "texts": [ + " More specifically, if both local stiffness and damping are considered for changing, each row of K has only 2 non zero elements. In these cases, each matrix XT is r 2. However, if the rows of XT are diminished to 2, this matrix is square and its inversion can be computed. As a consequence, the eigenstructure assignment can be achieved easier for 2 eigenvectors. Otherwise, more careful numerical solutions can be required. 2.3 Landing gear model and pseudo matrix of gain The proposed methodology is demonstrated herein for the landing gear model illustrated in Fig. 2. The state vector is defined by Eq. (25), where, _y\u00f0t\u00de and y(t) are, respectively, the lateral velocity and displacement of the landing gear center of gravity, _w\u00f0t\u00de and w\u00f0t\u00de are, respectively, the YAW angular velocity and angular displacement, and the a is the tire slip angle. The 5 5 dynamic matrix (Eq. 26) is written in the physical coordinate system, which makes simpler to define the matrix of the pseudo control gain. The input and output matrices are given by Eq. 27 and additional details can be found in reference [4]", + " In particular for the present system, the values of these incremental properties can be computed considering the following matrix K \u00bc Dky Dcy 0 0 0 0 Dkw Dcw \u00f028\u00de which implies to consider two pseudo inputs u\u00f0t\u00de \u00bc fFy TwgT and outputs y\u00f0t\u00de \u00bc fy0\u00f0t\u00de _y0\u00f0t\u00de w\u00f0t\u00de _w\u00f0t\u00degT such as Fy\u00f0t\u00de \u00bc Dkyy\u00f0t\u00de \u00fe Dcy _y\u00f0t\u00de and Tw \u00bc Dkww\u00f0t\u00de \u00fe Dcw _w\u00f0t\u00de, which in practice allow to compute the physical properties (incremental stiffness and damping) to assign the desired eigenvector. For this introduced context, to redesign a landing gear, the proposed methodology can be applied for two main conditions: (1) to reduce vibrations in a nearshimmy condition; (2) to eliminate the shimmy in a speed Vshi. For the case (1), specially to low damped systems, if V ! Vshi, a stable landing gear can exhibit a high vibrating behaviour, and redesign can improve it in this sense. To illustrate these idea, it is considered an initial landing gear (see Fig. 2) preliminarily defined through the following physical and geometrical properties: lateral stiffness ky \u00bc 1:5 106 N/m, yaw stiffness kw \u00bc 1:0 105 N/m rad, lateral damping cy \u00bc 520 N s/m, yaw damping cw \u00bc 520 N m s/rad, mass m \u00bc 200 kg, moment of inertia Iz \u00bc 81 kg m2, relaxation length r \u00bc 1:2 m, trailing arm total length et \u00bc e\u00fe q \u00bc 0:6 m. Based on a classical eigenvalue value analysis, it is established the stability map through Fig. 3, where the blank is stable. Considering V \u00bc 36 m/s, the pre-designed system is unstable and a time domain integration of Eqs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001855_s11249-017-0818-8-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001855_s11249-017-0818-8-Figure3-1.png", + "caption": "Fig. 3 Diagram showing the geometric quantities used to define the simulated periodic surface textures. Ri is the inner radius, Ro is the outer radius, Rc is the radius to the center of the texture, Rt is the radius of the texture, and u is the periodic spacing of the surface textures. Boundary conditions are also shown", + "texts": [ + " 2) are at z = 0: vh \u00bc rX and vr \u00bc vz \u00bc 0 at z \u00bc h : vh \u00bc vr \u00bc vz \u00bc 0 \u00f04\u00de which from integrating conservation of momentum (see Supplemental Information) gives the velocity field as vr \u00bc 1 2g op or z2 zh vh \u00bc 1 2g 1 r op oh z2 zh \u00fe rX h z h : \u00f05\u00de If the gap height h\u00f0r; h\u00de is prescribed (and boundary conditions for the pressure in the r and h directions are also prescribed), then Eq. (3) can be solved explicitly for the pressure and used in Eq. (5) to calculate the velocity field. To solve Eq. (3), boundary conditions on the pressure must be imposed, as shown in Fig. 3. In the h direction, we use periodic boundary conditions, given as p h \u00bc u=2\u00f0 \u00de \u00bc p h \u00bc u=2\u00f0 \u00de h3 h \u00bc u=2\u00f0 \u00de op oh jh\u00bc u=2 \u00bc h3 h \u00bc u=2\u00f0 \u00de op oh jh\u00bcu=2: \u00f06\u00de where u is the span in the h direction of the periodic texture, and the h3 front factor in front of the pressure gradient accounts for the possibility of a different gap height at the start and end of the periodic cell. At the inner radius, we impose a Neumann boundary condition, given as op or jr\u00bcRi \u00bc 0 \u00f07\u00de to impose symmetry and regularity in the physical space [33]", + "4 Pa s, and u = p/5 rad, X = 10 rad/s, h1 = 1 mm, and h2 = 0.5 mm. The expected exponential decay in the error is observed [28\u201330], verifying the numerical method. To validate the predictive capabilities, we compare to experiments with textured disks. In our previously measured experimental results for a surface-textured thrust bearing [1], the surface textures are cylindrical holes cut at an angle b, which creates an elliptical top profile. We model our surface textures with an elliptical top profile, similar to the experiments. Figure 3 shows the geometric quantities used to define the surface textures; the finite inner radius is needed so that the 1/ r terms do not diverge. We found that the normal force converged when Ri\\ 0.02 mm (Ri/Ro\\ 0.1%); details on this are given in Supplemental Information. The geometric values for all the simulated textures are given in Table 1. Examples of the simulated texture surfaces are given in Fig. 5. The boundary conditions in the h direction used to solve the Reynolds equation are given in Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002046_ijmmm.2017.088895-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002046_ijmmm.2017.088895-Figure3-1.png", + "caption": "Figure 3 Cutting tool insert and the tool holder (see online version for colours)", + "texts": [ + " The worn tools were later analysed using a Gemini scanning electron microscope (SEM) for both morphological features and elemental variation. A Ti-6Al-4V solid bar of about 60 mm diameter and about 150 mm in length was used for the machining trials. The microstructure of the workpiece is shown in Figure 2, and the hardness of the workpiece was 358 \u00b1 25 HV0.5. The average elemental composition (mass percentage) of the workpiece is given in Table 1. Uncoated tungsten carbide tools CNMX1204A2-SM H13A manufactured by Sandvik Coromant along with the Sandvik tool holder were used for this study (shown in Figure 3). They had a rake angle of +15\u00b0,1 entry angle of 45\u00b0, corner chamfer of 50\u00b0, major cutting edge angle of 95\u00b0, and corner radius of 0.8 mm. Machining was performed under dry cutting conditions (DM) during LAM trials. A depth of cut of 1 mm and feed rate of 0.214 mm/rev was maintained for all trials. Three cutting speeds, 125 m/min, 150 m/min, and 175 m/min were selected.2 The length of cut was kept between 95 to 110 mm after which the wear on the flank face of the cutting insert was measured under an optical microscope" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001715_978-3-642-21922-1_6-Figure6.1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001715_978-3-642-21922-1_6-Figure6.1-1.png", + "caption": "Fig. 6.1 Two legged body", + "texts": [ + " As a result, capacity measure exploiting cubes to cover the attractor region and Lyapunov dimensions computed based on the Lyapunov exponents are used throughout this section. Lyapunov exponents, being the unique measures to determine the chaotic phase and cell to cell mapping leading to a strong tool to take the picture of the phase portraits, are employed in subsequent sections, associated with fractal dimensions in order to investigate the stability of nonlinear robotic systems. The dynamic model of a two legged robot (Fig. 6.1) is composed of a rigid body and a pair of legs whose weights are lumped within the robot body so that they can be modeled as weightless. In the system, W = Mg is the body weight, J the moment of inertia, O the hip joint, the length from center of mass c to O; L is the step length and S is the support shift giving the horizontal distance from joint O to the forward step. During the walking process, only one leg provides support at a time and the legs are lifted up to a maximum height h. The legged locomotion is considered to displace the body in a planar surface x-z with a velocity V " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002063_s00170-017-0656-8-Figure20-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002063_s00170-017-0656-8-Figure20-1.png", + "caption": "Fig. 20 An instantaneous cutting state", + "texts": [ + " In the processing of setting of the movement condition and the boundary condition, a rotary cutting movement with the velocity of 15.7 rad/sec and a linear feeding movement with the velocity of 30 mm/min are applied on the cutting tool. The initial temperature is set at 20 \u00b0C, the friction is 0.45, and the interface heat transfer coefficient is 30(SI); the tolerance of the contact boundary condition is 0.0002 mm.The cutting simulation could be carried out after the completion of the setup. An instantaneous state during processing is shown in Fig. 20. After completion of the calculation, the forces of the three directions are extracted as Fig. 21. As illustrated in Fig. 21, the three directions of cutting forces all present a rising trend. Among them, the X direction force increases gently, with a small rise from 150.8 N at the cutting-in point to 265.2 N at the cutting-out point. The Y direction force increases from 594.6 N at the cutting-in point to 852.3 N at the cutting-out point. The Z direction force increases dramatically, with a great rise from 276 N at the cutting-in point to 966 N at the cutting-out point" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003584_icra.2019.8793779-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003584_icra.2019.8793779-Figure1-1.png", + "caption": "Fig. 1: Example distance maps as the hand closes. Each red region on the hand generates one map (each red dot on the finger generates one pixel, map for circled finger link shown). Orange is further away, black touching.", + "texts": [ + " Broadly speaking, these representations capture the geometric relationships between the surfaces of the fingers and the object as they come into contact. By capturing data in a grid structure (rectangular set of samples per finger pad/palm) we can directly employ existing imagebased machine learning algorithms. More specifically, for every potential finger contact surface (typically the pad of each finger and the palm, but we can include all of the finger links if desired) we define a grid of points. For each of those points, we record the distance (and optionally, orientation) to the object (see Figure 1). All of these grids can be combined into a single 2D \u201cimage\u201d using tiling. Our two feature representations differ by how they calculate distance. In the first representation, called the signed *This work was supported in part by NSF grants CNS 1730126 and CNS 1659746 1E. Dessalene is with the department of Computer Science and Mathematics, George Mason University, 4400 University Drive, Fairfax USA edessale@gmu.edu 3Oregon State University, Collaborative Robotics and Intelligent Systems (CoRIS) ongyi,morrowj, Ravi" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000298_s10846-010-9465-0-Figure10-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000298_s10846-010-9465-0-Figure10-1.png", + "caption": "Fig. 10 Different reachable region depending on contact point and object configuration", + "texts": [ + " In the following sections, parameterization of boundary between mode i and j will be presented as zi, j(s) to identify modes. 3.6 Expression for Reachable Region For planning trajectories to realize the desired configurations of objects, it is required to identify where the object can reach (reachable region). In this section, a describing method is given for the reachable region for observation variable y( ) (corresponding to O ). Before this, let us consider the pushing task of a triangular object as an example (see Fig. 10). When the hand is touching the object at the left (, respectively right) leg, the reachable region of the object is restricted to the right-hand (, respectively left-hand) space of the current object position. This shows that the reachable region of an object changes according to at which part of object- (hand-) surface the contact takes place. In the general case, the reachable region (, say of object O ) is characterized by the following factors: (i) Set of objects forming a chain whose tip object is O and their arranging order (ii) Initial locations of the the objects in the chain (iii) Contact positions of any two mutually connected objects in the chain (i) is reflected in the observation on mode transitions since each mode corresponds to a chain" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001127_j.ijepes.2012.08.004-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001127_j.ijepes.2012.08.004-Figure3-1.png", + "caption": "Fig. 3. Half sectional end and sectional elevation", + "texts": [ + " (b) The value of the potential varies linearly between any two adjacent nodes on the element edges. (c) The value of the potential function in each element is determined by the order of the finite element. The order of the element is the order of polynomial of the spatial co-ordinates that describes the potential within the element. The potential varies as a quadratic function of the spatial coordinates on the faces and within the element. 4.2. Boundary conditions The details of the induction motor are shown in Fig. 3. In this analysis, the two-dimensional domain of core iron and winding chosen for modeling the problem is shown in Fig. 4 and the geometry is bounded by planes passing through the mid-tooth and the mid-slot. The temperature distribution is assumed symmetrical across two planes, with the heat flux normal to the two surfaces being zero. From the other two boundary surfaces, heat is transferred by convection to the surrounding gas. It is convected to the air\u2013gap gas from the teeth, to the back of core gas from the yoke iron" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003075_tasc.2019.2891534-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003075_tasc.2019.2891534-Figure6-1.png", + "caption": "Fig. 6 Magnetic cloud map and magnetic field line distribution in load operation. (a) Inverter power supply. (b) Sinusoidal power supply", + "texts": [ + "004 -4 -3 -2 -1 0 1 2 3 4 C ur re nt ( A ) Time (s) (a) 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 A m pl it ud e (A ) Harmonic order 2 3 4 5 6 7 8 9 10 0.00 0.05 0.10 0.15 0.20 (b) 1051-8223 (c) 2018 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. Fig. 4 Analysis of current waveform. (a) Current waveform. (b) Harmonic analysis Figure 6 shows the magnetic density and magnetic line distribution of the motor. The saturation magnetic density is 3.08T when the inverter is supplied, slightly higher than 3.04T when the sinusoidal wave is supplied. The loss of PMSM mainly includes copper loss, stator iron loss and eddy current loss of permanent magnet. When sinusoidal power is supplied, the copper loss is calculated according to equation (1). The copper loss is calculated at 34.77W. 2 13CuP I R (1) Where I is the current rms value; R1 is the phase resistance. When inverter power is supplied, the copper loss is calculated according to equation (2). The copper loss is calculated at 41.61W. 2 2 1 2 3 +3Cu k k k P I R I R (2) Where Rk is the harmonic equivalent resistance; k is harmonic number. Fig.6 (a) shows the iron loss curve of sine wave power supply, and the average value of iron loss is 11.15W. Fig.7 (b) shows the iron loss curve of inverter power supply, and the average value of iron loss is 15.87W. The formula of the eddy current loss of the permanent magnet is shown in equation (3). 2 2a V J P dv (3) Where J is the eddy current density of permanent magnet; is the conductivity of permanent magnet. As shown in Fig. 8, the average value of eddy current loss is 0.74W when sinusoidal power is supplied, and the average value of eddy current loss is 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000151_iemdc.2009.5075258-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000151_iemdc.2009.5075258-Figure1-1.png", + "caption": "Fig. 1. Schematic view of the motor", + "texts": [ + " Moreover general numerical studies on this special type of machine have not been carried out yet. Therefore a study has been made to find out, how helpful 2D- and 3D-FEM could be in our days, to solve the task of optimizing this special machine. Especially for small companies experimental field tests are not adequate and analytical design combined with modern numerical FEM software should be state of the art. Hybrid stepper motors work like a reluctance motor. They consist of at least two rotor stacks, which are placed left and right to an axially magnetized magnet (see fig. 1a and 1b). A special feature of these motors is that the two stacks are twisted against each other by tau2/2. Hence two quasi homopoles are formed. The magnet causes detent torque which normally should be minimized by an optimization but also helps to increase power density. For some special application it may be useful to maximize the detent torque to produce a strong self holding. First of all we start with a short analytical calculation to determine the expected air gap flux density and axial torque of the motor to have first reference values for further FEM simulations [7]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001045_jsea.2010.312134-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001045_jsea.2010.312134-Figure9-1.png", + "caption": "Figure 9. Deformation due to gravity.", + "texts": [], + "surrounding_texts": [ + "In this research, the proposed approach provides better formulation of cost-tolerance relationships for empirical data. BP network architecture of configuration 4-6-1 generates a suitable model for cost-tolerance relationship of R2 value 0.9997, there by eliminating errors due to curve fitting in case of regression fitting. And it also generates more robust outcomes of tolerance synthesis. The proposed non conventional optimization technique obtains an optimal solution better than that of simulated annealing [6] and Response surface methodology (RSM) [1].This study proposes a tolerance synthesis based on BP learning, a NSGA II based optimization algorithm and CAD interface, in order to ensure that the proposed values of controllable factors (tolerances) satisfies the assembly constraint, even before the start of manufacturing process. There by reducing scrap and rework cost." + ] + }, + { + "image_filename": "designv11_33_0001422_pime_conf_1964_179_264_02-Figure21-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001422_pime_conf_1964_179_264_02-Figure21-1.png", + "caption": "Fig. 21. I . Bearing diagram", + "texts": [ + " The film boundaries are taken to be at zero reference pressure, that is, at atmospheric pressure. Solution of a partial differential equation is required to establish the pressure distribution and the corresponding acceleration distribution with the assumed boundary conditions. The force on the journal is found from the resultant load of the pressure distribution. The steps in the calculation are summarized below. The calculation is made as far as convenient in non-dimensional terms. Consider a section of the film, as shown in Fig. 21.2, subtending a small angle S/l and of small axial width Sz. The circumferential length of the section is R Sj3. The film thickness h* at position /3 is small compared with R. Owing to the eccentricity of the journal, h* has changed to (h*+Sh*) at the other end of the section. Where the film is convergent, Sh* is negative. One of the assumptions made for calculation is that the oil is incompressible. If there had been flow of the oil across the faces of the section, the condition of continuity would have given the summation of volume flow as zero", + " This value is in the neighbourhood of the highest values likely to be reached within the viscous flow range on a turbine bearing of normal design, and therefore gives coefficients Proc Inrtn Mech Engrs 19646.5 cll, . . ., cz2 representative of the higher values which can be reached in practice. In calculations for particular cases involving other values of R,(C/D), the coefficients here calculated must be multiplied by R,(C/D). The results of the calculation are plotted as the 'wide bearing' curves of Fig. 21.3 for the range of eccentricity ratios 0.1 to 0-9. For a very narrow bearing, the approximations of the Ocvirk theory are used. The pressure gradients in the 1.5 1.0 c z w 0 LL LL w 0 0 0.5 0 -0.4 0 ECCENTRICITY RATIO Non-dimensional coefficients calculated for the case R,(C/D) = 1 Re = Reynolds number in film referred to bearing radial clearance = - PU(CI2). P Fig. 21.3. Force-acceleration coeficients for wide and narrow bearings Vol179 Pt 33 at Gazi University on March 2, 2016pcp.sagepub.comDownloaded from JOURNAL BEARING DYNAMIC CHARACTERISTICS-EFFECT OF INERTIA OF LUBRICANT 41 axial direction are taken as so steep compared with those in the circumferential direction that the acceleration of the lubricant corresponds to the axial pressure gradients alone. The pressure bearing film is taken as extending from p1 = 0 to p2 = 7~12. In equation (21.7), the terms in ap'/@ can be omitted", + " (21.22) and the coefficients in equations (21.11) and (21.12) are evaluated as For a narrow bearing in which L / R is a small quantity of the first order, the f coefficients are thus small quantities of the second order. Nevertheless the c coefficients, calculated from equation (21.18), have finite values, because the Ocvirk parameter s (y = s (&)' has finite values. Numerical calculations have been made for the same case, Re(C/D) = 1. The results are plotted as the 'narrow bearing' curves in Fig. 21.3. The acceleration coefficients thus calculated for very wide and very narrow bearings with R,(C/D) = 1 are superimposed on Fig. 21.3. It is seen that for very wide and for very narrow bearings, at the same eccentricity ratio, the coefficients are fairly near to one another, with the narrow bearing characteristics closer to symmetry than the wide bearing characteristics. It is reasonable to assume that for bearings of finite width the acceleration coefficients will lie within or close to the range between the wide bearing coefficients and the corresponding narrow bearing coefficients. The coefficients are much higher at low eccentricities than at high eccentricities", + " There would, however, have been great variation with LID ratio, the coefficients thus expressed becoming small quantities of the second order for a very narrow bearing. The coefficients cI1, . . ., ca2 are expressed relatively to horizontal and vertical axes, since in systematic investigation of rotor motion, these are the axes in which it is Proc Instn Mech Engrs 1964-65 generally convenient to write down the equations of motion of the rotor. For examining the physical character of the acceleration reactions, it is, however, useful to consider the maximum and minimum values of the acceleration coefficients in any orientation of axes. This is the object of Fig. 21.4, where maximum and minimum acceleration coefficients are plotted for wide and for narrow bearings. The angle of inclination of the principal axes of the acceleration terms is plotted in Fig. 21.5. COMPARISON OF DISPLACEMENT, VELOCITY, AND ACCELERATION TERMS VolZ79 Pt 3J at Gazi University on March 2, 2016pcp.sagepub.comDownloaded from 42 D. M. SMITH such a way that with simple harmonic vibration of frequency equal to the running speed, the maximum component disturbance forces due to displacement, velocity, and acceleration are in the same ratio as the corresponding displacement, velocity, and acceleration coefficients. Comparison of the acceleration coefficients as plotted in Fig. 21.3 with the displacement and velocity coefficients plotted in (I) shows that the highest acceleration coefficient at each eccentricity ratio is much less than the highest displacement and velocity coefficients. The acceleration coefficients for individual bearings are proportional to R,(C/D), but the value 1 for which the coefficients are plotted is not far short of the highest value likely to be attained in a bearing of turbine type with viscous flow in the film. Hence for vibration of frequency equal to or below the running speed, the acceleration force components are smaller than the displacement and velocity force components, so long as the bearing is operating under viscous flow conditions", + " An approximation has been made to the values of the ratio for a bearing with LID = 1 by interpolation of the c coefficients between those calculated for a very wide bearing and for a very narrow bearing at the same eccentricity ratio. The geometric mean has been taken for interpolation; this is an arbitrary assumption, but in view of the approach to one another of the c coefficients at the same eccentricity ratio, this approximation is unlikely to be in error by more than a few per cent. It gives adequate assessment of the trend of virtual inertia ratio with varying LID of the bearing. The maximum and minimum values of virtual inertia ratio thus calculated are plotted in Fig. 21.6. It thus appears that for bearings of the proportions normally used in turbines, the virtual inertia of the oil film is a multiple of the mass of the journal. Whether this Vo1179 Pr 33 at Gazi University on March 2, 2016pcp.sagepub.comDownloaded from 43 JOURNAL BEARING DYNAMIC CHARACTERISTICS-EFFEcr O F INERTIA OF LUBRICANT For journal bearings operating in the viscous flow regime under conditions of Sommerfeld similarity, the displacement and velocity coefficients are constant in the non-dimensional expression of dynamic characteristics; but the acceleration coefficients are not constant; they are proportional to R,(C/D)", + "1 Dimensional Symbols Any consistent system of absolute units is applicable. C D e h* L N n U W 8 81 + 8 2 + Diametral clearance, r = C/2 . J o u e l diameter, R = D/2 . = OA Eccentricity of journal. Film thickness at 8. Bearing length. Speed of rotation, w = 2nN. = e/r Eccenmcity ntio. Peripheral velocity of journal, U = mDN. Steady load applied vertically downwards. Angle from attitude line. Angular extent of pressure bearing film. Attitude angle. Horizontal and vertical axes x, y and attitude axes k, 1 as shown on Fig. 21.1. For local co-ordinate axes see Fig. 21.2. P,*, Py* x*, y* ax, ay FL Viscosity of lubricant. P Density of lubricant. P Pressure of lubricant. Disturbance force superimposed on steady load. Displacement of journal axis from steady-running Acceleration of journal axis. position. Non-dimensional Parameters S Re Sommerfeld duty parameter defined as - Reynolds number defined as 1 Ur = 5- NDC. \u201c P P U Vol 179 P1 3J at Gazi University on March 2, 2016pcp.sagepub.comDownloaded from 44 D. M. SMITH Non-dimensional Symbols APPENDIX 21.11 Sommerfeld (2) solved certain integrals which occur in the analysis of steady-running conditions in circular journal bearings, using the auxiliary angle y defined by INTEGRALS U S E D IN CALCULATION p,* P Y * Disturbance force P, = -9 P - -w y - w x* y* Displacement Y = --,y = - r n p Velocity -J - x a, Acceleration a,' = - = -, alv' = 2 = 3 w a u2r w2 war w w Coefficients (111," + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003196_transele.2018oms0005-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003196_transele.2018oms0005-Figure2-1.png", + "caption": "Fig. 2 Structure of the fuel gel sandwiched by the GCFC electrodes. The upper figure shows the GCFC surface.", + "texts": [ + " The optimum concentration of fuel [8], the optimum modification amount of enzyme of Bilirubin oxidase (BOD) [4], [18]\u2013[21], and the optimum pH [22]\u2013 [25] have been investigated. The optimum concentrations of fuel, gel and BOD were investigated in the proposed system. The mechanism of fuel cell power generation using fructose is shown in Fig. 1. A proton (H+) is generated at the anode by the fructose dehydrogenase (FDH) enzyme. The proton is reduced by dissolved oxygen by the BOD enzyme at the cathode. In this work, an agarose gel mixed with fructose and McIlvain buffer solution (MBS) was sandwiched between the anode and cathode electrodes as shown in Fig. 2. Agarose was dissolved in MBS containing 100 mM fructose at a concentration of 1.5 wt%. Wu et al. reported that pH 5.0 was optimal for the BOD enzyme [25]. Therefore, MBS of pH 5.0 was used. The agarose solution was boiled, poured Copyright c\u00a9 2019 The Institute of Electronics, Information and Communication Engineers into a plastic case, and then cooled and gelated. The anode and cathode were modified by the FDH and BOD enzymes. Table 1 lists the specifications of them. The electrode material was carbon fiber woven fabric modified by graphene on the surface, GCFC (Incubation Alliance Co" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000841_mmar.2012.6347891-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000841_mmar.2012.6347891-Figure3-1.png", + "caption": "Fig. 3. Thrusters Configuration", + "texts": [ + " In general to relate the generalized force vector \ud835\udf0f with the vector of forces applied individually to each thruster, we must consider that the vehicle has \u2019m\u2019 propellers, with m\u22656, for the six possible degrees of freedom. In this paper we will consider a vehicle that is capable of moving along the three dimensions and using their thrusters to provide only 5 degrees of freedom. We\u2019d prefer to not use the roll motion because, it is automatically stabilized by floaters of the vehicle, being unnecessary the spent computation efforts to control it. As shown in Fig. 3 based in work done by [2], we have the distribution of thrusters in the vehicle, that will provide de movement of the vehicle. Therefore we can write. \ud835\udf0f = \ud835\udc35\ud835\udc47. (14) Where \ud835\udf0f1\ud835\udc65\ud835\udc5a is the vector of the generalized forces, \ud835\udc356\ud835\udc65\ud835\udc5a is the matrix of the thruster configuration, that brings precious information of the geometric configuration of the thrusters, \ud835\udc47\ud835\udc5a\ud835\udc651, is the vector of the each thruster force applied, 14, can be rewritten as: ( \ud835\udf0f\ud835\udc3b \ud835\udf0f\ud835\udc49 ) = ( \ud835\udc35\ud835\udc3b \ud835\udc423\ud835\udc652 \ud835\udc422\ud835\udc654 \ud835\udc35\ud835\udc49 )( \ud835\udc47\ud835\udc3b \ud835\udc47\ud835\udc49 ) . (15) Where, \ud835\udf0f\ud835\udc3b = [\ud835\udc39\ud835\udc65, \ud835\udc39\ud835\udc66,\ud835\udc40\ud835\udc67]\ud835\udc47 is the horizontal components of generalized reactions, and \ud835\udf0f\ud835\udc49 = [\ud835\udc39\ud835\udc67,\ud835\udc40\ud835\udc66]\ud835\udc47 vertical components of generalized reactions, \ud835\udc47\ud835\udc3b = [\ud835\udc471, \ud835\udc472, \ud835\udc473, \ud835\udc474]\ud835\udc47 is the horizontal forces applied by each thruster respectively and finally, \ud835\udc47\ud835\udc49 = [\ud835\udc475, \ud835\udc476]\ud835\udc47 the vertical forces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003435_978-3-030-22747-0_49-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003435_978-3-030-22747-0_49-Figure5-1.png", + "caption": "Fig. 5. Schematic of the experimental set-up for AM weld tests.", + "texts": [ + " Three layers of alloy IN-738 were deposited in a SAE-AISI 1524 carbon steel substrate. The effect of convection and radiation heat loss from the layers surfaces were included into the FE analysis. To predict the evolution of temperature distribution in the entire weldment (substrate, two and three cast IN-738LC alloy layers) for the entire welding and cooling cycle of the process, a 3D transient nonlinear heat flow analysis was performed. To observe the heat transference among layers, all welding layers had gluing contacts. Figure 5 shows the experimental set-up, specimen dimensions, and x, y and z directions for AM weld tests. The initial temperature T0 was set to 20 \u00b0C. The designed mesh is shown in Fig. 6a for processes AM1 and AM2, and in Fig. 6b for processes AM3 and AM4. Numerically predicted thermal gradients, isotherms and thermal cycles produced by the AM process were calculated with the developed thermal model. For the first part of this work, the following results corresponding to the processes AM1 and AM2 were obtained in a mesh representing two wire layers (composed of four finite element rows) plus the substrate (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002731_s13198-018-0726-9-Figure13-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002731_s13198-018-0726-9-Figure13-1.png", + "caption": "Fig. 13 Photos of a outer race fault and b inner race fault", + "texts": [ + " It can be seen from the comparison of the two figures that the method proposed in this paper is also effective for another work condition. The dataset obtained from the Mechanical Failures Prevention Group (MFPT) (Fault data sets. Available online: http://www.mfpt.org/FaultData/FaultData.htm (accessed on 10 April 2013) is utilized in this case study. The three health conditions (baseline, inner race fault and outer race fault) are considered for evaluation and discussion in this research. The photos of inner race fault and outer race fault are shown in Fig. 13. The test bearings are radial ball bearings produced by RBC NICE and its parameters can be seen in Table 3. The description of the bearing dataset can be seen in Table 4. In order to test the effectiveness of the GRNN-based method to various operation conditions, all of the dataset are utilized. In addition, every data is divided to several samples which each sample has 36,621 points. Therefore, baseline has 48 samples, outer race fault has 76 samples and inner race fault has 28 samples. The first 10 samples of every health conditions are served as the train samples and the others are served as test samples" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002709_978-1-84996-220-9_5-Figure5.3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002709_978-1-84996-220-9_5-Figure5.3-1.png", + "caption": "Figure 5.3 (a) The humanoid robot ARMAR-III moving in a translational dimension. (b) The effect in workspace when changing the C-space value for the dimension associated with the torso pitch joint", + "texts": [ + " Therefore, an upper bound for the workspace movements of each limb is used for an efficient and approximated uniform sampling. A change \u03b5trans in a translational component of the C-space moves the robot in workspace by \u03b5trans. All other dimensions of the C-space have to be investigated explicitly to derive the upper bound of the robot\u2019s workspace movement. Table 5.1 gives an overview of the maximum displacement of a point on the robot\u2019s surface when changing one unit in C. The effects of moving one unit in the different dimensions can be seen in Figure 5.3. The different workspace effects are considered by using a weighting vector w whose elements are given by the values of the workspace movements from Table 5.1. In Equation 5.1 the maximum workspace movement dWS(c) of a C-space path c = (c0, ...,cn\u22121) is calculated: dWS(c) = n\u22121 \u2211 i=0 wici. (5.1) To sample a C-space path between two configurations c1 and c2, the vector vstep is calculated (Equation 5.2). For a C-space displacement of vstep it is guaranteed that the maximum workspace displacement is 1 mm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000945_ssd.2010.5585533-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000945_ssd.2010.5585533-Figure7-1.png", + "caption": "Figure 7. CPAES stator study domain considered in FEA. Legend: (1) magnetic collector, (2) lamination, (3) yoke.", + "texts": [], + "surrounding_texts": [ + "its DC-excitation winding is located in the stator rather than in the rotor in conventional machines. The proposed model takes into account for the saturation of the mag netic circuit. A special attention is paid to the distribution of the air gap flux density and the back-EMF. A validation of the results yielded by FEA is achieved considering both measurements carried out on a prototype of the studied machine and analytical results obtained by a reluctance model developed in a previous work.\nIndex Terms- Claw pole alternator, DC-excitation in the stator, 3D finite element analysis, flux path, air gap flux density, back-EMF, validation.\n1. INTRODUCTION\nDuring the last four decades, talking about automotive power generation, it is quite commonly dealing with claw pole alternators [1, 2, 3]. The popularity of such electric machine is mainly due to its heteropolar topology offering the possibility of integration of a high pole pair number (6 to 10) in a low volume, leading to high torque densities and interesting generation capabilities.\nHowever, the machine is penalized by a crucial main tenance problem due to brush-ring system through which the rotor field winding is fed by DC current. An approach to discard this drawback and hence increase the availabil ity of the machine has been reported in the literature [4, 5]. Basically, this approach consists in transferring the DC excitation winding from rotor to stator, yielding the so called \"Claw Pole Alternator with DC Excitation in the Stator\" (CPAES). Previous investigations of the CPAES have highlighted interesting features [5, 6]. These have been carried out considering reluctance models. However, accounting for the complicated magnetic phenomena involved in the CPAES, it is expected that the results yielded by reluctance models are more or less accurate. Therefore, finite ele ment analysis (FEA) tum to be a necessity for the sake of a deeper investigation of the CPAES features. The present paper develops this idea.\n978-1-4244-7534-6/10/$26.00 \u00a920 1 0 IEEE\nThe CPEAS concept presents a claw pole topology where the DC-excitation winding is located in the stator rather than in the rotor as in conventional claw pole machines. As a result, the brush-ring system and the associated mainte nance problem have been discarded, which represents cru cial cost and availability benefits.\nThe machine is equipped by a three phase armature winding. In the manner of conventional claw pole alterna tors, the stator is made up of a laminated cylindrical mag netic circuit as shown in figure 1.\nThe field winding is simply wound in a ring shape. Figure 2 shows the photo of one half of the stator field winding.", + "The two halves of the field winding are inserted in both sides of the machine between the armature end-windings and the housing as illustrated in figure 3. They are con nected in series in such a way to produce additive fluxes.\nFollowing the transfer of the DC-excitation winding from rotor to stator, appropriate changes of the magnetic circuit have been introduced. These concerned mainly the rotor where the two iron plates with overlapped claws facing the air gap tum to be magnetically decoupled. Figure 4 shows a photo of the rotor of the CPAES.\nFollowing the removal of the field winding from rotor to stator and for the sake of an efficient flow of the flux, two magnetic collectors have been included in the stator mag netic circuit. These guarantee the flux linkage between rotor and stator. They are embedded on the two flasks of the machine. Figure 5 shows a photo of one magnetic col lector.\nThe CPAES static and rotating components are illustrated in figure 6. One can notice that each claw plate includes six poles. Moreover it is to be noted that the two claw plates are magnetically decoupled.\nIn order to reduce the computation time, the FEA study domain is limited to a one pair of poles of the CPAES. Figures 7 and 8 show the stator and the rotor study domains, respectively.", + "The meshing of both stator and rotor study domains has been automatically generated in the Ansys-Workbench environment.\nFigure 9 shows a mesh of the stator yoke and of the two associated magnetic collectors.\nA mesh of the stator lamination is illustrated in\nfigure 10.\nFigure 11 shows a mesh of the rotor claws and the as sociated magnetic rings.\n3.2.1. Main Flux Paths\nThe flux moves through the magnetic circuit of the CPAES within three dimensional (3D) paths. These are described as follows:\n\u2022 axially in the stator yoke (see figure 12),\n\u2022 radially down through the collector and crossing the\nair gap down to the magnetic ring (see figure 12),\n\u2022 radially then axially in the magnetic ring (see figure 12),\n\u2022 axially-radially in the claws (see figure 13),\n\u2022 radially up in air gap facing the stator laminations\n(see figure 14),\n\u2022 radially up in the stator teeth and circumferentially\nin the stator core back (see figure 15),\n\u2022 radially down in stator teeth facing the two adjacent\nclaws (see figure 16),\n\u2022 radially-axially in the adjacent claws (see figure 16),\n\u2022 axially then radially in the magnetic ring\n(see figure 17),\n\u2022 radially up through the air gap and the mag\nnetic collector on the other side of the machine (see figure 17),\n\u2022 axially in the other side of the stator yoke\n(see figure 18).\n3.2.2. Leakage Flux Paths\nNot all the flux produced by the excitation winding and the armature contributes to the EMF generation. Different leakage fluxes have been distinguished in the CPAES. The two main ones are\n\u2022 the leakage flux linking adjacent claws,\n\u2022 tow dimensional flux paths which flow through the\nmagnetic circuit as follows:\n- axially in the stator yoke,\n- radially down through the magnetic collector and air gap el,\n- radially then axially in the magnetic rings holding the claws,\n- axially in the claw,\n- radially in air gap e2 and in the stator teeth.\nFigure 19 illustrates the flux vectors flowing within 20 paths." + ] + }, + { + "image_filename": "designv11_33_0000162_iemdc.2009.5075177-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000162_iemdc.2009.5075177-Figure4-1.png", + "caption": "Fig. 4. Flux density distribution of the IPM wind generator.", + "texts": [ + " The generated frequency \u03c9s is directly related to the shaft speed of the wind turbine by the following expression 2 2 s m s P f\u03c9 \u03c9 \u03c0= = (7) where P is the number of rotor PM poles of the IPM machine. The IPM wind generator is a salient pole ac machine in which the saliency is created by the orientation of magnets arrangement inside the rotor assembly in stead of built-in variable air gaps of a typical salient pole synchronous machine. Figure 3 shows the schematic of a 4-pole IPM wind generator. The NdBFe magnetic materials are arranged in straight magnet form below the squire cage type dampers. Figure 4 shows the magnet flux distribution of an IPM machine using a straight type arrangement of magnets inside the 4-pole rotor. Figure 5 shows the per phase voltage equivalent circuit diagram of an IPM generator. The steady state per phase voltage E0 at the air gap of a P-pole three phase balanced IPM generator can be expressed as ( )0 0 ( ). p s s q d d q E V I R jX I X X\u03b8= \u2220 + \u2220 \u00b1 + + \u2212 (8) The developed power in a three-phase IPM generator is expressed as 2 0 3 3 ( ) sin sin 2 2 p p d q d d d q V E V X X P X X X \u03b4 \u03b4 \u2212 = + (9) where Vp is the per phase terminal voltage, Is is the per phase current , Rs is the per phase stator resistance, Xd and Xq are the per phase direct and quadrature axis reactances, respectively, and \u03b4 is the torque angle between Vp and E0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002336_iccas.2017.8204221-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002336_iccas.2017.8204221-Figure6-1.png", + "caption": "Fig. 6 Wrist attachment Fig. 7 Birdview of apparatus", + "texts": [ + " These axes are passive ones and correspond to elbow and shoulder position respectively. The axes (iii), (iv) and (v) are actively actuated by the artificial muscle, whereas (iv) and (v) are guides of the same wire. The distance between each axis was determined empirically. The main body is worn on the body by the shoulder belt (a) and the waist belt (b). Also, by attaching the back pad (c), we made sure that the wearer's back face and the McKibben box (M-box) (d) where the artificial muscle is located, do not contact with each other when worn. As shown in Fig. 6, the wrist is fixed by using a stretchy supporter. Several persons' shoulder widths were measured and the shoulder width of the device was decided as 500 mm in order to secure a sufficient shoulder width. Note that, since the waist width of the muscle suit for lower back assistance is 388 mm, as shown in Fig. 7, by tilting M-box 35 \u00b0, a waist width of 388 mm and a shoulder width of 500 mm were realized. The weight is 6.5kg and it is possible to wear it in about 30 seconds. Integral electromyography (IEMG) which shows total muscle power used during experiments, is applied to evaluate the muscle suit\u2019s effectiveness" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002701_23311916.2018.1493672-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002701_23311916.2018.1493672-Figure1-1.png", + "caption": "Figure 1. Schematic representation of four-pad adjustable bearing (Source: United States Patent and Trademark Office, www.uspto.gov, U.S. Patent No. 5,772,334, 1998): (1) bearing casing, (2) adjustable pad, (3) tilting spacer, and (4) radial displacement spacer.", + "texts": [ + " The operating principles were further extended onto partial arc bearings with adjustable geometry, where bearing segments are actuated by employing piezoelectric actuators. The proposed bearing design was implemented on industrial turbines, where enhanced stability margins were observed at high rotational speeds (Chasalevris & Dohnal, 2016). The objective of this study is tomathematically formulate the film thickness in amultipad externally adjustable fluid film bearing. Four adjustable bearing segments circumferentially arranged with a pad angle of 48\u00b0 as shown in Figure 1 is considered for the analysis. These bearing segments are configured to have both radial and tilt adjustments as in Figure 2. In this paper the modified film thickness equation is derived theoretically incorporating symmetrical pad adjustments. The film thickness equation developed can be applicable to the multiple numbers of pads present in the bearing and provides an accurate approximation of the film thickness gradient in the bearing clearances. The present study can be further extended in determining the theoretical performance characteristics of a four-pad externally adjustable bearing by employing the modified film thickness equation defined in this paper", + " For an alternate form of adjustable bearing with a single adjustable pad, Shenoy (2008) determined the variation of fluid film thickness by superimposing the radial and tilt displacements of the bearing element on a conventional film thickness equation. This principle was adopted based on the superimposition of ramp depth of a tri taper bearing as described in Hargreaves (2016). Dimensionless film thickness equation developed by Shenoy (2008) by considering a straight-line approximation of the pad angle under different adjustments is given as follows: In the present study, a four-pad externally adjustable bearing configuration with L/D ratio of 0.53 is considered, as shown in Figure 1. The bearing geometry consists of four bearing pads with an angle of 48\u00b0 circumferentially spaced and capable of translating in radial and tilt directions. These pad adjustments act as an effective means to control the bearing radial clearance and circumferential film thickness gradient. Figure 2 illustrates the bearing pad adjustment configuration, where symmetrical adjustments will be provided to the remaining pads under different operating conditions. Under negative radial adjustment, the bearing pads are displaced to move inward and thereby reducing the radial clearance spaces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003813_s12206-019-1039-x-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003813_s12206-019-1039-x-Figure9-1.png", + "caption": "Fig. 9. Resultant force acting on the Segway (in case of left or counterclockwise rotation).", + "texts": [ + " The centripetal force acts toward the centre of the circle when the Segway runs in a circular motion, on the other hand, the centrifugal force is generated by the inertial force of the driver and it acts outwards from the centre of the circle. For the stable turning motion of the Segway, the centripetal force and the centrifugal force must balance. If the footplate of the Segway is lifted to the direction of the centrifugal force, the driver\u2019s center of gravity moves toward the centre of circle and the Segway can stably run along with the circular path as shown in Fig. 9. The centrifugal force can be calculated using the radius of rotation, R, and the velocity of the Segway as follows: 2 2 c c vF ma m r R w= = = r rr , (14) In other words, the value of how much the footplate of Segway has to be lifted can be calculated using the centrifugal force acting on the driver, and the angle, Rq , which moves the driver\u2019s center of gravity to the centre of the circle can also be calculated. If the Segway is moving at a fixed speed, v r , the force acting on the Segway is shown in Fig. 9. In Fig. 9, the normal force n r can be expressed as the sum of the vertical and horizontal vectors, and the resultant forces acting along the vertical axis are zero. That is, the vertical component of n r is equal and acts the opposite direction to the force of gravity as follows: cos 0Rn mgq - = r r . (15) Also, in Fig. 9, the horizontal vector of normal force acts toward the centre of the circular path and is equal to the centripetal force, Fc, and can be expressed as follows: sinc RF n q= r . (16) When the Segway is moving at a fixed speed, v, in a curved trajectory, the angle, Rq , which how much the footplate of the Segway has to be lifted can be calculated as: 2 arctan( )R v Rg q = r r . (17) which can be obtained by putting Eqs. (14) and (15) into Eq. (16). In this study, the SEA was designed to be used for the active suspension system and the Segway for experiments was self-manufactured as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002804_ls.1430-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002804_ls.1430-Figure1-1.png", + "caption": "FIGURE 1 First\u2010generation gas foil bearing cross\u2010section schematic, viewed at wileyonlinelibrary.com]", + "texts": [ + " However, when using the Newton\u2010Raphson method, the inverse lubrication problem needs to be solved instead of the direct problem; ie, the applied load is known but the corresponding Cartesian eccentricity components of journal need to be calculated iteratively. 2 | THEORETICAL ANALYSIS 2.1 | Modified Reynolds equation and its analytical perturbation process For inverse aerodynamic lubrication problem (ie when the static external load is imposed), the motion of the rotor in a compliant self\u2010acting gas film bearing is described in the frame (X, Y) related to the applied load W where the origin is located at the bearing centre as illustrated in Figure 1. The gas film pressure can be integrated over the surface of the bearing to obtain the lift force components: FX FY \u00bc Z L=2 \u2212L=2 Z 2\u03c0 0 p cos\u03b8 sin\u03b8 Rd\u03b8dz; (1) where p is the gas film pressure, R is the shaft radius, L is the bearing length, and \u03b8 is the circumferential coordinate so\u2010called the bearing angle originating at the X\u2010axis as indicated in Figure 1. The film pressure is obtained from the solution of the modified Reynolds' equation derived from modified and gas\u2010film and bump foil structure models [Colour figure can be Navier\u2010Stokes and continuity equations for a nonNewtonian couple\u2010stress fluid by using the V. K. Stokes micro\u2010continuum theory19: \u2207\u00b7 \u03c1Q\u00f0 \u00de \u00bc \u2212 \u2202 \u2202t \u03c1h\u00f0 \u00de; (2) where \u03c1 is the fluid mass density, h is the film thickness, t is the time variable, and Q \u00bc Qx Qz \u00bc \u2212 G h; l\u00f0 \u00de 12\u03bc \u2202p \u2202x \u2202p \u2202z 8><>: 9>=>;\u00fe h 2 \u03c9R 0 8<: 9=; (3) Qx and Qz being the volume flow rate components per unit length in x and z directions respectively", + " For a compliant cylindrical bearing, the film thickness h is defined by the following expression: h \u03b8; z; t\u00f0 \u00de \u00bc C \u00fe X t\u00f0 \u00de cos\u03b8\u00fe Y t\u00f0 \u00de sin\u03b8\u00fe Lp \u03b8; z; t\u00f0 \u00de; (8) where C is the bearing radial clearance, X and Y are the Cartesian coordinates of the shaft centre, and L \u00bc 2s E \u2113 tb 3 1\u2212\u03c32\u00f0 \u00de is the scalar compliance operator of the bump foil in (m Pa\u22121) according to the Heshmat's analytical model31,32 by modelling the corrugated sub\u2010foil as a simple Winkler elastic foundation with isotropic stiffness Kr as shown in Figure 1. The boundary conditions associated to the Reynolds Equation 7 may be classified as follows: \u2022 Boundary conditions related to the environment in which the system operates: p \u03b8; z \u00bc \u00b1 L 2 ; t \u00bc pa; at the bearing edges (9a) \u2022 Periodicity condition: p \u03b8 \u00bc 0; z; t\u00f0 \u00de \u00bc p \u03b8 \u00bc 2\u03c0; z; t\u00f0 \u00de (9b) \u2022 Boundary conditions related to lubricant flow (depression phenomenon): p \u03b8sub; z; t\u00f0 \u00de \u00bc pa \u2202p \u2202\u03b8 \u03b8sub; z; t\u00f0 \u00de \u00bc \u2202p \u2202z \u03b8sub; z; t\u00f0 \u00de \u00bc 0 8<: : (9c) At these conditions, we can add for aligned journal bearings the following condition: \u2202p \u2202z \u03b8; z \u00bc 0; t\u00f0 \u00de \u00bc 0; at the bearing centreline: (9d) Equation 9a results from the fact that the ends of the bearing are exposed to ambient pressure, while Equations 9c are the Reynolds (Swift\u2010Stieber) conditions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002099_amm.868.124-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002099_amm.868.124-Figure2-1.png", + "caption": "Fig. 2 Finite element model", + "texts": [ + " To simulate the distribution of the temperature field and variation of thermal deformation, the following assumptions are made: (1) Grooves of the nut and screw shaft are ignored; (2) The convection heat transfer coefficient is a constant value; (3) The heat conduction between solid structure and lubricant is negligible; (4) The heat generation of the motor cannot conduct into the ball screw system. Moreover, small structure of the ball screw system should be reasonably simplified, such as chamfer, thread hole, fillet and key groove. The tetrahedral and hexahedral element with 20 nodes was applied to mesh the solid structure, as shown in Fig. 2. There are 242,149 nodes and 134,455 elements in the FEA model. Heat generation of ball screw According to Refs. [16] and [17], the heat generation of ball screw can be expressed as 2 60 Q nM (1) where Q denotes the heat generation of ball screws; n denotes the ball screw\u2019s rotational speed; and M denotes the friction torque of ball screws and can be expressed as 2 ( )cos M z Me Mg (2) where z denotes the number of balls, denotes the helix angle of the raceway, and Me and Mg denote the friction torque caused by the external load and the friction torque 4 3 4 Q Me m (3) 2 5 3 2 16 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000451_978-3-642-29329-0_6-Figure6.4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000451_978-3-642-29329-0_6-Figure6.4-1.png", + "caption": "Fig. 6.4 Representation of the single-track model", + "texts": [ + " Additionally we will add the forces at the left and right side of an axle, obtaining Fx1 \u00bc Fx11 \u00fe Fx12 ; Fx2 \u00bc Fx21 \u00fe Fx22 ; Fy1 \u00bc Fy11 \u00fe Fy12 ; Fy2 \u00bc Fy21 \u00fe Fy22 ; Moreover, assuming that the steering angle d is small (typical values are between 15 and +15 degrees) we can set cos\u00f0d\u00de 1 e sin\u00f0d\u00de d. Thus m\u00f0 _u vr\u00de \u00bc Fx1 \u00fe Fx2 Fy1d 1 2 rSCxu 2; m\u00f0 _v\u00fe ur\u00de \u00bc Fx1d\u00fe Fy1 \u00fe Fy2 ; Iz _r \u00bc Fx1da\u00fe Fy1a Fy2b By means of the above described approximations we have obtained the so called single-track model, in which the vehicle is actually reduced to a rigid pole with two tires. It is worth noticing that the same equations can actually be obtained by starting directly from this assumption, namely by using the mechanical model of Fig. 6.4. The angles a1 and a2 in Fig. 6.4 are the lateral slip angles and play a fundamental role in the vehicle dynamics, as it will be detailed in Sect. 6.2.5. The centre of rotation of the vehicle model is denoted by C in Fig. 6.4. In the same figure the radius r of the trajectory of G is reported. In order to obtain the simplest possible set of equations for the subsequent bifurcation analysis, we now introduce the following additional assumptions: \u2022 The forward speed u is constant. \u2022 The drag forces (rolling resistances and aerodynamic resistances) are vanishing. \u2022 No longitudinal forces are acting at the wheels. The first assumption implies that the first balance equation actually reduces to an algebraic one, so that the mathematical model of the vehicle has only two differential equations, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000758_s11465-012-0317-4-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000758_s11465-012-0317-4-Figure6-1.png", + "caption": "Fig. 6 Contour atlas of norm\u00f0\u03c9\u00de when \u03b23 \u00bc 72\u2218", + "texts": [], + "surrounding_texts": [ + "As we known that the SNU 3-UPU parallel mechanism prototype will always exhibit the parasitic rotations in practice [17]. Now we will use the value of norm\u00f0\u03c9\u00de to find the bound of mechanism instability, which indicates there, will appear the parasitic rotation when the number exceeds the bound. For SNU 3-UPU parallel mechanism, the first revolute axes of universal joint attached to the fixed base intersecting at one common point. And the same is true for the last revolute axes of universal joint attached to the moving platform, as shown in Fig. 7(a). Take one limb as an example, the direction of the constraint couple applied on the moving platform is sketched in Fig. 7(b). It\u2019s easy to find that the angle \u03b2i from the direction of constraint couple to the moving platform keeps in a relative stable area. It is around \u03c0=2, and we express the angle \u03b2i as \u03b2i \u00bc \u00bd\u03c0=2 . Of course, the mechanism will be singularity when all the three constraint couples are parallel with each other. However, since the existing clearances, these three constraint couples will never truly be parallel with each other in practice. Since the value \u03b2i \u00bc \u00bd\u03c0=2 , we choose a values set of \u03b23, (80\u03c0=180, 85\u03c0=180, 87\u03c0=180, 89\u03c0=180), to get the contour atlas of norm\u00f0\u03c9\u00de, as shown in Fig. 8. Through observation of the above contour atlas of norm\u00f0\u03c9\u00de, and the obtained practical region of angle \u03b2i, \u03b2i \u00bc \u00bd\u03c0=2 , we conclude that the norm of parasitic rotation, norm\u00f0\u03c9\u00de, of SNU parallel mechanism is located in the area of around value 1.2. However, practically the manufactured hardware prototype of SNU parallel mechanism really revealed an unexpected parasitic rotation due to the joint clearance. When all the prismatic joints are locked, the mechanism behaves as if it has additional rotational degree of freedom. Hence we call the number 1.2 is the bound of instability. The parallel mechanism combined with three UPU limbs will be instability when the norm of parasitic rotation exceeds 1.2 with limited clearance \u03b5. Namely, if we still want to achieve the pure translation without parasitic rotation through the parallel mechanism with three UPU limbs, we should study the following two aspects further. One is obtaining three relative stable parameters \u03b2i by placing the universal joint at the proper position and orientation. These three parameters can make the value of norm\u00f0\u03c9\u00de be lower. Another is decreasing the possible clearance to the greatest extent." + ] + }, + { + "image_filename": "designv11_33_0003893_0954406219888240-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003893_0954406219888240-Figure3-1.png", + "caption": "Figure 3. Structure diagram of the universal joint.", + "texts": [ + " Thus, the workspace of the universal joint is important for the combined spherical joint, which will be discussed first. Although the workspace analysis of this class of universal joints has been discussed,11 they are too complex to use. Inspired by the cross-sectional area method, a concise projection method is presented to study the joint workspace in this paper, so that only two parameters can determine the non-interference domain of this class universal joint. Before the workspace analysis of the commonly used universal joint, some notations are firstly presented as shown in Figure 3. Clearly, the universal joint is composed of a bottom hinge, an upper hinge, and a cross trunnion that is made up of a block and two pins. For the convenience of description, two coordinate systems are established. The o-UVW coordinate system is fixed on the bottom hinge, and the o-uvw coordinate system is moving with the upper hinge. The V-axis is defined in the direction from the cross trunnion center to the Pin 1, and the W-axis is along the axis of bottom hinge. Simultaneously, the u-axis is defined in the direction from cross shaft center to Pin 2, and the w-axis is along the axis of upper hinge" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002343_icrera.2017.8191184-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002343_icrera.2017.8191184-Figure2-1.png", + "caption": "Fig. 2. SRM with a central-tapped node. (a)Windings with central-tapped node. (b)Motor drive topology", + "texts": [ + " The topology based on fullbridge converter with driving and charging functions. Comparing with conventional full-bridge converter, the topology include an additional relay. The simulation and experiments results confirm the proper performance of the converter. II. INTEGRATED TOPOLOGY WITH CENTRAL-TAPPED Conventionally, the phase wingdings of SRM consist of an even number of series connected windings [12], as shown in Fig. 978-1-5386-2095-3/17/$31.00 \u00a92017 IEEE 1. Thus, central-tapped windings are formed, which can be easily developed in SRM as shown in Fig. 2(a). Fig. 2(a) and Fig. 2(b) show the central-tapped node of the winding and the full-bridge converter. The investigated integrated topology is shown in Fig. 3. A 6/4 poles SRM has phase A, phase B and phase C. La1, La2, Lb1 and Lb2 are winding inductances of phase A and phase B. NA and Nb are winding central-tapped nodes of phase A and phase B. S0~ S11 are twelve MOSFETs, D0~ D11 are twelve parasitic diodes paralleled with MOSFETs. Compared with traditional full-bridge converter, a relay is added between switch S8 and switch S10 to divide the voltage stage of dc-link and battery" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002873_978-3-030-00232-9_56-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002873_978-3-030-00232-9_56-Figure2-1.png", + "caption": "Fig. 2. Experimental test rig.", + "texts": [ + " To avoid errors and reduce noise, provided that the scenario viewed by the camera is steady, each distance in the Depth image has been evaluated as the mean value on 20 consecutively acquired distance values. The noise of the acquired data can provide inaccurate depth results if a single frame is considered [14]. In the hypothesis that scene observed by the Kinect, is stationary, it is possible to reduce the measurement errors, evaluating each distance in the Depth image as the mean value on more data; in this application 20 consecutively frames are used. The experimental setup, Fig. 2, comprises of an underactuated mechanical finger that is constrained to a rigid support by the proximal phalanx end part. The finger is linked to a pulley by means of two inextensible wires that represents the traction and antagonist tendons [3, 5]. An analog servomotor moves the pulley and sets the finger tendon displacements. A myRIO Embedded Device controller (National Instruments) controls the servomotor rotation. Moreover, the experimental setup is equipped with an encoder to receive a response on the motor angular position" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003909_icems.2019.8922564-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003909_icems.2019.8922564-Figure1-1.png", + "caption": "Fig. 1. The schematic diagram and definition of coordinate system", + "texts": [ + " Finally, comprehensive performances including the estimated angle errors, torque-per-ampere and load capacity are respectively evaluated. Importantly, as a convincing reference, an experimental comparison with a commercial ABB standard driver is carried out. In this section, the model of the rotor position and speed observer with MRAS , EEMF-based and direct stator flux calculation methods are derived respectively. The schematic diagram and definition of coordinate system of the SynRM involved in this paper is described in Figure 1. 978-1-7281-3398-0/19/$31.00 \u00a92019 IEEE The current equations of Permanent Magnetic Assisted Synchronous Reluctance Machine (PMaSynRM) in d-q coordinate system is as follows: 1 0 1 0 qsd e d d dd d q q q e fd s e qq q LRdi i L uL Ldt di i uL R LL Ldt \u2212 = + \u2212 \u2212 \u2212 (1) For the SynRM, the voltage equations are the same except that the permanent magnet flux linkage \u03c8f =0. The MRAS adjustable model can be expressed as: \u02c6\u02c6 1\u02c6 \u02c6 0 \u02c6\u02c6\u02c6 \u02c6 \u02c6 \u02c6\u02c6 \u02c6 1 0\u02c6 \u02c6\u02c6 \u02c6 qs sd e sd d sdd d sqsq sqd s e q q q LRdi i L uL Ldt udi iL R Ldt L L \u2212 = + \u2212 \u2212 (2) Where \u02c6 \u02c6 \u02c6\u02c6 \u02c6 \u02c6 \u02c6, , , \u02c6 \u02c6 f f sd d sq q sd d s sq q d d i i i i u u R u u L L = + = = + = (3) The adaptive law of can be derived as follows: ( ) ( ) ( ) \u02c6 \u02c6 \u02c6 \u02c6 \u02c6 \u02c6\u02c6 \u02c6 0 \u02c6 \u02c6 qi d e p sd sd sq sq sq sd e d q LK L K i i i i i i s L L = + \u2212 \u2212 \u2212 + (4) The rotor position can be directly achieved through integral operation as: \u02c6 \u02c6 e edt = (5) The voltage equations of PMaSynRM in d-q coordinate system is as follows: ( ) d d s d d e q q q q s q q e d d f di u R i L L i dt di u R i L L i dt = + \u2212 = + + + (6) The relationship between rotating d-q coordinate system and stationary \u03b1-\u03b2 coordinate system can be expressed as follows: cos sin cos sin , sin cos sin cos d d q q i ui u i ui u = = \u2212 \u2212 (7) sin cos cos sin cos sin sin cos d e q di di idt dt di i di dtdt \u2212 = + \u2212 \u2212 \u2212 (8) from the following expression: ( ) ( ) sin tan cos s q e d q e e e s q e d q di u R i L L L i dt di u R i L L L i dt \u2212 \u2212 \u2212 \u2212 = = \u2212 \u2212 + \u2212 (9) The extended electro-motive force in \u03b1-\u03b2 coordinate system can be expressed as: ( ) ( ) s q e d q s q e d q di e u R i L L L i dt di e u R i L L L i dt = \u2212 \u2212 + \u2212 = \u2212 \u2212 \u2212 \u2212 (10) The rotor position angle and angular speed can be directly calculated through tane e arc e = (11) e e d dt = (12) The flux linkage in \u03b1-\u03b2 coordinate system can be expressed as follows: ( ) ( ) s s s s u R i dt u R i dt = \u2212 = \u2212 (13) The rotor position angle and angular speed can be directly calculated through tan s e s arc = (14) e e d dt = (15) In this section, the comprehensive performances including the estimated angle error, torque-per-ampere and load capacity are respectively evaluated and compared" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001761_15325008.2012.700384-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001761_15325008.2012.700384-Figure2-1.png", + "caption": "Figure 2. Test system configuration: (a) SMIB system and (b) SMSL system.", + "texts": [ + " : (24) With the existing third-order model ignoring the stator transient and rotor speed, both efDQ and Te are not directly related to ! and are not valid when the rotor speed change D ow nl oa de d by [ G eo rg e M as on U ni ve rs ity ] at 2 2: 35 0 7 Ja nu ar y 20 15 is included. The proposed reduced model considering the rotor speed neither introduces new state variables nor requires a small step size, and it is easy for actual application. Two system configurations are tested as shown in Figure 2. The first is a single-machine infinite-bus (SMIB) system, where the SG connects an infinite system. The second is a single-machine single-load (SMSL) system, where the SG connects a load given by constant impedance. They correspond to the extreme system conditions, i.e., with or without synchronism capability and requirement. The SG parameters are SGN D 555 MVA, cos'N D 0:9, Ra D 0:003 p.u., Ld D 1:81 p.u., Lq D 1:76 p.u., Ll D 0:15 p.u., L00 d D 0:23 p.u., L00 q D 0:25 p.u., T 0 do D 8:0 sec, T 00 do D 0:03 sec, T 00 qo D 0:07 sec, TJ D 7:05 sec, and D D 2:0 p" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003375_ilt-10-2018-0390-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003375_ilt-10-2018-0390-Figure4-1.png", + "caption": "Figure 4 The force diagram on the rolling element", + "texts": [ + " When the bearing temperature increases, the expansion difference d r of inner ring and outer ring in the radial direction is: d r \u00bc 0:5as diDTi doDTo\u00f0 \u00de (7) The thermal expansion d b of the ball can be expressed as: Figure 1 Schematic diagram of bearing under initial preload and assembly stress diagram of bearing under initial preload and assembly stress 2 2 d a\u03b4 2 d 2 id oC 0\u03b1 1iC 1oC iC 0 aF I 1iC iC 0\u03b1 1\u03b1 2A iu 1A \u03b1 oC a\u03b4 1oC iu Figure 2 Thermal deformations of the bearing od ox \u03b1 id ix Angle contact ball bearing Pingping He, Feng Gao, Yan Li, WenwuWu and Dongya Zhang Industrial Lubrication and Tribology D ow nl oa de d by B os to n C ol le ge A t 2 0: 42 1 6 M ay 2 01 9 (P T ) d b \u00bc abdbDTb (8) It is assumed that the curvature center position of bearing outer raceway is unchanged, the center position of the ball and the displacement of the curvature center of the inner raceway is further changed because of the centrifugal and thermal effect as shown in Figure 3. The geometric equation of the raceway curvature center and the center of the ball is: A1j \u00bc loj sinaoj 1 lijsinaij A2j \u00bc lojcosaoj 1 lijcosaij ( (9) Here, lij \u00bc fi 0:5\u00f0 \u00dedb 1 d ij, loj \u00bc fo 0:5\u00f0 \u00dedb 1 d oj. Under the combined action of centrifugal load, thermal load and preload, the geometric relationship between the curvature center of the raceway and the center of the ball is: A1j \u00bc Asina1 d a A2j \u00bc Acosa1 d r 1 d c ( (10) The force of analysis the ball is shown in Figure 4. The equilibrium equation for each ball j (j = 1, . . ., Z) can be established as follows: Qijcosaij Qojcosaoj Tijsinaij 1Tojsinaoj 1Fcj \u00bc 0 Qijsinaij Qojsinaoj 1Tijcosaij Tojcosaoj \u00bc 0 ( (11) Above parameters are explained in detail (Zhang et al., 2017a, 2017b). Under the fix-position preload, the actual axial load of the bearing can be written as: Fa \u00bc Xz j\u00bc1 Qojsinaoj\u00f0 \u00de (12) The heat generated in the bearing mainly derived from the friction between the ball and the raceway. For grease lubricated angular contact ball bearing, the frictional heat mainly includes the heat generated by the differential sliding of the contact area between the ball and the raceway, the heat by spin sliding of the ball, the heat by the lubrication drag friction, the heat by the gyroscopic motion, and the heat by sliding between the ball and the cage" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002319_i2017-11597-1-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002319_i2017-11597-1-Figure1-1.png", + "caption": "Fig. 1. The wire trap experimental setup: a 50 \u03bcm current carrying copper wire suspended over a glass substrate traps superparamagnetic beads. Additional pairs Helmholtz coils induce uniform fields in the x, y and z direction that modulate the field of the wire.", + "texts": [ + " Susceptibilities are defined as a response function of the magnetization M or the magnetic moment m = MVb to the external magnetic field m = \u03c7\u0303bBe = \u03c7bVbHe and M = \u03c7bHe. In the following we omit the index in Be and He. Bead trapping on the wire. The colloidal suspension is placed inside a silicone chamber (Coverwell perfusion chamber, height 1mm) that was glued to a microscope glass slide. Due to their large density the colloidal beads sediment on the microscope glass slide. An uninsulated copper wire (diameter dw(\u2261 2rw) = 50\u03bcm connected to a constant current generator passes through the experimental cell along the z-direction, as sketched in fig. 1. Spacers of height \u2248 100\u03bcm ensured that the wire is suspended above the microscope slide surface, and its height h was measured before each experiment. In order to generate magnetic fields able to attract particles at the surface of the wire, significant currents must go through the wire, generating high current densities up to 200\u2013300A/mm2, giving rise to fields of several mT at the wire surface). The thin wire could support currents up to 0.5A for several minutes and up to 0.8A for short times without deterioration. Additional fields generated by a set of commercial and custom made Helmholtz coils along the x, y or the z direction were superimposed with the field of the wire (see fig. 1). The coils were driven by a computer controlled signal generator linked to a custom amplifier that allows to choose the amplitudes and phase differences of the magnetic fields in the frequency range of 0.1 to 20Hz. Observation of the beads was performed on an inverted microscope, Eclipse Ti-S (Nikon), equipped with a PlanFluor 10\u00d7/0.30 objective. Images were time lapse recorded with a CCD camera (Hamamatsu) and subject to image analysis in NIH-ImageJ \u2014a Java-based, open-source software package", + " In the simplest case of a straight long wire the induced field is HI(x) = (I/2\u03c0r)e\u03c6 with r = |x| the radial distance to the wire center, \u03c6 the azimuthal angle and e\u03c6 = ey sin \u03c6\u2212 ex cos \u03c6 the azimuthal unit vector. In general, a spherical, paramagnetic bead placed in an external magnetic field H(x) has a magnetic energy2 given by W (x) = \u2212\u03bc0 2 m(x)H(x) = \u2212\u03bc0 2 \u03c7bVbH2(x), (1) where m(x) \u2261 \u03c7bVbH(x) the magnetic moment of the bead, \u03c7b is the bead\u2019s susceptibility and Vb its volume. The bead experiences a magnetic gradient force F = \u2212\u2202W/\u2202x. We have built a simple setup consisting of a 50\u03bcm (radius 25\u03bcm) thick long slender copper wire suspended above a glass surface, see fig. 1. After switching-on the current through the wire, superparamagnetic beads are attracted to the wire with a force acting radially to the wire (in er direction) given by F = \u2212\u03bc0\u03c7bVb 4\u03c02 I2 r3 er. (2) 2 The energy of a magnetizable bead in an external field is in fact a free energy F = E \u2212 TS and contains contributions from the magnetic energy and the entropy \u2212TS of magnetic monodomains. For simplicity we refer to it here as (magnetic) energy. If the wire is sufficiently close to the substrate plane (typically < 100\u2013150\u03bcm) and the current is high enough (I > 100mA) the beads that have reached a close proximity to the wire begin to lift off the glass surface and attach to the bottom of the wire" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000239_sii.2010.5708334-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000239_sii.2010.5708334-Figure1-1.png", + "caption": "Fig. 1. Take off and landing of tail sitter VTOL", + "texts": [ + " There are several ways to perform VTOL maneuver such as tilt-rotor, tilt-wing, thrust-vectoring, and tail-sitting etc. The simplest way is tail-sitting since it needs no extra actuators the VTOL maneuver unlike the tiltrotor, tilt-wing and thrust-vectoring. A simple mechanism is preferable for aerial robots, because weight saving is crucial for the VTOL maneuver. Tail-sitter VTOL aircraft switches between level flight and hover modes by changing its pitch angle of the fuselage by 90 [\u25e6] as shown in Fig. 1. Recently, different types of tail-sitter VTOL UAVs have been developed. Green et al. developed a simple tail-sitter VTOL UAV which airframe is single propeller R/C airplane [1]. Stone has developed the T-Wing, which has a canard wing and tandem rotors [2]. US Air Force Research Lab and AeroVironment Inc. developed SkyTote which is equipped with a coaxial contra-rotating propeller [3]. These tail-sitter VTOL UAVs were developed based on fixed-wing aircraft. They control their attitude by use of control surfaces (rudder, aileron and elevator)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003541_iemdc.2019.8785161-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003541_iemdc.2019.8785161-Figure8-1.png", + "caption": "Fig. 8 Inner rotor structure of the magnetic-geared motors.", + "texts": [ + " On the other hand, deterioration of torque characteristics is concerned since saturation magnetic flux density is lower than the conventional Si steel. Fig. 6 shows the comparison of iron losses in each core material. The iron losses is reduced about half by using the 6.5 % Si steel as shown in the figure. Next, Fig. 7 indicates the comparison of torque characteristics. The figure reveals that the decrease in torque is slight by employing the 6.5 % Si steel. Next, a method for reducing eddy current loss of inner rotor magnets is described. Fig. 8 shows two structures of the inner rotor. One is a surface permanent magnet (SPM) type shown in Fig. 8(a), the other is an interior permanent magnet (IPM) type shown in (b), respectively. The SPM type is the most conventional type and employed in the previous magnetic-geared motor. However, the eddy current loss is large since the magnets facing to the air gap are affected by the slot harmonics. Thus, the IPM type that the magnets are embedded in the inner rotor core is employed for the improved machine. Fig. 9 shows the comparison of eddy current loss when a current density is 5 A/mm2. It is clear that the eddy current loss of the IPM type is reduced to 1/4 in comparison with the SPM type" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002949_detc2018-85642-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002949_detc2018-85642-Figure1-1.png", + "caption": "Figure 1 Experimented Part (a: Printed Part; b: Part CAD Model; c: Part Point Cloud)", + "texts": [ + " The developed model is then used to optimize the energy consumption per part given a specific design provided the constraints based on final part quality requirements. The study addresses the quantification of AM energy consumption with part geometric accuracy consideration. Three process parameters are included as control variables and AM energy consumption is the response. Each of these parts was then sent to 3D scanner to generate point cloud for geometric accuracy evaluation. Table 1 presents the process parameters and their corresponding levels. The part shown in Figure 1 is used for experiments, and the design is similar with the one used in [27]. Using full factorial design, 27 cases with 2 replications for each case are developed and conducted by MakerGear M2e FDM 3D printer (Fig. 2a). The HOBO UX 120 Plug Load logger (Fig. 2 d) is used to collect the energy consumption data during the AM process. This device was used as a data logger to record the instant energy consumption data with sampling frequency of 1 Hz. The energy consumption data were collected through HOBO data logger software (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001656_s0219843612500156-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001656_s0219843612500156-Figure3-1.png", + "caption": "Fig. 3. ZMP (a) stable; (b) unstable.", + "texts": [ + " Ground reference point stability methods Ground reference point methods are the most widely applied methods in DHM.21 23,29,30 These ground reference points include the ground projected center of mass,39 the zero moment point (ZMP),33 and the foot rotation indicator point (FRI).31 The most famous ground reference point method is the ZMP stability criterion. It states that for a posture to be stable, the location of the ZMPmust fall within the convex contact region created by the body-ground interface, as shown in Fig. 3. By classical de\u00afnition, the zero-moment point is a point on the ground in which the tipping moments, or the moments about the ground parallel axes, are zero. In other words, it is the point on the ground where the net reaction forces act, which implies its coincidence with the center of pressure of the contact area. Also, in cases where no external loads are present on the body and postures are static, it is the ground projection of the whole body center of mass.31,32 The advantages of the ZMP are two-fold" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002597_978-981-13-0107-0_22-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002597_978-981-13-0107-0_22-Figure3-1.png", + "caption": "Fig. 3 Schematic diagram of orthogonal scanning strategy and overview of specimens", + "texts": [ + ", LTD, which equips with a fiber laser with maximum power output of 500 W, maximum scanning speed of 7 m/s. Before SLM experiment, the base plate was pre-heated to approx. 373 K to reduce thermal gradient and stress. During the process, the build chamber was filled with argon gas and the oxygen content is below 100 ppm. The as-cast specimen, which is used to study the difference from SLMed condition, is prepared by vacuum arc furnace. The specimen is cylinder with diameter of 10 mm and height of 50 mm. In this study, orthogonal scanning strategy is chosen as shown in Fig. 3. The overview of specimens is also revealed in Fig. 3. Cylindrical specimens with diameter of 10 mm and height of 15 mm are produced. Scanning vector changed 90\u00b0 every two layers. The influence of process parameters including velocity v, laser power p and hatching space h on the properties of parts manufactured by SLM are investigated. The other parameters are constant. The layer thickness t is fixed at 40 lm in order to obtain accurate dimension parts. The energy density(E), which represents a measure for evaluating the average energy transmitted to per area of single-track, is calculated by Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000400_978-90-481-8764-5_3-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000400_978-90-481-8764-5_3-Figure4-1.png", + "caption": "Fig. 4 The principle of the stabilizer bar. Due to the high inertia the stabilizer bar lags behind the roll/pitch movement, applies a cyclic pitch input to the rotor and creates a redress moment", + "texts": [ + " 2 G = A BJ \u00b7 \u23a1 \u23a3 0 0 mg \u23a4 \u23a6 = mg \u00b7 \u23a1 \u23a3 \u2212 sin \u03b8 cos \u03b8 sin\u03c6 cos \u03b8 cos\u03c6 \u23a4 \u23a6 (13) With all forces and moments defined, the next step is to model the dynamics of the stabilizer bar, swash plate and the electro motors. 3.3 Stabilizer Bar and Swash Plate An important part of the system model is the stabilizer bar. In simple words this stabilization mechanism gives cyclic pitch inputs, similar to the swash plate, to the upper rotor to stabilize the helicopter in flight. The stabilizer bar has a high inertia and lags behind a roll or pitch movement of the fuselage as shown in Fig. 4. Through a rigid connection to the rotor, this time delay results in a cycling pitching of the rotor blades and therefore to a tilting of the TPP. If the stabilizer bar is adjusted correctly the thrust vector shows in the opposite direction of the roll or pitch movement causing a redress moment. Reprinted from the journal 33 The stabilizer bar following the roll/pitch movement can be modeled as a first order element as \u03b7\u0307bar = 1 Tf,up (\u03c6 \u2212 \u03b7bar) , \u03b6\u0307bar = 1 Tf,up (\u03b8 \u2212 \u03b6bar) , (14) with angles \u03b7bar and \u03b6bar" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000303_ichr.2010.5686287-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000303_ichr.2010.5686287-Figure1-1.png", + "caption": "Fig. 1. Free-body diagram of object B, showing a force, fi, and a torque, ti, at contact point i. O and G represent the reference point and center of mass, respectively.", + "texts": [ + " In the literature, such a transformation has been formulated by combining open-chain Jacobian matrices for each manipulator with a grasp matrix relating the velocity of each end-effector to the velocity of the object (both linear and angular). This tacitly assumes that all object degrees of freedom are specified in the primary task. To free up some object DOFs, we adopt a closed-chain perspective that collectively modifies the grasp and Jacobian transformations. We first review the open-chain transformations and then we combine them to form the closed-chain transformation for a reduced set of object DOFs. The free-body diagram of the object and the coordinate system are shown in Fig. 1, where N and B represent the ground and body reference frames, respectively. Suppose that there are n points of contact between the robot and the manipulated object. For i = 1, . . . , n, let ri be the position vector from the reference point to the ith contact point and let vi and \u03c9i be the velocity and angular velocity of an end-effector frame whose origin coincides with that point. The acceleration of the body B and the acceleration of the ith contact frame of reference are related as: v\u0307i = v\u0307 + \u03c9\u0307 \u00d7 ri + \u03c9 \u00d7 (\u03c9 \u00d7 ri) + 2\u03c9 \u00d7 vreli + areli \u03c9\u0307i = \u03c9\u0307 + \u03b1reli ", + " (12) one may take E as E = rT 1 0 0 0 rT 2 0 rT 2 rT 1 0 \u03b1I3 \u2212 \u03b3r\u00d7 2 \u03b2I3 + \u03b3r\u00d7 1 \u2212I3 (13) While the derivation of E in this case is not obvious, one may check that it annihilates GS\u22a5 and that it is full rank. Before turning to the control law, we still need to understand the net contribution of the contact forces on the object. This includes both the external dynamics as well as the internal forces acting on the object. For the external dynamics, consider once again the freebody diagram in Fig. 1. The equation of motion for the object can be expressed as follows. F ma = F + GT f + mg\u0302 (14) F ma . = ( maG IG\u03c9\u0307 + \u03c9 \u00d7 IG\u03c9 + rG \u00d7 maG ) g\u0302 . = ( g rG \u00d7 g ) Here, F ma is the inertial forces written in terms of: m, the mass of the object; IG, the moment of inertia about the center of mass, G; aG, the acceleration of point G; and rG, the position vector from the reference point O to G. On the right-hand side, f is the column matrix of contact forces, f i, and contact torques, ti, (see Fig. 1) arranged to mirror the list of velocities, vi, and angular velocities, \u03c9i, that appear in x\u0307. Also, F is the net external wrench (force and moment) about point O, and g is the gravity vector. For the internal forces, one can see from (14) that they are defined by the null space of GT . Our approach is to use the relative acceleration terms to control the internal forces; hence, they too must lie in the same space. For the sake of this work, we will control the interaction forces between the contacts" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003842_012008-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003842_012008-Figure1-1.png", + "caption": "Figure 1. Schematic of TC4 powders prepared by EIGA.", + "texts": [ + "999%) supplied by Sichuan Pangang Messer Gas Products Co., Ltd. During the experiment, the TC4 rod rotated at low speed into the water-cooled copper induction coil for heating. The molten metal droplets were continuously dripped and atomized by the high-speed argon gas sprayed from the nozzle. Then, spray droplets were spheroidized and solidified into spherical particles. At last, TC4 powders were collected in the powder storage tank through the cyclone. Schematic of the atomizing processes was illustrated in figure 1. The morphology and microstructure of TC4 powders were investigated using scanning electron microscopy (SEM, MLA650F, FEI, USA). The particle size distributions of TC4 powders were observed using a laser micron sizer (LMS, ZS90, Malver, UK). The phase constituents of TC4 powders were analyzed by X-ray powder diffraction (XRD, Empyrean, PANalytical B.V., The 1st International Conference on Metals and Alloys IOP Conf. Series: Materials Science and Engineering 668 (2019) 012008 IOP Publishing doi:10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003075_tasc.2019.2891534-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003075_tasc.2019.2891534-Figure9-1.png", + "caption": "Fig. 9 Platform of prototype test", + "texts": [ + " Copper loss, stator loss and eddy current loss of the permanent magnet can cause the motor temperature to rise. In this paper, the heat conductivity and heat generation rate of each part were calculated by using loss as heat source. And the three-dimensional temperature field simulation model of the motor was established in FLUENT. The test platform bracket will have a great impact on the motor temperature rise. Therefore, a simulation model of temperature field including bracket is established in this paper, as shown in Fig. 9. During the simulation, the initial ambient temperature is 293k and the simulation duration is 30min. The temperature distribution of the whole motor and stator core is shown in Fig. 10. The temperature rise is 35K when sinusoidal power is supplied, and the temperature rise is 43K when inverter power is supplied. Tem (K) 328 325 323 320 318 315 313 311 Tem (K) 336 333 330 327 324 321 318 316 (a) (b) Tem (K) 328 326 324 323 321 320 318 317 Tem (K) 336 334 332 330 328 326 324 323 (c) (d) Fig. 10 Temperature distribution of the whole motor and stator core" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001949_12.2257550-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001949_12.2257550-Figure4-1.png", + "caption": "Figure 4: Coordinates for location.", + "texts": [ + "20,21 In this section, we first present a method to estimate a sphere-shaped target\u2019s relative location in the UAV\u2019s inertial frame using a single camera carried by the UAV. We then present an overview of UAV\u2019s kinematic model. In the end, a vision-based control law is developed for the UAV and gimbal to together track the mobile target. To find the sphere-shaped target\u2019s relative location in the UAV inertial frame XiYiZi (North-West-Up), we first locate the target in the camera frame, and then transfer it to the body frame and finally the inertial frame. As shown in Figure 4, a sphere with radium r projects an ellipse P with center OB\u00b5\u03bd onto the image plane \u00b5\u03bd. Proc. of SPIE Vol. 10169 101690A-6 Downloaded From: http://proceedings.spiedigitallibrary.org/pdfaccess.ashx?url=/data/conferences/spiep/92321/ on 04/26/2017 Terms of Use: http://spiedigitallibrary.org/ss/termsofuse.aspx We first locate the sphere\u2019s position in the camera frame XcYcZc using A and the position (mB , nB) of OB\u00b5\u03bd measured on the image plane and also the sphere\u2019s radius r. We use (xcB , ycB , zcB) to denote the sphere\u2019s position in XcYcZc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002839_j.mechmachtheory.2018.06.021-Figure10-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002839_j.mechmachtheory.2018.06.021-Figure10-1.png", + "caption": "Fig. 10. Simplified example of traction drive.", + "texts": [ + " Furthermore, the logarithmic CVT is actually an optimization of the half-toroidal CVT, and the inner cone and outer cone CVTs are two optimizations of the logarithmic CVT. Because of the intimate connection, the half-toroidal CVT, logarithmic CVT, inner cone CVT and outer cone CVT are selected to compare with each other in this study. The comparison is conducted in the four aspects of spin ratio, efficiency, contact ranges of components and manufac- turability. Spin motion, which is directly proportional to spin loss, is expressed by a spin ratio that can be computed by a widely accepted mathematical model. Spin speed ( Fig. 10 ) can be expressed as \u03c9 spin = \u03c9 dg sin \u03b8dg \u2212\u03c9 dn sin \u03b8dn , (43) where \u03c9 spin is the spin speed; \u03c9 dg and \u03c9 dn are the rotational speeds of the driving and driven components, respectively; \u03b8dg is the angle between the contact area and the rotational axis of the driving component; and \u03b8dn is the angle between the contact area and the rotational axis of the driven component. As such, \u03c9 dg sin \u03b8dg and \u03c9 dn sin \u03b8dn represent the vertical velocity components of driving and driven rotational speeds, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000477_s11771-011-0748-9-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000477_s11771-011-0748-9-Figure2-1.png", + "caption": "Figure 2 shows the sketch of the PECF system for finishing the SBG surface. The gear and the tool were securely held by a universal dividing head and a clamp, respectively. Both the gear and the tool were insulated from the main body in order to only focus an electrochemical reaction between them. As the tool scanned over the tooth surface, the whole surface was finished.", + "texts": [], + "surrounding_texts": [ + "J. Cent. South Univ. Technol. (2011) 18: 685\u2212689 686\nunder different working conditions [10]. But there existed more assumptions in the model. Consequently, an exact mathematical model is necessary for controlling the surface quality effectively in the PECF.\nLeast square support vector machine (LSSVM) has been successfully applied in pattern recognition and function estimation problems [11\u221213]. It was developed from the standard support vector machine (SVM) [14\u221215]. The LSSVM is computationally more efficient than the standard SVM method, since the LSSVM training only requires the solution of a set of linear equations instead of the long and computationally hard quadratic programming problem involved in the standard SVM [16].\nAs a complicated and nonlinear process, the results of PECF are difficult to control. Usually, the trial-anderror method is used to select the finishing parameters. However, it is difficult for this method to obtain the desired surface roughness. In this work, the LSSVM method with a radial basis function (RBF) kernel is employed to develop an offline model which can predict the surface roughness and select the finishing parameters in PECF.\nConsidering a given training set {xi, yi} (i=1, 2, \u2026, N), where xi\u2208Rn represents a n-dimensional input vector and yi\u2208R, which is a scalar measured data, represents the system output. The form of this function is defined as [17]\nwhere w is the weight vector, b is the bias term and \u03c6(\u00b7) is a nonlinear function that maps the input space into a higher dimension feature space.\nThe regression model (Eq.(1)) can be constructed with a nonlinear mapping function \u03c6(\u00b7). By mapping the original input data into a high-dimensional space, the nonlinear separable problem becomes linearly separable in space. Now, the objective is to find the optimal parameters that minimize the prediction error of the regression model. Hence, consider the following optimization problem with equality constraints:\n( )\n( )\nT 2\n1\nT\n1 1min , 2 2\ns.t.\n1, ,\nN\ni i\ni i i\nJ e\nb e i N\n\u03b3\n\u03d5\n= = +\n= + + =\n\u2211w e w w\ny w x L\n(2)\nwhere ei is the random error and \u03b3\u2208R+ is a regularization parameter in determining the trade-off\nbetween minimizing the training errors and minimizing the model complexity.\nTo solve the above optimization problem, the following Lagrange function is constructed:\n( )T\n1 ( , , ; ) ( , ) { }\nN\ni i i i i L b J a b e\u03d5 = = \u2212 + + \u2212\u2211w e w e w x y\u03b1 (3)\nwhere \u03b1 is the Lagrange multiplier. The solution of Eq.(3) can be obtained by partially differentiating with respect to w, b, ei and \u03b1i:\n1\n1\nT\n0 ( )\n0 0\n0 ( 1, , )\n0 ( ) 0 ( 1, , )\nN\ni i i\nN\ni i\ni i i\ni i i i\nL\nL b L e i N e L b e i N\n\u03b1 \u03d5\n\u03b1\n\u03b1 \u03b3\n\u03d5 \u03b1\n=\n=\n\u23a7 \u2202 = \u21d2 =\u23aa\u2202\u23aa \u23aa\u2202\u23aa = \u21d2 = \u2202\u23aa \u23a8 \u2202\u23aa = \u21d2 = =\u23aa\u2202 \u23aa \u2202\u23aa = \u21d2 + + \u2212 = =\u23aa\u2202\u23a9 \u2211 \u2211 w x w w x y L L\n(4)\nIn the form of matrix, the above equations can be\nexpressed as\nT\n1\n0 0b \u03b3 \u2212 \u23a1 \u23a4 \u23a1 \u23a4 \u23a1 \u23a4 =\u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5 + \u23a3 \u23a6 \u23a3 \u23a6\u23a2 \u23a5\u23a3 \u23a6 yI \u0398 \u03b1\u0398 \u03a9\n(5)\nwhere y=[y1, \u2026, yN]T, \u0398=[1, \u2026, 1]T, \u03b1=[\u03b11, \u2026, \u03b1N]T, \u2126 is a square matrix, and I is an identity matrix. The elements \u2126ij=\u03c6(xi)T\u03c6(xj)=\u03ba(xi, xj), i, j=1, \u2026, N. Finally, the estimated values of \u03b1 and b can be obtained by solving the linear system of Eq.(5), and the resulting LSSVM model can be expressed as\n1 ( ) ( , )\nN\ni i i b\u03b1 \u03ba = = +\u2211y x x x (6)\nwhere \u03ba(x, xi) is the kernel function. In comparison with some other feasible kernel functions, the radial basis function (RBF) is a more compact supported kernel and is able to reduce the computational complexity of the training process and improve the generalization performance of LSSVM. As a result, the RBF kernel is selected as the kernel function:\n2 2 1( , ) exp 2i i\u03ba \u03c3 \u239b \u239e= \u2212 \u2212\u239c \u239f \u239d \u23a0\nx x x x (7) where \u03c3 is the kernel width parameter.\nNote that the parameters \u03b3 and the value of \u03c3 from the kernel function have to be turned. Using the above LSSVM model with RBF kernel, the offline nonlinear model of the controlled system can be expressed as\n2\n2 1 1( ) exp 2\nN\ni i i b\u03b1 \u03c3= \u239b \u239e= \u2212 \u2212 +\u239c \u239f \u239d \u23a0 \u2211y x x x (8)", + "J. Cent. South Univ. Technol. (2011) 18: 685\u2212689\n687\nThe scheme of SBG tooth surface finishing by PECF is illustrated in Fig.1. Two electrodes were connected to a pulse power supply. The tool was the cathode and the gear was the anode. The electrolyte was pumped at high velocities through the interelectrode gap in order to remove the reaction products and to dissipate the heat generated.\nFig.1 Scheme of SBG tooth surface finishing by PECF\nAccording to the Faraday\u2019s law and Ohm\u2019s law, the interelectrode gap thickness can be calculated as [18]\n2 0 e v2 ( )S S k U U t\u03b7\u03ba \u03b4= + \u2212 \u0394 \u22c5 \u22c5 (9)\nwhere S0 is the initial interelectrode gap, \u014b is the current efficiency of anodic dissolution, \u03bae is the electrolyte conductivity, kv is the volumetric electrochemical equivalent, U is the applied voltage, \u0394U is the over potential value, \u03b4 is the pulse duty factor, and t is the finishing time. 3.2 Experimental setup\nThe pulse DC power supply was adopted in this system. The main ingredient of electrolyte was NaNO3. The experimental and workpiece parameters are listed in Table 1 and Table 2. The workpiece chemical compositions are listed in Table 3.\nAccording to the principle of PECF, the finishing process is influenced by tool rotating speed (nt), applied voltage (U), initial interelectrode gap (S0), pulse duty factor (\u03b4) and finishing time (t). In order to assess the effect of finishing parameters, the Taguchi method L16(45) was introduced and the interaction effect of parameters was estimated by carrying out 16 experiments. The main effect shows the contribution of individual parameter to the surface roughness in the experimental results. For generating the training data, all parameters were set by four levels, as listed in Table 4.\nThe input data {xi, yi} of LSSVM consist of five experimental parameters, including the tool rotating speed (nt), applied voltage (U), initial interelectrode gap (S0), pulse duty factor (\u03b4) and finishing time (t). The output data yi is the result of experiments, that is, the surface roughness Ra. Table 4 lists the experimental results of PECF. The sixteen groups of experimental data", + "J. Cent. South Univ. Technol. (2011) 18: 685\u2212689 688\nselected from Table 4 were used to perform the training. The values of \u03b1 and b could be obtained from Eq.(5). Toolbox of MATLAB was adopted for the algorithm of LSSVM. The nonlinear radial basis function (RBF) kernel was chosen, and variables \u03b3=15, \u03c3=2.45 were fixed. Consequently, the predictive LSSVM model can be constructed, as shown in Fig.3.\nThe other four groups of experimental results in Table 4 were used for verifying the availability of the\npredictive model. From Fig.4, the errors between the predicted values and the experimental values can be acquired. It can be concluded that the model can predict the surface roughness with reasonable accuracy.\nFig.4 Comparison between predicted values and experimental values of surface roughness\nThe mean absolute percent error (MAPE) is a very common criterion for evaluating the performance of a forecasting system since it gives the mean of the absolute. The MAPE is computed as\n1\n1 abs l\ni\ni E l =\n\u2212\u239b \u239e = \u239c \u239f \u239d \u23a0 \u2211 y y y (10)\nwhere y is the experimental value, yi is the LSSVM predicted value and l is the number of validation data. It can be seen form Fig.4 that the model provides good prediction performance with MAPE of 2.43% for the surface roughness of SBG in PECF.\nThe actual manufacturing process was constrained by the desired surface roughness and the machining cost. Usually, the trial-and-error method was used to select the finishing parameters. However, it is difficult for this method to obtain the desired surface roughness with a higher machining efficiency. Therefore, the prediction of the finishing parameters is also necessary.\nTable 5 lists the factorial analysis results of the orthogonal experiments of the PECF. It can be seen that the applied voltage has more effect on the surface roughness than the other finishing parameters.\nTherefore, the applied voltage (U) was chosen as the output data, and the surface roughness and the other four parameters were regarded as the input data. By applying the above procedure, a new LSSVM model was constructed to obtain the optimum applied voltage matching the desired value of surface roughness. By adjusting the variables \u03b3=35 and \u03c3=2.24, the predictive" + ] + }, + { + "image_filename": "designv11_33_0002586_apec.2018.8341172-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002586_apec.2018.8341172-Figure1-1.png", + "caption": "Fig. 1. Machine topologies of 12/8-pole (a) SSHE-SRM and (b) CSRM.", + "texts": [ + " Many different types of hybrid SRMs with PMs have been proposed to improve output torque, efficiency and power density [9-13]. In this paper, a 12/8-pole segmented-stator hybrid-excitation SRM (SSHE-SRM) with a novel boost converter (NBC) drive is proposed to boost the phase voltage and hence improve its dynamic performance. Detailed simulations and experiments of the SSHE-SRM drive with NBC are carried out as well as compared with AHBC. II. MACHINE TOPOLOGY AND MAGNETIC CHARACTERISITCS The machine topologies of 12/8 SSHE-SRM and CSRM with the same basic size are shown in Fig. 1. The stator of the SSHE-SRM is comprised of six segments. The PMs are embedded between two poles in each stator segment. There is neither yoke iron nor other ferromagnetic material between two adjacent C-core segments. The six C-shaped segments are evenly arranged along the rotor, which are characterized by their own independent magnetic structures. The flux generated by the PM is closed through the U-shaped segment and does not pass the rotor core and air gap. However, when the winding is excited with a current, the flux produced by excited coils and the flux produced by PM are added together" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001223_0954406211409805-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001223_0954406211409805-Figure1-1.png", + "caption": "Fig. 1 Planar revolute joint with clearances between bodies i and j", + "texts": [ + " and are positive constants that represent the feedback control parameters for the velocities and position constraint violations. In the real mechanism system, joint clearances influence the mechanism outputs with both kinematic and dynamic aspects. Therefore, it is necessary to introduce the joint clearance contact models (JCCM) in the mechanism model, which well depicts the real contact state of the clearances. This article uses two planar circles (denoted by pin circle and bushing circle) to represent the pin and bushing, which are, respectively, attached to rigid bodies i and j, as shown in Fig. 1. Then, contact model of the two planar circles is built to simulate the real contact state of joint clearances. Flores and Ambrosio [9] compared several kinds of the impact force models and pointed out that the impact force model proposed by Lankarani and Nikravesh was more reasonable, since it was the only model that accounted for the energy dissipation during the impact process. The formula of this impact force model is as follows FN \u00bc K 1:5 1\u00fe 3\u00f01 e2\u00de 4 _ _ \u00f0 \u00de \u00f02\u00de where is the penetration, e the coefficient of restitution, _ the relative penetration velocities, _ \u00f0 \u00de the impact velocities, and K a constant which is dependent on the material properties of the components and their geometry" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002159_gt2017-63208-Figure18-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002159_gt2017-63208-Figure18-1.png", + "caption": "Fig. 18 Oil distribution (simulation), oil-jet supply 3.7\u00d72L/min, input 10000rpm", + "texts": [ + " This is probably because the eddies in the cavity between the gear teeth are not well resolved in the simulation. Influence of Temperature Setting. The influence of the temperature setting was not substantial because the temperature dependence of the loss was small (double circles and open circles in Fig. 17(a), a double square and an open triangle in Fig. 17(a-1)). As a consequence, in fluid dynamic loss and windage loss, the experimental values approximately agree with the simulated values. Consideration on the correspondence between the lossclassification and the CFD. Figure 18 shows the simulated oil distribution (iso-surface of the oil volume fraction in the cells = 0.5). In Shroud 2, oil is observed below the smaller gear. It is considered that oil that is not discharged remains in the shroud because the opening area is small in Shroud 2. Possibly, the remaining oil causes the increase of oil reacceleration loss, as shown in Fig. 7(b). The validity of the loss-classification method suggested in this paper can be supported with the consistency between the increase of loss in the experiment and the flow visualization in the CFD" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003437_ceit.2018.8751810-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003437_ceit.2018.8751810-Figure7-1.png", + "caption": "Fig. 7. Membership functions for e\u0307x", + "texts": [], + "surrounding_texts": [ + "In Artificial Neural Networks (ANNs), the learning is realised by the adaptation of the network weights based on the input data and the corresponding desired outputs. At each iteration, the parameters are updated such that the output of the network will match with the desired value. Different algorithms for the adaptation of the weights have been proposed in the literature. Among those, gradient descent method is the most widely used one, in which the gradient of a cost function utilizing the difference between the desired and network outputs has been determined, and then the weights are updated in the opposite direction of this gradient to minimize the cost function [12]\u2013[14]. The main drawback of this method is that the network can stuck at a local minima, and as a result the network output will not converge to the desired value. To alleviate this problem, VSS-based learning algorithms can be employed, in which a sliding surface defined by an error term and its derivative has been generated for each layer of the network. The main advantage of this method is that it guarantees the robustness, stability and fast convergence of the overall system [15], [16]. Consider a fully connected, three layer neural network, where there are p neurons in the input layer, n neurons in the hidden layer and one neuron in the output layer. In this network, the vector X(t) = [x1(t), . . . , xp(t)] T is corresponding to the input signals (outputs of the input layer neurons), the vector UH(t) = [uH1 , . . . , uHn ]T is standing for the outputs of the hidden layer neurons and u(t) is the scalar output of the network. The matrix W1(t) (having a dimension of (nxp)) denotes the weights between the input layer and hidden layer, where each entry w1i,j(t) is corresponding to the weight of the connection between ith neuron in the hidden layer and ith neuron in the input layer. Similarly, the n-dimensional vector W2(t) represents the weights of the connections between the hidden layer and output layer. In this network structure, the outputs of the hidden layer neurons and the output of the overall network can be computed as follows: uHi = f p\u2211 j=1 w1i,jxj (1) u(t) = n\u2211 i=1 w2iuHi (2) To be able to use VSS-theory in the learning of the network, a sliding surface in the following form has been defined: s = e\u0307+ \u03bbe (3) where e, e\u0307 and \u03bb are the discrepancy between the desired and actual outputs of the network, the derivative of this discrepancy, and a positive constant corresponding to the slope of the sliding surface, respectively. For this sliding surface to be attractive, it should satisfy the following two conditions: \u2022 Once the sliding surface is reached, the system trajectories should keep moving along this surface for all subsequent times (s = 0 and s\u0307 = 0) \u2022 If the system trajectories are not on the sliding surface, they should approach it at least asymptotically (lims\u21920+ s\u0307 < 0 and lims\u21920\u2212 s\u0307 > 0) To be able to enforce the second condition, Lyapunov stability method is used [17]. In this method, if a Lyapunov candidate function that will satisfy the following two conditions can be determined, then the convergence of the system states to the sliding surface followed by sliding along it is guaranteed: V (s) = 1 2 s2 \u2265 0 and V\u0307 (s) = s \u2202s \u2202t \u2264 0 (4) With the following weight update rules, these conditions can be satisfied, and thus the convergence of the system states will be guaranteed: w\u03071i,j = \u2212 ( w2ixj XTX ) \u03b1sgn(e) (5) w\u03072i = \u2212 ( uHi UH TUH ) \u03b1sgn(e) (6) In this study, a neural network structure with two input neurons (one for the deviation from the reference trajectory in x- or y-direction and one for the derivative of this deviation), five hidden neurons and one output neuron has been employed for each control command signal. Simulations and experiments have been carried out for different number of hidden layer neurons, however no significant effect of this parameter on the overall system response has been observed. The learning rate is specified as \u03b1 = 15 for both controllers using trial-anderror method and the slope of the sliding surface has been set to \u03bb = 10 The weights are initially set in a random manner to show that a priori knowledge about the system is not required in ANNs. Simulations and experiments have been carried out for different number of hidden layer neurons, however no significant effect of this parameter on the overall system response has been observed. The results of the simulation and real time experiments are presented in Figures 12 and 13, respectively. For the simulations, the obtained results are similar to the ones of the other two controllers. In the experiments, deviations from the reference trajectory become especially salient at the corners. These are points where there is a sudden change in either x- or y-direction, and because of this change the network needs to readjust the weight values. The time required for the adaptation and delays associated with the communication give rise to these circular patterns at the corners. The advantage of the ANN with VSS-based learning algorithms becomes prominent in Figure 13. As can be observed the discrepancy between the reference and the system output is rather smaller compared to PID and fuzzy logic controllers." + ] + }, + { + "image_filename": "designv11_33_0001580_iros.2011.6094712-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001580_iros.2011.6094712-Figure1-1.png", + "caption": "Fig. 1. Schematic illustration of the adopted continuum model. The red and blue lines denote multi-articular muscles.", + "texts": [ + " Then, the joint torque on the ith joint, \u03c4i, is obtained as \u03c4i = k(vd \u2212 v)(\u03c6i\u22121 \u2212\u03c6i), (5) where k is the control gain that governs longitudinal acceleration and vd is the desired longitudinal velocity. Equation (5) simply denotes P control of the posterior joints, whose references are the anterior joint angles. In this section, we extend the continuum model described above to the case where the body is actuated by multiarticular muscles. Here, we consider two-dimensional motion, i.e., locomotion on a flat terrain. The adopted model is shown in Fig. 1. The body consists of a continuous curve of length L and zero thickness; hereafter, this curve is referred to as the backbone curve. The backbone curve is parameterized by the arclength s \u2208 [0,L], where s = 0 and s = L denote the head and tail ends, respectively. The curvature of the backbone curve is denoted by \u03ba(s). Rigid rectilinear links of length 2r are aligned such that they are perpendicularly bisected by the backbone curve. Each link is numbered as i (1 \u2264 i \u2264 N), and it intersects the backbone curve at s = (i\u22120" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003616_j.promfg.2019.08.022-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003616_j.promfg.2019.08.022-Figure3-1.png", + "caption": "Fig. 3. Scheme of cutting samples for mechanical tests on static bending from blanks obtained by laser deposition of Inconel 625 alloy on steel 15Cr11MoW", + "texts": [ + " To determine the values of impact toughness of welded joints, blanks were cut from witness samples, which made it possible to manufacture impact specimens with a V-notch located in the zone of fusion of the base metal and cladding. Samples of welded joints made of steel 15Cr11MoW with overlaying with Inconel 625 alloy in the initial state after overlaying and after overlaying and subsequent high-temperature tempering were tested. The tests were carried out at room temperature. According to the test results, the ability of the deposited layer obtained by the method of laser-powder surfacing to withstand a given plastic deformation (bending angle) was evaluated. Samples of type XXVI measuring 4\u00d720\u00d7120 mm (Fig. 3) were cut from the witness sample in such a way, that the weld metal, fusion zone and base metal were located along the sample and were in the working zone during bending. The thickness of the samples corresponded to the thickness of the entrance edge of the blade after recovery. This allowed us to simultaneously evaluate the plastic properties of all zones of the welded joint, including the weld and M. Kuznetsov et al. / Procedia Manufacturing 36 (2019) 163\u2013175 167 Author name / Procedia Manufacturing 00 (2019) 000\u2013000 5 the base metal" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002503_1350650117753915-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002503_1350650117753915-Figure7-1.png", + "caption": "Figure 7. Transient pressure distribution.", + "texts": [ + " The calculation considering the rotation of the ring holes makes the calculation of the new mesh in every time step difficult to converge. Although this simplification might hide the fluctuations of the flow field caused by the rotation of the floating ring, this method is adopted considering the feasibility in practice. In addition, this simplification is considered a good approximation when considering vibration problems. The detailed results are analyzed in the following section. The transient calculation results demonstrate that the impact of the motion of the floating ring and shaft is significant. Figure 7 shows the pressure distribution at different time steps at 10 kr/min shaft speed, whereas Figure 8 shows the 10 kr/min shaft speed air entrainment condition. Given the transient effect, the position of the high pressure zone fluctuates with time. Air volume fraction also fluctuates with the fluctuation of the low pressure zone and with the effect of rotation. DFT is a signal processing method that converts time domain signals into frequency domain. This method is adopted to analyze the vibration frequency of the rotor using transient CFD results" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002149_gt2017-63495-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002149_gt2017-63495-Figure6-1.png", + "caption": "Figure 6 Three-pad HAFB with three orifice tubes", + "texts": [ + " Externally pressurized air is injected into the bearing clearance through those orifice tubes. Figure 4 shows the top foil with welded orifice tube. When the top foils are assembled on to the bearing sleeve, the orifice tubes are located at 60 , 180 (direction of gravitational loading) and 300 . Figure 5 shows orifice tubes configuration for the three-pad HAFB. Orifice tubes are connected to the main supply pressure line, and the controlled hydrostatic injection is achieved by controlling the air pressure to the orifice tube located at 180 through a separate on/off valve. Figure 6 shows the three-pad HAFB. 4 Copyright \u00a9 2017 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 08/23/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use The rotor is constructed with several components. Bearing journals are at each side of induction motor element, and thrust runner and end cap (with the same weight as the thrust runner) are at the each ends and they are all assembled with two tension bolts, ensuring the first bending critical speed of the rotor assembly far above the maximum operating speed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002730_s13369-018-3419-4-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002730_s13369-018-3419-4-Figure1-1.png", + "caption": "Fig. 1 Modification curves on the rotary project plane", + "texts": [ + " To reduce vibration and noise and improve the scuffing load capacity of face gear drives, a multi-objective optimization model is established with uniform TSLD, minimumWALTE and minimum TSFT as the three objectives based on tooth contact analysis (TCA) and loaded tooth contact analysis (LTCA). Topography modification has been considered the composition of a tooth profile modification and an axial modification, which can reduce meshing impact and vibration and evenly distribute tooth load, ensuringmeshing stability. In this paper, a tooth profile and an axial modification curve on a spur pinion are designed with a straight line and two sections of parabola, as shown in Fig. 1, where l1 and l3 are the modification amounts in the profile direction, l2 and l4 are the corresponding modification lengths, and l5, l7 and l6, l8 are the modification amounts and the corresponding modification lengths in the axial direction. Symbols ra and rh denote the diameter of the addendum circle and the tangent of the fillet and involute profile on the rotary projection plane. The modification curve was designed with three segments, with the advantage of the free combination of different values li (i = 1 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001993_tac.2017.2707602-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001993_tac.2017.2707602-Figure4-1.png", + "caption": "Fig. 4. The i-th agent will increase its communication range until it discovers all of its neighbors in the topology of V\u03d1(Z 0) (or, equivalently, in the topology of X\u03d1(Z 0)). To simplify this illustration, we have assumed that \u00b5i \u03d1 = \u00b5 j \u03d1 = 0, which implies that the ellipsoids E i \u03d1(\u03b7i), E j \u03d1(\u03b7i) and E j \u03d1(\u03b7i) are all equal modulo a linear translation.", + "texts": [ + " In the light of the Rayleigh quotient inequality together with the definition of E j \u03d1(\u03b7i), we have that \u03b4\u0304i\u03d1(\u03b7i) \u2265 |\u03a0 1 2 (\u03bej\u03d1 \u2212 x)|2 \u2265 \u03bbmin(\u03a0)|\u03bej\u03d1 \u2212 x|2, for all x \u2208 E j \u03d1(\u03b7i), which in turn implies that |\u03bej\u03d1 \u2212 x| \u2264 \u221a \u03b4\u0304i\u03d1(\u03b7i)/\u03bbmin(\u03a0) =: \u03c8i \u03d1(\u03b7i), (28) for all x \u2208 E j \u03d1(\u03b7i). It follows immediately that B(\u03bei\u03d1;\u03c8 i \u03d1(\u03b7i)) \u2287 E j \u03d1(\u03b7i). Now, let us consider the stripe S that is formed by the collection of all balls of radius \u03c8i \u03d1(\u03b7i) which are translations of B(\u03bei\u03d1;\u03c8 i \u03d1(\u03b7i)) and are tangent to 0018-9286 (c) 2016 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. the boundary of X i \u03d1(z 0 i ; \u03b7i), as is illustrated in Fig. 4 (in this figure, the balls of the stripe S are depicted with dashed lines). Note that any generator \u03be\u2113\u03d1 \u2208 {\u039e\u03d1} that does not belong to X i \u03d1(z 0 i ; \u03b7i) and whose distance from bd(X i \u03d1(z 0 i ; \u03b7i)), in terms of \u03b4\u03d1(\u00b7; z0 \u2113 ), is less than or equal to \u03b4\u0304i\u03d1(\u03b7i) will belong to S. In addition, the stripe S will be contained itself in the closed ball B(x0 i ; \u03c8\u0304 i \u03d1(\u03b7i)), which is centered at the initial location of the i-th agent and has radius \u03c8\u0304i \u03d1(\u03b7i) := \u03c8i \u03d1(\u03b7i) + d\u0304i\u03d1(\u03b7i), (29) where d\u0304i\u03d1(\u03b7i) denotes the maximum (Euclidean) distance between x0 i and the boundary of X i \u03d1(z 0 i ; \u03b7i), that is, d\u0304i\u03d1(\u03b7i) := max{|x0 i \u2212 x|, x \u2208 bd(X i \u03d1(z 0 i ; \u03b7i))}" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003397_s00170-019-03894-w-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003397_s00170-019-03894-w-Figure3-1.png", + "caption": "Fig. 3 Free-form milling of spur gear with residual error using end mill", + "texts": [ + " By changing the tooth profile generating parameter u, the rotation angle of workpiece \u03c6 at any point on the tooth profile, as well as the tool setting parameters (a, e), can be obtained. Consequently, the cutter location point at any point on the tooth profile can be derived. In the process of free-form milling of gears, rather than to calculate all the cutter location points on tooth surface, the tool path is calculated to approach the theoretical tooth surface on the basis of satisfying tooth accuracy. Taking the common free-formmilling of spur gear with end mill as an example, the calculation of tool path including residual error is illustrated in Fig. 3. Note that point A is the first cutter location point which is tangent to tooth profile, point B is the adjacent cutter location point which is tangent to tooth profile as well, point C is the intersection of these two tangent lines AC and BC, and point D is the normal intersection of tooth profile curve in the plane of the drawing. The length of line CD represents the residual error \u0394t in the plane of the involute profile in case of spur gear, but in case of helical gear, the surface error should be measured in orthogonal direction to the surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003195_s1068798x18120420-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003195_s1068798x18120420-Figure1-1.png", + "caption": "Fig. 1. Plane burnishing machine.", + "texts": [ + " In the present work, we investigate the influence of the basic parameters of transverse burnishing on the surface roughness. On that basis, we may assess the utility of transverse burnishing by f lat plates as a method of finishing and strengthening. In determining the influence of the burnishing conditions on the surface microprojections, the basic parameters considered in the experiments are the compressive strain and the speed of workpiece rotation. A plane burnishing machine is used for transverse burnishing (Fig. 1). The system operates as follows: rotation is transmitted from electric motor 1 through 1 worm gear 2 to working screw 3. Rotation of screw 3, which passes through a threaded hole in mobile plate 4, is converted to linear motion. Plate 4 moves from left to right and, in tandem with plate 5, is responsible for transverse burnishing of part 6. Plate 4 is returned to its initial position by reversing motor 1. For the experiments, we use circular samples made by turning. The metals employed differ markedly in mechanical properties: \u2014steel 45: E = 200 GPa, \u03c3y = 360 MPa; \u03c3u = 590 MPa; HB = 174\u2013217 MPa; \u2014BrAZh9-4 bronze: E = 116 GPa, \u03c3y = 280 MPa; \u03c3u = 540 MPa; HB = 100\u2013120 MPa; \u2014D1 duralumin: E = 72 GPa, \u03c3y = 275 MPa; \u03c3u = 410 MPa; HB = 45\u201395 MPa" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000598_icems.2011.6073667-Figure10-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000598_icems.2011.6073667-Figure10-1.png", + "caption": "Fig. 10. FEM model: a) the PM machine 4 with 24slots/28poles; b) the PM machine 5 with 27 slots and 24 poles.", + "texts": [ + " Stator diameter, slot height, rotor diameter, air gap length, magnet thickness, ratio of slot opening and tooth pitch and ratio of pole width and pole pitch are the same. Therefore, material cost of all PM machines is nearly the same. Table IV gives some major specifications of studied PM machines. TABLE IV: MAJOR PARAMETERS OF STUDIED PM MACHINES Rated current, A 10 Tooth width, mm 5 Magnet thickness, mm 2 Stator length, mm 35 Magnet remanence, T 1.2 Outer rotor diameter without flywheel, mm 170 Fig. 9 and Fig. 10 present flux density distribution using nonlinear transient 2D-FEM simulation for the studied PM machines, but that of the machine 2 is not shown here because it has the same combination to the machine 1. Performances are presented in this section, that is, electromotive force (EMF), rotor eddy current loss, other losses, mean torque, torque density, power density and efficiency. A1. EMF and Cogging Torque Fig. 11 presents EMF rms versus combination of slot and pole number. For this calculation, PM machines have the same number of turns per phase" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003466_j.mechmachtheory.2019.06.017-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003466_j.mechmachtheory.2019.06.017-Figure2-1.png", + "caption": "Fig. 2. A face-gear mating with a tapered involute pinion.", + "texts": [ + " However, the aforementioned face-gear drives can rarely satisfy the requirements of independent backlash adjusting in different power flow branches in the concentric split-torque system. As a result, the load sharing between the idlers and pinions will be seriously affected. A new kind of face-gear drive has been presented to implement the independent backlash adjusting in the concentric split-torque transmission system [26] . A tapered involute gear was introduced as the pinion member. The face-gear was generated by a tapered involute shaper ( Fig. 2 ). Due to the variation in tooth thickness of the tapered involute pinion, axial movement of the pinion provides feasibility for backlash adjusting of the face-gear drive. This allows independent setting of backlash for each of the idlers and pinions in the concentric split-torque transmission system. Though brief introduction about tooth geometries of the tapered involute pinon and the mating face-gear have been performed [27,28] , the concrete tooth generations, tooth characters, contact simulations, and the stress analyses were yet to be conducted" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001147_demped.2011.6063611-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001147_demped.2011.6063611-Figure4-1.png", + "caption": "Figure 4. 3D finite-element model for alternator with stator turn fault", + "texts": [ + " Therefore, in the voltage frequency spectrum, the stator turn fault will give rise to the harmonic component at one third of the rectifier ripple frequency. Since a large capacitive load (the battery) is always connected to the alternator output, such harmonic component is even more evident in the current spectrum of the alternator. IV. FINITE-ELEMENT MODELING OF ALTENATORS WITH STATOR TURN FAULTS In order to further validate the proposed fault signature, an alternator with stator turn fault is modeled and simulated in Ansoft Maxwell 12.2 \u00a9. Fig. 4 shows the 3D finite element model of one pole-pair of the claw-pole alternator. Although the specific alternator being studied has a wave-wound stator winding, the finite element model uses a lap wound coil with the same number of conductors to help reduce the computational time while maintaining the same physics of the alternator. As shown in Fig. 4, the Phase B and Phase C winding each contains only one lap coil. The Phase A winding, however, is split into two lap coils. One of the lap coils is short-circuited to itself through a small resistor to simulate the stator turn fault. The winding connections are illustrated more clearly in Fig. 5. In Fig. 5, the three-phase windings are connected in delta. A fraction of turns in Phase A (depending on the severity of the turn fault) is split from the entire winding and is shorted through a small resistor. The shorted winding is still electrically connected in series with the rest part of Phase A, and is still magnetically coupled with all three phases as shown in the 3D model in Fig. 4. Fig. 6 shows the waveforms of alternator output current obtained from the finite-element simulation. Running at 3139 rpm, the alternator has a ripple frequency of 1883 Hz. It is clear in Fig. 6 that not all ripples have the same amplitude. There exists a repeating pattern of a large ripple followed by two small ripples, which agree well with the analytical analysis in section III. When the short circuit in the simulation is removed, the alternator output current is shown in Fig. 7. The repeating pattern is not obvious any more" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000238_icca.2010.5524364-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000238_icca.2010.5524364-Figure4-1.png", + "caption": "Fig. 4. Trajectory of SDRE Guidance", + "texts": [ + "2: Although maneuvering target was not formulated in section IV, we still consider the case that target has a spiral motion as shown in Fig. 3. The radius of spiral is 300m at a constant speed of 400m/sec in the Y direction, the magnitude of velocity is 500m/sec. The scenario is the same as that given in example 5.1. The results are shown in Table II. Unfortunately, both SDC forms A1 and A2 couldn\u2019t capture target under these conditions. while SDC form A3 performs well even though target maneuvers. After some observations, we conclude that the regulation of \u03c1\u0307 via SDC forms A1 and A2 would result in miss of capturing target. Fig. 4 depicts the trajectories of one of the cases in the 3- dimensional space. It can be observed that missile guided by SDC forms A1 and A2 has a motion similar to that of target. Example 5.3: A most practical missile guidance law, namely PPN, is herein compared to SDC form A3. By using PPN law (aM = N\u2126 \u00d7 vM ) with navigation gain N = 5 and same scenario given in example 5.1, the simulation results are shown in table III. These results reveal that in the performance index time to go, missile guided by SDC form A3 outperforms the one guided by PPN" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003893_0954406219888240-Figure19-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003893_0954406219888240-Figure19-1.png", + "caption": "Figure 19. Diagram of the joint clearance and contact deformation.", + "texts": [ + " Once the dimensional constraint of spherical joint is a constant angle, the non-interference domain will be limited in a circular domain whose diameter is the square root of 2d2 (shown in Figure 18) instead of a square domain whose side length is 2d (shown in Figure 16), and the workspace of the spherical joint is still a part of the complete workspace. Approximate clearance and contact deformation models of the spherical joint As connectors for parallel manipulators, joints are used to transmit forces and motion, so that the joint contact deformations and joint clearances are unavoidable, as shown in Figure 19. Due to the complexity of the accurate mathematical models of the joint clearance and the contact deformation, the stiffness magnitude is the main concern in this paper, and the given joint clearance and contact deformation models are approximate. Since the joint clearances and joint contact deformations are closely related to the joint structure, the approximate clearance model and the contact deformation model of the ideal and combined spherical joints will be respectively discussed as follows" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002118_iemdc.2017.8002091-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002118_iemdc.2017.8002091-Figure2-1.png", + "caption": "Fig. 2. The cross section of the 12/8-pole WFDSM.", + "texts": [ + " The theoretical analysis on the new control scheme is laid. By the finite element analysis on the 12/8-pole WFDSG prototype, the energy conversion loops can be obtained. The correspondent experiment has been done to verify the simulation and analysis. Sponsored by Jiangsu Key Laboratory Research Fund of Hi-Tech Research for Wind Turbine Design. II. THEORETICAL ANALYSIS ON THE CONDUCTION ANGLE CONTROLLED RECTIFIER In this paper, a three-phase 12/8-pole prototype WFDSM is adopted in the generating system. The cross section of the WFDSM is plotted in Fig. 2. There are four large slots in the stator for placing two field coils. Four vent holes in the stators are helpful to heat dissipation. In the prototype, the stator pole arcs have same width as those of rotor pole arcs. The main parameters of the 12/8-pole WFDSM are shown in Table I. The wiring diagram with the controlled rectifier is shown in Fig. 3. Three power switches labeled as Qa, Qb and Qc, which can be IGBTs and MOSFETs, are added to make phase currents controllable. When the stator poles align with the rotor poles, the correspondent phase power switch turns on to commutate", + " The current variation rate changes to 0/ ( ) /a o adi dt e u L Therefore, in stage [2T/3, tz], ia would drop fast. The energy loop curve ABCD is plotted in Fig. 5. For comparison, the curve for the HWR is also plotted, which is labeled as AED. The area surrounded by an energy loop curve represents the accepted electric power of the corresponding phase in an electric period. It is clear that the CAC scheme is helpful to improve the generator power. To research the effect of CAC scheme, the further simulation and experiments have been accomplished. The 12/8-pole WFDSG in Fig. 2 has been selected as the object. The simulation results will be presented in part III. III. SIMULATION OF THE CAC SCHEME ON A PROTOTYPE WFDSG The WFDSG FEM model and the control circuit model are set up by Maxwell 2D software. The simulation results are plotted in Fig. 6. The WFDSG speed is 2000rpm, and the conduction angle is set to 8\u03c0/15. For both CAC scheme and the HWR topology, the filter capacitors are chosen to 2000\u03bcF, and the loads are 0.5\u03a9 resistors. In simulation, the windings resistance may cause additional voltage drop, and the forward voltage of the power switch Qa is not zero" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002319_i2017-11597-1-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002319_i2017-11597-1-Figure3-1.png", + "caption": "Fig. 3. A perpendicular field Hx superimposed to the wire field HI generates a complex trapping field geometry. Left column: the total squared field strength H2, color code blue (red): low (high) H2. Right column: potential energy of a superparamagnetic bead (color code blue: low energy) (A) For a strong enough Hx > HI the normal component of the magnetic gradient force changes sign and is radially repulsive on one side. (B) For sufficiently weak external field Hx < HI the trapping field is everywhere attractive towards the wire surface. The beads remain on the surface but experience azimuthal forces towards a single minimum energy position.", + "texts": [ + " (3) remain unchanged with respect to the pure wire field case (H0z = 0). That is, while chains become reoriented by the presence of the z-field, single particles experience no change of trapping force or distribution on the surface. In a second interesting case we impose an external field orthogonal to the wire axis (e.g., in the x-direction) H(r, \u03c6) = HI(r)e\u03c6 + H0xex. (4) Note that now the field combination loses azimuthal symmetry and depends both on the radial distance r and the azimuthal angle \u03c6, see fig. 3. A single bead in this field has an energy W (\u03c6, r) = \u2212\u03c7b\u03bc0Vb 2 ( I2 4\u03c02r2 \u2212 H0xI \u03c0r cos \u03c6 ) . (5) From the azimuthal component F\u03c6 = \u2212(\u2202W/\u2202\u03c6)/r of the gradient force F\u03c6 = \u03c7b\u03bc0Vb 2\u03c0 H0xI r2 sin \u03c6 (6) and its radial component Fr = \u2212(\u2202W/\u2202r) Fr = \u2212\u03c7b\u03bc0Vb 4\u03c02 I2 r3 ( 1 \u2212 2\u03c0H0xr I \u00b7 cos \u03c6 ) (7) we see that depending on the magnitude of H0x vs. I and the angular position \u03c6 the wire surface can be either attractive or repulsive in the radial direction. In the following we will focus entirely on the weak external field, H0x < I 2\u03c0r , where the wire is radially attractive3, i", + ", b)) the dynamical behavior is very much richer than in case a). The fields of the wire and external field, as well as the induced dipolar fields between beads, dynamically interfere in an interesting manner. The perpendicular field H0\u22a5(t) superimposes with the wire field HI(x) = I 2\u03c0|x|e\u03c6 and weakens it on one side but strengthens it on the other side of the wire leading to the formation of a wave-like field structure on the surface of the wire with an energy of the form W (\u03c6 \u2212 \u03c9t, r) (eq. (5)), see fig. 3 and the previous section for the static case. In the following we investigate how the propagation of spinning waves leads to a complex dynamical phenomenology, see figs. 5, 6. The behavior of clusters in spinning fields depends on two parameters: the field spinning angular frequency \u03c9 and the relative field intensity h0 = H0 HI . We find that when the frequency of the spinning wave is sufficiently small \u03c9 < \u03c9c, with \u03c9c a critical frequency, one observes clusters of beads rotating uniformly with the field around the wire without a change of shape" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002525_28465-ms-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002525_28465-ms-Figure6-1.png", + "caption": "Figure 6\u2014Schematic of the Safer Plug design showing the complex internal channels.", + "texts": [ + " Due to the wide range of manufacturing methods and variables, certification requires a level of bespoke tailoring to ensure the goal of the relevant standards is achieved without compromising on the integrity or overly restricting the manufacturer. It is envisaged that in the future, with the development of standards, there will be more prescriptive guidance, however, there will always be a need for a more goal based philosophy. This approach has been successfully applied to the first oil and gas component to be certified - a titanium gateway manifold for pipelines used in the oil and gas industry, manufactured by the Safer Plug Company (Davies, S.,2017), and independently certified by Lloyd's Register, Figure 6 and 7. This product is used in a high integrity application and took full advantage of AM design flexibility. The internal channeling would not have been possible without the use of AM as the manufacturing process. It is clear that the adoption of additive manufacturing by the offshore industry still has challenges. With the need to improve the quality control and production consistency as well as develop the full supply chain to support the industry. It is clear from the example set by other industries that have already adopted this technology, such as aerospace, medical and the early adopters in the offshore industry, these challenges can be overcome" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001063_robio.2011.6181298-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001063_robio.2011.6181298-Figure1-1.png", + "caption": "Fig. 1. Displacement of the center of gravity", + "texts": [ + " In such cases, it is difficult for humanoid robots to move using a conventional stepping walk, as higher stability is necessary. It has been suggested that walking and turning using a foot sliding motion could provide this stability. In a sliding walk, it is not necessary to lift the legs or stand on one leg. Using a sliding walk, the robot also has a larger support polygon in terms of its Zero Moment Point (ZMP)[2]. As the target ZMP becomes wider, it is possible to achieve a smaller swinging motion, in comparison to conventional stepping walk (Fig. 1(a)). Besides, as the robot is supported by both legs during the sliding walk(Fig. 1(b)), it is possible to walk with a small displacement of the center of gravity, resulting in a shorter robot movement time. Recently, there has been a great deal of research on narrow space motions and slip motions of biped robots. Harada et al. [3] realized narrow space movement with a conventional S. Tsuichihara, S. Sugiyama, and T. Yoshikawa are with the College of Information Science and Engineering Ritsumeikan University, 1-1-1, Nojihigashi, Kusatsu, Shiga 525-8577, Japan. satoki-t@is.naist" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000451_978-3-642-29329-0_6-Figure6.1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000451_978-3-642-29329-0_6-Figure6.1-1.png", + "caption": "Fig. 6.1 Reference system", + "texts": [ + " All the control systems for enhancing stability (such as ABS and ESP) are often designed by means of the simple model presented in this chapter, the validation tests being performed on slippery surfaces [2]. The derivation presented hereafter refers to [16]. The reader who is not an expert in vehicle dynamics can follow step by step the derivation of the ordinary differential equations which describe the motion of the vehicle and which will be used in the bifurcation analysis. The reference system adopted to write the equations of motion and to describe the fundamental kinematics is depicted in Fig. 6.1. The x-axis is directed towards the longitudinal axis of the vehicle body, whereas the y-axis is orthogonal to the previous one and passes through the centre of gravity G. The vertical z-axis (not drawn in the planar figure) passes throughG too. The three reference axes are defined by the respective vectors i; j; k. The velocity at the centre of gravityG isvG and its components along x and y are denoted as u e v. The centre of gravity G lies at the ground level. Since the motion takes place exclusively in the x y plane, the vertical component z vanishes", + " The respective distances a and b of G from the centre of the fore and of the rear axle define the position of the centre of gravity, i.e., l \u00bc a + b. The distance between the centres of the two wheels of the same axle is the track t. For simplicity, the front and rear tracks are supposed to be equal (this is almost true for most vehicles). The mass of the vehicle ism, and I is the inertia tensor of the vehicle. We assume that I is of the form I \u00bc Ix 0 0 0 Iy 0 0 0 Iz 2 4 3 5; which implies that the principal inertia axes are coincident with the axes of the reference system in Fig. 6.1. The error introduced by this simplification is negligible for the presented analyzes. The steering angle of a wheel is the angle between a line directed as the meridian plane of the wheel and the longitudinal axis of the vehicle. To preserve the correct kinematics (at least at vanishing speed), the two steering wheels have different steering angles. The inner wheel is steered by the angle di, which is larger than the one of the outer wheel de (see Fig. 6.2). From Fig. 6.2 we obtain tan\u00f0de\u00de \u00bc l r\u00fe t=2 ; tan\u00f0di\u00de \u00bc l r t=2 : Eliminating r from the above equations, we obtain 1 tan\u00f0de\u00de \u00bc t l \u00fe 1 tan\u00f0di\u00de ; which, expanded in Taylor series, gives de \u00bc di t l d2i \u00fe O\u00f0d3i \u00de: For practical values of the steering angles (not larger than a few degrees) we can fairly assume that the steering angles of the outer and inner wheels are equal, i", + " Thus the differential equations describing the motion of an automobile running on an even surface are m\u00f0 _v\u00fe ur\u00de \u00bc Fy1 \u00fe Fy2 ; Iz _r \u00bc Fy1a Fy2b: (6.1) The non linearities of such a second order autonomous system can be found in the expressions of the forces Fy1 and Fy2, which depend on the lateral slipsa1 anda2, which in turn can be expressed as function of the state variables. The lateral slip a of a wheel is the angle between the line directed as the meridian plane of the wheel and the velocity vector at contact point of the wheel on the ground. In general a vehicle model like the one depicted in Fig. 6.1 has four different lateral slips, one per each wheel (see Fig. 6.5). We have assumed that the vehicle can be described by one single rigid body. Thus the speed of the contact point on the ground of each wheel can be derived as a function of the speed of the centre of gravity vG and of the yaw velocity r. For a conventional two wheel steer vehicle, the following relationships hold: \u2022 Front left wheel: v11 \u00bc \u00f0u 1 2 rt; v\u00fe ra\u00de \u00bc) tan\u00f0d a11\u00de \u00bc v\u00fe ra u 1 2 rt \u2022 Front right wheel: v12 \u00bc \u00f0u\u00fe 1 2 rt; v\u00fe ra\u00de \u00bc) tan\u00f0d a12\u00de \u00bc v\u00fe ra u\u00fe 1 2 rt \u2022 Rear left wheel: v21 \u00bc \u00f0u 1 2 rt; v rb\u00de \u00bc) tan\u00f0 a12\u00de \u00bc v rb u 1 2 rt \u2022 Rear right wheel: v22 \u00bc \u00f0u\u00fe 1 2 rt; v rb\u00de \u00bc) tan\u00f0 a22\u00de \u00bc v rb u\u00fe 1 2 rt In general, the radius of the path is much bigger than the track (r t) so that, recalling that u \u00bc |r| R, we have u rj jt=2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003893_0954406219888240-Figure23-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003893_0954406219888240-Figure23-1.png", + "caption": "Figure 23. Parameters of the universal joint.", + "texts": [ + " As a result, the clearance model of the revolute joint can be approximated by the maximum values as Pcle R \u00bc c2, c2, c3, LR \u00fe ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi L2 R \u00fe 4dRc2 p 2dR , LR \u00fe ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi L2 R \u00fe 4dRc2 p 2dR , 0 2 6664 3 7775 T \u00f023\u00de For the universal joint, since the two rotating axes (Axis I and Axis II) of the two hinges are always orthogonal, the joint clearances will be discussed in the O-u0v0w0 system. According to the clearance of the revolute joint, the clearance model of the universal joint can be expressed as \u00f0 Pcle U \u00de 0 \u00bc c4 \u00fe c5, c4 \u00fe c5, 2c4, 0, 0, LU \u00fe ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi L2 U \u00fe 4dUc4 p dU \" #T \u00f024\u00de where c4, c5, LU and dU are given in Figure 23. Note that we assume that the gap c6 is greater than the gap c5. Translating the results into the O-uvw coordinate system, the joint clearance is Pcle U \u00bc R1\u00f0 \u00de\u00f0 Pcle U \u00de 0 \u00f025\u00de where R1( ) is the transformation matrix. Finally, taking equations (23) and (25) into equation (20), the maximum clearance model of the combined spherical joint can be obtained. Approximate contact deformation model of the ideal spherical joint For some heavy load parallel manipulators, the reaction force in the joint is large; therefore, the tiny contact deformation of the joints cannot be ignored" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000776_msec2010-34325-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000776_msec2010-34325-Figure2-1.png", + "caption": "FIGURE 2: MILLING GEOMETRY [7]", + "texts": [ + " Prediction of cutting forces in milling process is used for two purposes: (i) to determine the power usage and compare to the machine limit, and (ii) to monitor the tool wear, since an increase in the forces indicates increase in tool wear. Analytical, numerical or empirical approaches were proposed. An analytical force model for both end and face milling was developed by Seethaler and Yellowley (2000) by assuming that the major component of forces are proportional to the undeformed chip thickness [7]. Figure 2 shows the milling geometry. The mean cutting forces are determined as follows: ( ) ( )( ) ( ) ( ) ( )( ){ } ( ) ( ) ( ) ( ) ( )( )( ){ } 1 2 1 1 2 1 1 2 1 2 cos 2 cos 2 2 sin 2 sin 2 8 cos 2 cos 2 2 sin 2 sin 2 , 8 t x s t y s KaNs F r KaNs F r \u03c6 \u03c6 \u03c6 \u03c6 \u03c6 \u03c0 \u03c6 \u03c6 \u03c6 \u03c6 \u03c6 \u03c0 = \u2212 + + \u2212 = \u2212 \u2212 + + \u2212 (2) where K is the specific cutting pressure, a is axial depth, N is the number of flutes, st is the feed per tooth, r1 is the force ratio, \u03a61 is the insert entry angle , \u03a62 is the insert exit angle, and \u03a6s is the swept angle of cut [7]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001973_10402004.2017.1323146-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001973_10402004.2017.1323146-Figure4-1.png", + "caption": "Figure 4. Photograph of high-speed motor setup.", + "texts": [ + " The high-speed test rig used here is a modified EHL rig (manufactured by PCS Instruments, Acton, UK) as described in detail in an earlier publication (Hili, et al. (23)) and is, in essence, a higher speed version of those used in many previous investigations into emulsions (Ratoi-Salagean (6); Wan, et al. (11); Ratoi-Salagean, et al. (16), (17); Yang, et al. (18)). This is shown in Fig. 3. The test contact is that between a glass disc and a 19.05-mm-diameter steel ball, each driven by a separate electric motor, the latter being capable of 20,000 rpm (Fig. 4), so that, in pure rolling, the entrainment speed approached 20 m/s. The contact was illuminated with white light and the interference image viewed through an optical microscope. This allowed film thickness to be measured by an ultra-thin-film interferometry technique (Johnston, et al. (24)). It should be noted that the surface energy of the silica-coated disc used for these measurements was approximately 300 mmJ/m2, which is close to that of the glass disc used for the fluorescence tests. As a result, the effect on emulsion entrainment due to the different wetting properties of the two surfaces can be considered to be negligible", + " The dichroic filter allows this fluorescent light to reach the charge-coupled device of an intensified camera, while removing light with wavelengths that correspond to the incident laser beam. Fluorescent images obtained in this way are either presented directly (such as those shown in Figs. 6 and 8) or are averaged using Matlab software to give the intensity values (as shown in Fig. 11). Because the fluoresce technique and the optical interferometry technique require different microscopes, it was not possible to make these measurements simultaneously. As a result, the fluctuations observed in Fig. 4 do not correspond to the fluorescence measurements in Fig. 11. Figure B1. Schematic diagram of fluorescence microscopy technique." + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001039_09544119jeim730-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001039_09544119jeim730-Figure6-1.png", + "caption": "Fig. 6 Determining whether a point lies inside or outside the vasculature", + "texts": [ + " Thus, a trade-off between performance and efficiency should be considered. The following method indicates how to determine whether a point lies inside or outside the vasculature. Traverse the OBB tree from the root node until arriving at a node where the point does not lie inside the OBBs of its two child nodes, or reaching a leaf node. Among all the patches contained in these two OBBs or the OBB of the leaf node, find the one for which the distance from its centre to the endpoint is minimal. In Fig. 6, n is the normal vector of theFig. 5 The movement vector points into a GCP JEIM730 Proc. IMechE Vol. 224 Part H: J. Engineering in Medicine at NANYANG TECH UNIV LIBRARY on May 23, 2015pih.sagepub.comDownloaded from patch pointing into the inside of the vasculature. For point A, the angle between OA and n is larger than 90u, and thus it lies outside the vasculature. For point B, the angle between OB and n is smaller than 90u, and thus it lies inside the vasculature. The minimal distance is used to represent the distance between the endpoint and vascular wall" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001529_icphm.2011.6024357-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001529_icphm.2011.6024357-Figure5-1.png", + "caption": "Fig. 5 Planet gears with artificially created pitting damage", + "texts": [ + " The 1st stage planet gears are mounted on the 1st stage carrier which is connected to the 2nd stage sun gear by shaft #2. The 2nd stage carrier is located on shaft #3. Four accelerometers named LS1, HS1, LS2 and HS2 are located on the housing of the planetary gearbox as also shown in Fig. 4. The pitting damage was artificially created on one of the four planet gears in the 2nd stage planetary gearbox. Four levels of pitting damage were tested, i.e. baseline, slight, moderate, and severe pitting (Fig. 5). Vibration data were collected from four accelerometers with a sampling frequency of 10 KHz at two load conditions (i.e. no load and 10,000 lb load), and four speeds of the driven motor (i.e. 300, 600, 900, and 1200 revolutions per minute, RPM). For each damage level, 10-minute data were recorded at each of the eight combinations of speed and load conditions, and were split equally into 20 samples. Totally 160 samples were collected for each pitting damage level. IV. DIAGNOSIS OF PITTING DAMAGE LEVELS The traditional techniques for vibration-based gear damage diagnosis are typically based on statistical properties of the collected vibration signal" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000631_j.proeng.2012.04.108-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000631_j.proeng.2012.04.108-Figure1-1.png", + "caption": "Fig. 1. (a) Experimental setup for impact testing; (b) definition of maximum ball deformation", + "texts": [ + " Therefore, the balls were individually transported in an insulated container to the testing laboratory (22 \u00b1 2\u00b0C) where they were each subjected to a single impact. The mean and standard deviation for the time between removal and impact was 3 \u00b1 1 minute. The surface temperature of each ball was measured in the climate chamber and immediately after impact using a digital thermocouple (Table 2). When referring to ball temperature the mean value prior to impact will be used (i.e. -0.8\u00b0C, 22\u00b0C, 35.9\u00b0C). A modified pitching machine (BOLA, UK) was used to project the golf balls without spin, normal to a fixed rigid surface (Fig. 1a). The range of impact speeds was 25 to 48 m\u00b7s-1, which was comparable to the values used by Strangwood et al. [11] when investigating golf ball dynamics. The impacts were filmed using a high speed camera placed perpendicular to the plane of motion and operating at 37,000 Hz (Phantom V4.3, Vision Research). The camera started filming when the ball passed a light trigger in the barrel of the pitching machine. The video footage was manually analyzed using in-house software. COR was calculated from the inbound and rebound velocity. The definition of maximum deformation is shown schematically in Fig. 1b. Impact duration was measured visually from the video footage, with an error of one frame corresponding to 27 s. A repeatability study was undertaken to assess the uncertainty in the manual analysis. An impact at a low, medium and high impact speed was analyzed ten times. The standard deviation was taken as a measure of the uncertainty for each impact. Averaging (mean) across the three impacts gave the uncertainty associated with each measure (Table 3). An FE model of the golf ball and fixed rigid surface was produced using Ansys/LS-Dyna", + " COR, maximum deformation and impact duration are each plotted against inbound velocity. The three series correspond to the temperatures under investigation, -0.8\u00b0C, 22\u00b0C and 35.9\u00b0C. COR and impact durations decreased with increasing inbound velocity while maximum deformation increased. The results also show temperature to have an effect on golf ball dynamics, although there is a relatively large amount of scatter in the data. Increasing temperature increased COR, impact duration and maximum deformation. Fig. 1 shows the FE model predicted COR and maximum deformation reasonably well, while impact duration was under predicted. Step-wise multiple linear regression analysis suggests that both inbound velocity and temperature had a significant effect on COR, impact duration and maximum deformation (p<0.001 in all cases). The regression equations were: Adjusted R2 = 0.928 (2) Adjusted R2 = 0.840 (3) Adjusted R2 = 0.785 (4) where Vin is inbound velocity in m\u00b7s-1, T is temperature in \u00b0C, ID is impact duration in s and DEFmax is maximum deformation in mm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001836_s40598-017-0071-0-Figure11-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001836_s40598-017-0071-0-Figure11-1.png", + "caption": "Fig. 11 The pullback map \u03c3g", + "texts": [ + " This star has at most two triangles and at most four vertices. Since \u03c3g and \u03bcg are conjugate on Q via p/q \u2192 \u2212q/p, this proves statement 1. Now suppose that g is not combinatorially equivalent to a rational map. So g has an obstruction. The pullback map \u03c3g fixes the negative reciprocal of the slope of this obstruction. We find it convenient for this fixed point to be\u221e. So we replace the map f two paragraphs above by a conjugate so that the new pullback map is \u03c3 f conjugated by z \u2192 \u22121/z. Arguing as two paragraphs above, we find that Fig. 11 describes \u03c3g in the same way that Fig. 10 describes \u03c3 f . The n in Fig. 11 is an integer. Any integer is possible. The case n = 0 is the case in which g is conjugate to f . Arguing as in the case in which g is unobstructed, we find that every element of Q eventually enters the star in Tg of\u221e under the iterates of \u03c3g . Hence it only remains to determine the action of \u03c3g on integers. We have that \u03c3g(\u22121) = n. Using the reflection principle, we see that \u03c3g(m) = m + n + 1 for every odd integer m. Similarly, \u03c3g(m) = m + n + 1+ i for every even integerm. Furthermore, the stabilizer of\u221e in the subgroup of modular group liftables for g has a generator which acts on H as z \u2192 z + 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002009_icmsao.2017.7934926-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002009_icmsao.2017.7934926-Figure6-1.png", + "caption": "Fig. 6. Schematic of the sampling lines.", + "texts": [ + " Global mesh refinement is used to improve the calculation precision. For more obvious result, the thickness of the substrate is chosen as 6 mm, and large welding simulation parameters are used, which are the same as the experiments. The cooling time is set after each layer deposition to reduce the heat accumulation. IV. RESULTS AND ANALYSIS The simulation parameters were the same ones except the clamping forms. To compare and analyze the results, four sampling lines were chosen in the longitudinal and transversal direction, and Fig. 6 shows their distribution. Sampling line 1 and 2 are the central lines while line 3 and 4 are the offset lines. Transversal sampling line 2 and 4 contain the sampling points in each layer. Total distortion and residual stress had been investigated to analyze each clamping form. Total distortion is an essential evaluation index of WAAM. It makes a strong impact on the dimensional precision. Fig. 7 plots the distortion of the substrate in four clamping forms after cooling. The results show that the effect of clamping form 1 is the worst in all sampling lines, and the distortion can reach nearly 6 mm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000760_imece2011-62445-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000760_imece2011-62445-Figure1-1.png", + "caption": "FIGURE 1. SCHEMATIC OF THE EBF3 PROCESS USING SOLID WIRE FEED.", + "texts": [ + " In addition, as part of the NASA exploration mission, the researchers at NASA LaRC have been conducting research to show the usefulness of the EBF3 process in space applications. In the future, NASA is planning to establish bases on the moon, which will require methods to minimize cargo mass and utilize recycling methods. The EBF3 process has the potential for use on the moon to re-fabricate broken metal parts and recycle scrap metal parts [8]. EBF3 is a flexible process with many different control variables, including beam accelerating voltage, beam current, beam spot size and shape, translation speed, wire diameter, and wire feed rate (Fig. 1). These variables determine the thermal history of the deposited components, which in turn determine the microstructure and final mechanical properties. In this work, we investigated the feasibility of producing single wall samples with AISI 316L stainless steel metal wire using the EBF3 process. The objective was to study the role of the beam current, wire feed rate, and translation speed on the build-up geometry, microstructural characteristics, and microhardness of the samples. In the present study, the substrate material used for deposition was a commercially available, 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003466_j.mechmachtheory.2019.06.017-Figure19-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003466_j.mechmachtheory.2019.06.017-Figure19-1.png", + "caption": "Fig. 19. Schematic of the feeding motion with worm plunging.", + "texts": [ + " The angular velocity ratio of the worm and the tapered involute pinion during the generating process can be expressed as follows: i wp = \u03c9 w \u03c9 p = N p N w (42) where \u03c9 p and \u03c9 w are angular velocities of the pinion and the worm, respectively. For a helical tapered involute pinion, an additional rotation \u03d5p of the pinion should be performed along with transla- tional distance s w and the shifting distance of the apex s 0 . Here, \u03d5 p = ( s w \u2212 s 0 ) \u00b7 tan \u03b2 r pp (43) Longitudinal crowning of the tapered involute pinion is achieved by plunging of the worm, which changes the generating path of the feed motion from u - O u - u to v - O v - v ( Fig. 19 ). The shortest distance between the axis z p and the axis z w is constantly changing during the generation. The pinion surfaces p is obtained as the envelope set of the worm surfaces w , and can be expressed by: \u23a7 \u23a8 \u23a9 r p ( u p , l p , \u03d5 w , s w ) = M pe M eu M u v M v t ( s w ) M tc M cw ( \u03d5 w ) r w ( u p , l p ) f 1 wp ( u p , l p , \u03d5 w , s w ) = n w \u00b7 v ( w p , s w ) w = 0 f 2 wp ( u p , l p , \u03d5 w , s w ) = n w \u00b7 v ( w p , \u03d5 w ) w = 0 (44) M cw ( \u03d5 w ) = \u23a1 \u23a2 \u23a3 cos \u03d5 w sin \u03d5 w 0 0 \u2212 sin \u03d5 w cos \u03d5 w 0 0 0 0 1 0 0 0 0 1 \u23a4 \u23a5 \u23a6 (45) M tc = \u23a1 \u23a2 \u23a3 1 0 0 0 0 cos \u03b3wr sin \u03b3wr 0 0 \u2212 sin \u03b3wr cos \u03b3wr 0 0 0 0 1 \u23a4 \u23a5 \u23a6 (46) M v t ( s w ) = \u23a1 \u23a2 \u23a3 1 0 0 \u2212 E 0 1 0 0 0 0 1 s w 0 0 0 1 \u23a4 \u23a5 \u23a6 (47) M u v = \u23a1 \u23a2 \u23a3 1 0 0 0 0 1 0 0 0 0 1 \u2212 s 0 0 0 0 1 \u23a4 \u23a5 \u23a6 (48) M eu = \u23a1 \u23a2 \u23a3 0 1 0 0 \u2212 cos \u03b4 0 sin \u03b4 \u2212E 0 sin \u03b4 0 cos \u03b4 r pw sin \u03b4 0 0 0 1 \u23a4 \u23a5 \u23a6 (49) M pe = \u23a1 \u23a2 \u23a3 cos \u03d5 p sin \u03d5 p 0 0 \u2212 sin \u03d5 p cos \u03d5 p 0 0 0 0 1 0 0 0 0 1 \u23a4 \u23a5 \u23a6 (50) Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002007_kem.736.91-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002007_kem.736.91-Figure1-1.png", + "caption": "Figure 1. Multi-powder system", + "texts": [ + " (3) where \u2206P \u2013 the pressure difference between the area to be processed and the rotor inner volume; \u03c1 \u2013 powder density; g \u2013 the free-fall acceleration; \u2206R \u2013 the rotor wall thickness; R \u2013 the rotor tube radius. The main specifications of the developed multi-powder system allow producing heterogeneous powder structures of two different metals or alloys. The thickness of the forming layer ranges from 40 to 200 microns and the building field is 500x420mm. General view of the multi-powder system is shown in Figure 1. The window in the body (1) is used for scanning the surface by laser beam. The carriages (5, 6) on linear guide (4) are moving by drives (2, 3). Both carriages have rotor drives (7). The carriages (5, 6) are the most important part of the multi-powder system. They are bunker feeding devices with collecting rotor, which is described above. Carriage view is shown in Figure 2. Powder is fed from the bunker (2). Rotor (1) removes the \u201cunattached\u201d powder, it is moving by using sheave (5) and tension roller (6)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002839_j.mechmachtheory.2018.06.021-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002839_j.mechmachtheory.2018.06.021-Figure4-1.png", + "caption": "Fig. 4. Structural schematic of half-toroidal CVT.", + "texts": [ + " Thus, sin (\u03b8 + \u03b3 ) cos \u03d5 = cos (\u03b8 + \u03b3 ) sin \u03d5 (4) tan (\u03b8 + \u03b3 ) = tan \u03d5 (5) Furthermore, \u03b8 + \u03b3 = \u03d5 (\u03b8 + \u03b3 ) \u2208 [ \u2212\u03c0 2 , \u03c0 2 ] , \u03d5 \u2208 [ \u2212\u03c0 2 , \u03c0 2 ] (6) By substituting Eq. (6) into Eq. (2) , we have { x = \u2212 (D + r) cos \u03d5 y = \u2212 (D + r) sin \u03d5 + y \u2217 (7) The last step is designing the shapes of input and output discs. Eq. (7) represents a circle with the centre at (0, y \u2217) and radius of D + r , as shown in Fig. 3 (b). By using this circle as the generatrix of the disc, we can obtain a traditional half-toroidal CVT, as shown in Fig. 4 . In Eqs. (2) and (7) , let y \u2217 be equal to zero to obtain the full-toroidal CVT. The properties of the toroidal CVTs have been well studied, and thus, repeating the conclusions is unnecessary. The toroidal CVTs are designated as the comparative object in this study. A rolling cone with its apex on the rotation axis of the disc is considered in this study. In this section, we present the calculation of an envelope of a family of straight lines and propose a new type of zero-spin CVT referred as inner cone CVT", + " 10 ) can be expressed as \u03c9 spin = \u03c9 dg sin \u03b8dg \u2212\u03c9 dn sin \u03b8dn , (43) where \u03c9 spin is the spin speed; \u03c9 dg and \u03c9 dn are the rotational speeds of the driving and driven components, respectively; \u03b8dg is the angle between the contact area and the rotational axis of the driving component; and \u03b8dn is the angle between the contact area and the rotational axis of the driven component. As such, \u03c9 dg sin \u03b8dg and \u03c9 dn sin \u03b8dn represent the vertical velocity components of driving and driven rotational speeds, respectively. Spin ratio is the ratio of the spin speed to the rotational speed of the driving component: \u03c3spin = \u03c9 spin \u03c9 dg = sin \u03b8dg \u2212 \u03c9 dn \u03c9 dg sin \u03b8dn = \u03c9 spin \u03c9 dg = sin \u03b8dg \u2212 sin \u03b8dn i , (44) where i is the speed ratio and equal to \u03c9 dg / \u03c9 dn . As shown in Fig. 4 , the spin speed of the half-toroidal CVT between the input disc and the roller is \u03c9 spinT = \u03c9 i sin ( \u03b8 \u2212 \u03b3 ) \u2212 \u03c9 r cos \u03b8, (45) where \u03c9 spinT is the spin speed of the half-toroidal CVT between the input disc and the roller, \u03c9 i and \u03c9 r are the rotational speeds of the input disc and the roller, respectively; and \u03b8 and \u03b3 are the half-cone and the tilting angles of roller, respectively. We assume a no-slip condition in the traction between the input disc and the roller, such that \u03c9 i [ ( e + R ) \u2212 R cos ( \u03b8 \u2212 \u03b3 ) ] = \u03c9 r R sin \u03b8, (46) where R is the cavity radius and e is the distance of the toroidal cavity from the disc rotation axis (see Fig. 4 ). The aspect ratio is k = e / R . By substituting Eqs. (45) and (46) into Eq. (44) , we derive \u03c3spinT = sin ( \u03b8 \u2212 \u03b3 ) \u2212 1 + k \u2212 cos ( \u03b8 \u2212 \u03b3 ) tan \u03b8 , (47) where \u03c3 spinT is the spin ratio of the half-toroidal CVT between the input disc and the roller. In accordance with Fig. 11 , the spin speed of the logarithmic CVT between the input disc and the roller can be derived as \u03c9 spinL = \u03c9 i cos ( \u03b8 \u2212 \u03b3 ) \u2212 \u03c9 r cos \u03b8 (48) where \u03c9 spinL is the spin speed of the logarithmic CVT between the input disc and the roller", + " Nikas [38] developed a numerical model to study the fatigue life of the full-toroidal CVT at fixed speed ratio, and the study found that roller life is always longer than disc life. However, in the case of variable speed ratio, Lee et al. [39,40] observed that the rollers always spall from the surface crack before the other components are damaged, which they attributed to the rollers that constantly accumulate contact fatigue damage. In contrast, the discs undergo discrete damage at each speed ratio because the positions of the contact points relative to the rollers and discs ( Fig. 4 ) are fixed and unfixed, respectively. In Fig. 11 , the contact ranges of the logarithmic CVT components are same as those of the half-toroidal CVT. On this basis, the durability performance of the logarithmic CVT has not been improved relative to the half-toroidal CVT. In Fig. 6 , the contact positions relative to the inner cone roller and disc both vary with the speed ratio. Thus, all regions of the working surfaces of the inner cone roller and the disc may experience contact. A similar conclusion can be derived with mathematical equations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003380_1.4043844-Figure14-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003380_1.4043844-Figure14-1.png", + "caption": "Figure 14. Plastic equivalent strain (\ud835\udf00\ud835\udc52 \ud835\udc5d ) in stub-end after application of torque Tb and internal pressure Pi at outer temperature To = 60 C, with corresponding safety factors SFL and SFY.", + "texts": [ + "org/about-asme/terms-of-use -19- \ud835\udc46\ud835\udc39\ud835\udc3f = \ud835\udf0e\ud835\udc50 \ud835\udc4e\ud835\udc63\ud835\udc52 \ud835\udf0e\ud835\udc35\ud835\udc48\u2044 at 40 C (a) \ud835\udc46\ud835\udc39\ud835\udc3f = \ud835\udf0e\ud835\udc50 \ud835\udc4e\ud835\udc63\ud835\udc52 \ud835\udf0e\ud835\udc35\ud835\udc48\u2044 at 60 C (b) (a) (b) (c) Figure 12. FEA results showing gasket stress \ud835\udf0e\ud835\udc50 \ud835\udc4e\ud835\udc63\ud835\udc52 vs. Tb and Pi at: (a) 40 C, (b) 60 C and (c) 80 C. Acc ep te d Man us cr ip t N ot C op ye di te d Downloaded From: https://pressurevesseltech.asmedigitalcollection.asme.org on 06/11/2019 Terms of Use: http://www.asme.org/about-asme/terms-of-use -20- \ud835\udc46\ud835\udc39\ud835\udc3f = \ud835\udf0e\ud835\udc50 \ud835\udc4e\ud835\udc63\ud835\udc52 \ud835\udf0e\ud835\udc35\ud835\udc48\u2044 at 80 C (c) Figure 13. Leakage assessment charts: (a) 40 C, (b) 60 C and (c) 80 C. In Figure 14 the deformed shape of the manhole assembly is shown at operational conditions , at various torque levels Tb and internal pressure levels Pi at 60 C. The contour FEA results pertain to the plastic equivalent strain, where \ud835\udf00\ud835\udc52 \ud835\udc5d > 0 indicates plastic deformation due to yielding is prevalent. One can see for low pressure and torque levels as in Figure 14(a) with Tb = 500 Nm and Pi = 4 Bar, no plastic deformation (e.g. \ud835\udf00\ud835\udc52 \ud835\udc5d = 0) is present and the safety factor against yielding SFY = 5.09. However, leakage will take place as SFL = 0.16. If the torque level is kept low (e.g. Tb = 500 Nm) and the pressure is further increased, i.e. Pi = 8 Bar as in Figure 14(b), both leakage and yielding takes place (SFY = 0.29 and SFL = 0.04), which is also evident by the deformed shape of and the widespread plastic strain in the stub-end. Hence, this rotation or bending deformation mode of the stub-end collar, which we here will denote cupping, is caused mainly by an increased internal pressure at lower torque levels leading to leakage and prevalent plastic deformation at the neck and collar of the stub-end. On the other hand, when the torque level is increased in order to prevent leakage as shown in Figure 14(c)-(d) with Tb = 3000 Nm, significant plastic deformation develops in the collar of the stub-end, which leads to a deformation mode we here will denote step formation, which is apparent between the edge of the ring flange and the collar of the stub-end. Hence, step formation is primarily caused by excessive torque levels. Acc ep te d Man us cr ip t N ot C op ye di te d Downloaded From: https://pressurevesseltech.asmedigitalcollection.asme.org on 06/11/2019 Terms of Use: http://www.asme.org/about-asme/terms-of-use -21- Downloaded From: https://pressurevesseltech", + " The safety factor is defined as the ratio between the yield strength of the HDPE material at the given temperature and the von Mises effective stress in the neck region of the stub-end obtained from the FEA results (i.e. \ud835\udc46\ud835\udc39\ud835\udc4c = \ud835\udc46\ud835\udc66/\ud835\udf0e\ud835\udc63\ud835\udc40). Similarily, SFY is presented in so called yielding assessment charts in Figure 16 at various temperatures. As evident from Figure 16 the yielding assessment charts, the trend in non-satisfactory safety factor (SFY < 1) is of two-fold with respect to Tb and Pi. One at low Tb and pressure Pi levels which leads to i.e. cupping and conforming to the mode of deformation in Figure 14(b). The other is at high Tb leading to step formation and conforming to the mode of deformation in Figure 14(c)-(d). In addition, the temperature has a significant effect on SFY, as evident from Figure 15 and Figure 16. As the temperature increases, SFY decreases. Thus step formation and the cupping failure mode become prominent at rather low Tb and Pi levels when temperature is increased (i.e. Figure 16(c)). (a) (b) (c) Figure 15. Yielding safety factor in stub-end SFY vs. Tb and Pi at: (a) 40 C, (b) 60 C and (c) 80 C. Acc ep te d Man us cr ip t N ot C op ye di te d Downloaded From: https://pressurevesseltech" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003886_s40684-019-00170-w-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003886_s40684-019-00170-w-Figure7-1.png", + "caption": "Fig. 7 Parallel gear dynamic model", + "texts": [ + " The rotation speed of a typical wind turbine is of the order of a few revolutions per minute to a few hundred revolutions per minute depending on its dimensions, whereas the optimum speed of a conventional generator is between 800 and 3600 revolutions/min [10]. As a result, a multiplier is usually required to accommodate the two rotational speeds. It is a component integrated in the kinematic chain of a wind turbine as shown in Fig.\u00a05. Figure\u00a06 presents the kinematic scheme of a gear train used to provide a ratio of 60.5 in a wind turbine with a rated power of PN = 750 KW. Figure\u00a07 illustrates dynamic model of parallel gear involving damping and meshing coefficients. Gearbox subsystem is composed of: \u2022 Planetary gear connected to the output rotor shaft; \u2022 Parallel gear connected to the output planetary gear shaft; Figures\u00a08 and 9 show BG detailed model of the studied transmission. This model is characterized by the addition of passive elements R where Sanchez and Al. [18] considered only the energy storage elements C. In order to validate the above model, the scheme of Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000737_s12206-012-0825-5-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000737_s12206-012-0825-5-Figure1-1.png", + "caption": "Fig. 1. Geometrical schematic of a hydrodynamic journal bearing.", + "texts": [ + " In section two, the optimized method NSGA-II is introduced in detail and a case analysis is made in section three. In section four, an experiment is made taking the optimized bearings as objectives to verify those numeric results given in the paper are right. At last, a conclusion is drawn. Assuming the full gap is filled with the lubricating liquid and the shape is constant in axial direction (Z axial), the bearing shape can be studied in a two-dimensional polar coordinate system, as shown in Fig. 1. The bearing is fixed and the journal rotates with an angular speed \u03c9 clockwise about Z axial. The journal center is taken as the origin to establish a polar coordinate system, where r is the journal radius, ( )\u03c1 \u03b8 is the polar radius of the bearing profile and ( )h \u03b8 is a gap function between the bearing and the journal, which also is called the film thickness at angle \u03b8 . Their relation is: ( ) ( )r h\u03c1 \u03b8 \u03b8= + . (1) From the beginning of design, it is difficult to decide the bearing shape in advance" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000740_j.tcs.2012.04.038-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000740_j.tcs.2012.04.038-Figure1-1.png", + "caption": "Fig. 1. Observation error.", + "texts": [ + " Angle error Any observed angle has an additive error within \u00b1\u03b80. That is, letting angle formed by v and V be \u03b8 , cos \u03b8 \u2265 cos \u03b80 is satisfied.2 2 Note that \u03b8 always takes the smaller angle from the two formed by v and V because of cos \u03b8 \u2265 cos \u03b80 . We explain the influence that these parameters give by showing an example: consider a situation where a robot s1 observes s2. Let V be a vector representing the location of robot s2 in terms of s1\u2019s coordinate system, and v be the vector representing the location of s2 in s1\u2019s observation result. (See Fig. 1.) Then, the grey area is the possible observation area. That is the coordinate v is necessarily contained in the area at s1\u2019s observation result. In this subsection, we define the non-uniform error model and the uniform error model. Non-uniform error model [1] If a robot observes other two or more robots, the observation result can involve different distance/angle errors for each observed robot. Fig. 2.a shows an example. In this example, robot s0 observes all other robots, and the observation error occurs differently for two robots s1 and s2 : Distance error ratio \u03b51 and angle error \u03b81 is associated with robot s1 and \u03b52 and \u03b82 with robot s2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000546_j.1460-2695.2010.01540.x-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000546_j.1460-2695.2010.01540.x-Figure2-1.png", + "caption": "Fig. 2 Spur gear BE model, schematized representation of crack modelling with the FSM and SIF calculation with QPE.", + "texts": [ + " The factorial design requires 36 crack propagation simulations. Calculation results are a series of points defining the crack propagation path (x and y coordinates) with associated K I and K II values B E N D I N G F A T I G U E B O U N D A R Y E L E M E N T M O D E L Gear boundary element model Numerical simulations are performed with an in-house BE program offering: reduced discretization order and accurate modelling of high stress gradients.20 In all cases, the simulation involved a seven-tooth gear representation (Fig. 2). Model accuracy is ensured with an adequate mesh refinement of the cracked region using quadratic BEs under a plain strain assumption. The unit load W is applied at the highest point of single tooth contact of the cracked tooth. This has been proven to be the position controlling the entire bending fatigue propagation process.21 It is well known that the BE formulation mathematically degenerates in the presence of co-planar crack surfaces.22 Usually, this is circumvented with mixed-formulation or multi-domain approaches", + "20 However, a new simple BE crack modelling technique called the Finite Separation Method (FSM) has been recently proposed,23 and prevents BEM degeneration by imposing a negligible gap (\u03b4) between crack faces. This approach limits the model to a single domain, consequently to the standard displacement BEM formulation, thus leading to simpler BE representation and reduced computation times.23 The FSM is used in this work for crack modelling. The crack face separation \u03b4 is fixed at 1 \u00d7 10\u22122 \u03bcm. Accurate solutions are obtained with a gradually refined mesh imposed at each new crack segment (da), with hs = da/10 and ht = da/20, while previous segments are uniformly discretized23 (Fig. 2). The crack tip 1/ \u221a r stress singularity is numerically introduced with the well-known QuarterPoint Elements (QPE) (Fig. 2). SIF K I and K II are evaluated using the displacement extrapolation method. With the 1/ \u221a r singularity selfcontained in the QPE, SIF are calculated using only a twopoint extrapolation (Eq. (1)), where variables are defined according to Fig. 2. Under plain strain conditions, \u03bc = G (shear modulus) and \u03ba = 3\u20134\u03c5 . KI = \u03bc \u03ba + 1 \u221a 2\u03c0 ht [ 4 ( uB 2 \u2212 uD 2 ) + uE 2 \u2212 uC 2 ] KII = \u03bc \u03ba + 1 \u221a 2\u03c0 ht [ 4 ( uB 1 \u2212 uD 1 ) + uE 1 \u2212 uC 1 ] . (1) The crack propagation direction (\u03b8 ) is determined with the MTS criterion16 Eq. (2). \u03b8\u03c3\u03b8\u03b8 max = 2 tan\u22121 \u23a1 \u23a31 4 \u239b \u239d KI KII \u00b1 \u221a( KI KII )2 + 8 \u239e \u23a0 \u23a4 \u23a6 . (2) In order to ensure smooth crack trajectories, propagation steps (da) are fixed to 0.5 mm. Simulations end at complete fracture or, from a numerical point of view, when the propagation point falls outside the geometry" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000239_sii.2010.5708334-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000239_sii.2010.5708334-Figure6-1.png", + "caption": "Fig. 6. Simulation of vertical to horizontal flight transition.", + "texts": [], + "surrounding_texts": [ + "We study the effectiveness for energy consumption of the fixed-wing with simulator. We simulate hovering flight, level flight and transitional flight. The UAV hovers for 10 seconds, and it transfers from hovering flight to level flight. Simulation is made up of two steps. In the first step, the UAV hovers for 10 seconds while maintaining initial attitude and altitude. The initial condition is that the roll angle is 0 [\u25e6], the pitch angle is 0 [\u25e6], the yaw angle is 0 [\u25e6], the thrust is 0 [N], the flight speed is 0 [m/s] and the altitude is 0 [m]. In the second step, the UAV switches from hovering flight to level flight in 10 seconds. Reference states are that the angle of attack is 15 [\u25e6], the roll angle is 0 [\u25e6], the yaw angle is 0 [\u25e6] and the altitude is 0 [m]" + ] + }, + { + "image_filename": "designv11_33_0000945_ssd.2010.5585533-Figure14-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000945_ssd.2010.5585533-Figure14-1.png", + "caption": "Figure 14. 3-D flux path trough the north claw and the stator lamination.", + "texts": [ + " Figure 11 shows a mesh of the rotor claws and the as sociated magnetic rings. 3.2.1. Main Flux Paths The flux moves through the magnetic circuit of the CPAES within three dimensional (3D) paths. These are described as follows: \u2022 axially in the stator yoke (see figure 12), \u2022 radially down through the collector and crossing the air gap down to the magnetic ring (see figure 12), \u2022 radially then axially in the magnetic ring (see figure 12), \u2022 axially-radially in the claws (see figure 13), \u2022 radially up in air gap facing the stator laminations (see figure 14), \u2022 radially up in the stator teeth and circumferentially in the stator core back (see figure 15), \u2022 radially down in stator teeth facing the two adjacent claws (see figure 16), \u2022 radially-axially in the adjacent claws (see figure 16), \u2022 axially then radially in the magnetic ring (see figure 17), \u2022 radially up through the air gap and the mag netic collector on the other side of the machine (see figure 17), \u2022 axially in the other side of the stator yoke (see figure 18). 3.2.2. Leakage Flux Paths Not all the flux produced by the excitation winding and the armature contributes to the EMF generation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002264_1.5008043-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002264_1.5008043-Figure1-1.png", + "caption": "FIGURE 1 Example scanning path in gauge length of sample", + "texts": [ + ", medical devices and power plants) thanks to its outstanding corrosion resistance at low temperatures, and high oxidation resistance and superior creep-fatigue resistance at high temperatures. Cylindrical test pieces, with a parallel length of 18mm, a gauge diameter of 2mm and with M6 threaded end-grips, were printed by power-bed fusion printer (Concept Laser Mlab cusing R) in the vertical (build) direction. The powder size distribution is from 10micro to 45micron. On each layer, a bi-directional hatch pattern (see FIGURE 1) was used to melt powder onto existing solidified layer. The bi-directional hatch pattern was rotated 90\u00b0 compared to that of the previous layer. The machine parameters (provided by the machine manufacturer named CL20 \u201cperformance\u201d) for the pattern were as follows: 90W laser power, 600mm/s scan speed, 84\u00b5m hatch spacing, 50\u00b5m spot size and 25\u00b5m layer thickness. In addition to hatching two contour scans were performed around circumference of the cylindrical samples with a reduced laser power of 60W", + " The plastic anisotropy is one of main barriers to the application of advanced metals and alloys, [6]. Obviously, the observed columnar grains and texture can be strongly influential to the deformation response and failure behaviour of 040017-3 the printed steel, in particular during multi-axial deformation. Therefore, it would be very interesting to study the influence of texture and grain morphology on the plastic anisotropy during multi-axial deformation of AM metals. FIGURE 4b shows the cross-section of microstructures in melt pool. Because of the scanning strategy (FIGURE 1, the scanning pattern of a new layer was rotated 90\u00b0 compared to the last previous layer), a checkerboard pattern is seen in FIGURE 4b. High magnified observations show that the checkerboard pattern consists of relatively big grains (10-50\u00b5m in diameter) surrounded by elongated fine grains (5-10\u00b5m in width), FIGURE 4c. Very fine equiaxed grains (1-5\u00b5m) are seen in the intersection between four domains of larger grains. Combining the observations in BD (FIGURE 4a) and cross sections (FIGURE 4b and c), most grains aligned in column have a 3D spherical bowl structure with varied thickness: bowls get thicker towards their bottom" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000228_b905459f-Figure11-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000228_b905459f-Figure11-1.png", + "caption": "Fig. 11 Schematic plot of cross-sections assumed in the filament tension models, labelled with circumferences. The circumference of the corrugated filament can be estimated as p/2 times the circumference of the cylinder. The circumference of filaments with other shapes may be larger or smaller than derived from the apparent diameter and an assumed cylinder shape.", + "texts": [ + " 8) and the mass density as r\u00bc 1 g cm 3, then one obtains the solid curves c f d 1/2 in Fig. 7 and Fig. 8. It is evident that the agreement with the experimental data is far from being satisfactory. All data except of one single datum point are far below the prediction. This discrepancy has been mentioned earlier,9 on the basis of relaxation data after electrical excitation. The restoring force must be larger than assumed in the surface tension model. If we take into account that the filament surface is actually corrugated (Fig. 11, middle), then the surface may be larger by a factor of about p/2, and c increases by ffiffiffiffiffiffiffiffi p=2 p (dashed lines in Fig. 7 and Fig. 8). Thin filaments which have been checked to be cylindrical in good approximation are in reasonable quantitative agreement with the corrugated surface model predictions. Cylindrical filament shapes could be achieved only with radii of 50 mm and less. For the larger filaments, we find systematically too high resonance frequencies f. The comparison of electric and acoustic excitation of oscillations of the same filament showed in all cases very good agreement between the two methods" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002934_s10846-018-0935-0-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002934_s10846-018-0935-0-Figure2-1.png", + "caption": "Fig. 2 The leg structure of TMUBOT", + "texts": [ + " Also, the process of designing TESO bandwidth is investigated. To have a proper assessment of the effectiveness of the proposed theory, TMUBOT robot has been used as a real platform (see Fig. 1). Simulation and implementation results of the proposed algorithm on TMUBOT robot have been presented through some real scenarios. TMUBOT has been designed and built in the intelligent control systems laboratory (ICSLAB) at Tarbiat Modares University and it is a quadruped robot with twelve degrees of freedom (DOF) with each leg of the robot having three DOFs (see Fig. 2). Table 1 shows some physical and mechanical features of the system [31]. Since the considered system is a MIMO system, decentralized structure of the system has been derived and used therefore, the system dynamic equation has converted to SISO sub-systems. Finally, for each sub-system, a TESO has been designed in order to estimate and compensate internal disturbance caused the couplings between the subsystems and also, external disturbance due to the collision of the robot\u2019s foot to the ground and the friction force" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001093_icra.2012.6225142-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001093_icra.2012.6225142-Figure4-1.png", + "caption": "Fig. 4. Overview of the First Prototype Model of Planar Version", + "texts": [], + "surrounding_texts": [ + "In this section, we describe the configuration of the planar version of the omnidirectional driving gear mechanism. The actuators that drive the mechanism are placed in world coordinates so the plate that moves omnidirectionally on the flat plane can be configured in a lightweight manner." + ] + }, + { + "image_filename": "designv11_33_0000263_robio.2010.5723411-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000263_robio.2010.5723411-Figure2-1.png", + "caption": "Fig. 2. Generated paths for hip and swing ankle joints in one step of walking", + "texts": [ + " 1 (5) To eliminate Lagrange multipliers one may pre multiply Eq. 4 by transpose of orthogonal compliment of *; i.e. matrix *2 which satisfies **2 3. As *2 is a 4 matrix, the resulting equations represent a set of six new equations with eight unknown of 5 . In order to complete these equations one should append second time derivative of constraint equations, Eq. (3), to obtain where 6, 8 and : are as follows: * > : ;*2+3 < (7) In this section, two paths are specified for hip and swing ankle joints (Figure 2). To have smooth motion for hip joint, a trigonometric function is used as follow [5]: ?@ A % BCD % BEFG % BGFH I@ JCD K % LMN \"DOP A(Q BGR BCR % JC % AS (8) where Ls is one step length, T is one step period, Lfot is foot length, Ltoe is toe length, Lth is thigh length, Lsh is shank length, hs is vertical width of hip path curve and tl is a tolerance inserted in yH equation to prevent the model from singularity condition in knee joint. In this research torso in Figure 1 remains vertical during walking", + " 8 can be used as path of the center of mass with a vertical shift. Trajectory of swing ankle is considered to have two sections related to double support and single support phases. During DSP toe angle value (q1) begins from an initial value and linearly up to zero at the end of this phase. Then the ankle tracks a circular arc with a variable radius: ?TU %BC % BT LMN VT ITU BT NWX VT (9) where subscript \u2018d\u2019 refers to DSP, La and YZ are respectively the variable radius and its angle with the horizon, shown in Fig. 2. for state that toe-joint doesn\u2019t bend the path will be a circle with radius La=Lfot+Ltoe. During SSP a cubic polynomial in following form is used to define ankle path: ITC [ ?TC [ ?TC [ ?TC [\\ (10) In this relation subscript \u2018s\u2019 refers to SSP. Coefficients a3-a0 should be obtained from continuity condition of the curve and its derivative also the path should satisfy the condition of yas and its derivative being zero at the end of SSP. By using these paths and inverse kinematic, angles trajectory of the model can be obtained in one step of walking", + " Consumed energy as a function of time can be calculated by integrating Eq. 11 to get: b^ A c _`a a_deG \\ ) f A f P (12) where d is time differential variable. Energy consumption index for joint \u2018i\u2019 defined as total consumed energy in the joint during one walking step and can be written as: gh^ c _`a a_dA \\ (13) Total energy index is sum of the energy consumption index in joints: gh i gja ak (14) In this section, two models are considered: model with toe bending, and a toe-less model. Motion paths of hip and swing ankle for these two systems are the same as Figure 2. Physical parameters of the model in this simulation are shown in table I. to make two models comparable the sevenlink model is obtained from a nine-link model whose toejoint angle are fixed during motion. Simulation is done for step length Ls=50 cm, and walking speed of 72 cm/s. total step period is T=0.7 s. Duration of DSP is 0.1 s and remaining 0.6 s is SSP time. Joints angle variations in one step obtained using inverse kinematic and are shown with robot motion for two models in Figure 3 and Figure 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001520_pes.2011.6039209-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001520_pes.2011.6039209-Figure1-1.png", + "caption": "Fig. 1. Schematic diagram of the three-mass equivalent model", + "texts": [ + " A WTG is a quite complex system where wind energy is converted into mechanical energy by the wind turbine and mechanical energy is converted into electrical energy by a generator. The drive train system (also called shaft system) which is between the wind turbine and induction generator, is composed of wind turbines blades, wheel hub, low-speed shaft, gearbox, high-speed shaft and generator rotor. The three-mass equivalent model of the shaft system, where the blades are presented as the first mass, wheel hub and low-speed shaft as the second mass, and the others as the third mass [14], is shown in Figure 1. D 978-1-4577-1002-5/11/$26.00 \u00a92011 IEEE The differential equations of the three-mass model are: ( ) ( )\u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23a9 \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23a8 \u23a7 \u2212= \u2212= \u2212\u2212= \u2212\u2212= \u2212\u2212= 12 12 2 2 1112 1 1 22122 2 2 2 2 2 2 tts t tgs gt ttmt t t tttgt t t gggtg g g dt d dt d DTT dt dH DTT dt dH DTT dt d H \u03c9\u03c9\u03c9\u03b8 \u03c9\u03c9\u03c9 \u03b8 \u03c9\u03c9 \u03c9\u03c9 \u03c9 \u03c9 (1) ( ) ( )\u23a9 \u23a8 \u23a7 \u2212+= \u2212+= 1212121212 22222 tttttt ggtgtgtgt DKT DKT \u03c9\u03c9\u03b8 \u03c9\u03c9\u03b8 (2) where 1tH , 2tH and gH are the inertia constants of the wind turbine blades, the wheel hub and the generator rotor, respectively; 1tD , 2tD and gD are the damping coefficients of the wind turbine blades, the wheel hub and the generator rotor, respectively; 1t\u03c9 , 2t\u03c9 and g\u03c9 are the angle speeds of the wind turbine blades, the wheel hub and the generator rotor, respectively, and fs \u03c0\u03c9 2= are the synchronous angle speed; 12tD , 12tK , 12t\u03b8 and 12tT are the damping coefficient, the shaft stiffness coefficient, the shaft twist angle and the shaft torque between the blades and the wheel hub, respectively; gtD 2 , gtK 2 , gt 2\u03b8 and gtT 2 are the damping coefficient, the shaft stiffness coefficient, the shaft twist angle and the shaft torque between the wheel hub and the generator rotor, respectively; mT is the mechanical torque of the wind turbine; gT is the electromagnetic torque of induction generator" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000933_s11249-010-9690-5-Figure14-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000933_s11249-010-9690-5-Figure14-1.png", + "caption": "Fig. 14 Micro-oven used for the calibration experiments", + "texts": [ + " Behind the contact area, the lubricant\u2019s thickness reduction corresponds to the lubricant meniscus break-down also clearly seen on the optical micrograph (Fig. 13). 4.3 Simultaneous Measurements of Pressure Film Thickness and Temperature in an EHL Contact As mentioned in Sect. 3.2, the evolution of ln(IAS/IS) as a function of the inverse of temperature has to be established experimentally to validate this approach and also to determine the ordinate at the origin, which can be affected by the spectrometer. For this purpose, a volume of 1 mm3 of 5P4E is introduced in a small oven which is temperature controlled with an accuracy of 0.5 K (Fig. 14). After stabiliseion of the temperature, the laser probe is focussed in the 5P4E sample, and an anti-Stokes and Stokes spectrum is recorded. The evolution of ln(IAS/IS) as a function of 1/T is presented in Fig. 15. The linear relationship which was theoretically predicted is effectively obtained. As previously deduced, the experimental slope -A = -14.3 102 \u00b1 0.4 102 K is equal to the theoretical one, whereas the B0 constant 0.0 \u00b1 0.2 is slightly different from the theoretical one (0.4) resulting from the influence of the spectrometer" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003587_1350650119868910-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003587_1350650119868910-Figure1-1.png", + "caption": "Figure 1. Schematics of bearing with compound texture: (a) three-dimensional schematic, and (b) two-dimensional schematic.", + "texts": [ + " In this research, the influences of typical operation parameters on the acoustic performances of the compound textured journal bearing are investigated through CFD method using the associated acoustic theory. In the investigation, the thermal effect of the lubricant is concluded, since the thermal effect can change the bearing\u2019s tribological performances. Further, an experimental verification is also carried out. Finally, the relevant conclusions are given. The three-dimensional schematic of the bearing with the inner radius R1 is shown in Figure 1(a). On the bearing inner surface, the compound textures with the starting angle \u2019s are uniformly fabricated. The number and interval angle of the texture are represented with N and , respectively. The shaft with the radius r rotates around its geometric center O1 at the rotational speed n, as seen from Figure 1(b). The shaft center deviates from the bearing center O2 at the applied load Fz. The center deviation is evaluated through the eccentricity distance e and attitude angle . The lubricant is filled in the clearance c between the shaft and bearing to balance the load Fz. The clearance can be computed according to the relation c \u00bc R1 r, whose largest value hmax \u00bc c\u00fe e and whose smallest value hmin \u00bc c e. The symbol \u2019 represents the circumferential angular position, which is calculated from the vertical position of zero angle along the circumferential direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002063_s00170-017-0656-8-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002063_s00170-017-0656-8-Figure7-1.png", + "caption": "Fig. 7 Cutting depth and thickness Fig. 9 The cutting depth calculation model", + "texts": [ + " Projecting the curve AB to the plane T, the equation of the curve A\u2217B\u2217can be expressed as x1 uGi; \u03b8Gi\u00f0 \u00de \u00bc l1 \u00fe m1z uGi; \u03b8Gi\u00f0 \u00de \u00fe n1z2 uGi; \u03b8Gi\u00f0 \u00de \u00f04\u00de The surface cone generatrix equation in coordinate system S2 can be written as x2 \u00bc za \u00fe tan \u03b4 f z \u00f05\u00de where za is the face apex beyond the crossing point, and \u03b4fis the face angle. The height difference of the lineMN and the curve A\u2217B\u2217can be expressed as \u0394hi \u00bc x2 zi\u00f0 \u00de\u2212x1 zi\u00f0 \u00de Li < zi < Lo\u00f0 \u00de \u00f06\u00de where Loand Liare the outer cone distance and inner cone distance, respectively. The cutting area, determined by the cutting depth and thickness, is a key factor influencing the cutting force. Therefore, to calculate the cutting force, the cutting depth and thickness should be deduced firstly. The illustration of the depth and thickness is shown in Fig. 7. In this paper, the cutting tool effective angle means that only within the scope of this angle can the cutting tool process the workpiece. Based on the case of processing right-hand gears using right-hand cutting tools, the cutting-in and cutting-out angle diagram is established in Fig. 8. \u03b81 and \u03b82 are the cutting-in angle and cutting-out angle; ro is the cutting tool radius; Sr and q are the original radial distance and cradle angle, respectively; Li, Lm, and Lo are the inner cone distance, mean cone distance, and outer cone distance, respectively; B2 is the tooth width; and \u03b21 and \u03b22 are the inner end spiral angle and outer end spiral angle, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000237_niss.2009.249-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000237_niss.2009.249-Figure1-1.png", + "caption": "Figure 1. Finding an empty point", + "texts": [ + " For a regular coverage hole, the literature [12] presented a calculation method of the number of empty points, whose sensing circles are inscribed by hexagonal cells. But it is complex and not distributed. In the following section, we propose a new energy-efficient coverage hole self-repair algorithm. In this section, we first describe the basic idea of our algorithm and then propose the self-detecting algorithm, called DSEPA, to efficiently detect empty points in coverage holes. We also present the repair scheme considering power saving in this section. In Fig. 1, four sensors A, B, C, and D are given and they intersect at points a, b, c and form orifices, ab and bc. Assume ab and bc belong to a boundary of a coverage hole. For orifice ab, if we draw two circles with radius r = Rs at intersection points a and b, the two auxiliary circles intersect at points d1 and d2. Between d1 and d2, we can find one point (d2) locates inside the coverage hole. If we pick d2 as an empty point, the new sensor can definitely cover orifice ab. We plot a new auxiliary circle at intersection point c, which intersects the first auxiliary circle (centered at point a) at point e" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000809_acc.2010.5531262-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000809_acc.2010.5531262-Figure2-1.png", + "caption": "Fig. 2. Force Diagram of the Magnetic Bearing for a Single Axis", + "texts": [ + " In this paper, the analog compensators are replaced by digital controller implemented by National Instruments LabVIEW Real-Time and/or Mathworks xPC Targets. Both systems use the processors in dedicated target computers for realizing digital control. In order to get an intuitive understanding of the Magnetic bearing system, a model is derived from rigid body dynamics [4]. For simplicity, assume that motion in the x (vertical) and y (horizontal) directions are identical and independent. As seen in Figure 2 the magnetic bearing applies forces F1 and F2 at x1 and x2 respectively. Positions x1 and x2 are measured to be a distance of l1 from each end of the shaft. The sensors measure at positions X1 and X2, which are at a distance of l2 from each end of the shaft. The differential equations governing the rigid body motion are obtained by summing the forces and moments around the 978-1-4244-7427-1/10/$26.00 \u00a92010 AACC 2206 center of mass located at x0 to yield \u2211 F = mx\u03080 = F1 + F2 (1) \u2211 M = I0\u03b8\u0308 (2) = F2 ( L 2 \u2212 l1 ) cos \u03b8 \u2212 F1 ( L 2 \u2212 l1 ) cos \u03b8(3) The electromagnetic force is given as: Fj = ke (ij + ib) 2 (xj \u2212 xg)2 \u2212 ke (ij \u2212 ib) 2 (xj + xg)2 (4) where ke = 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003110_j.cad.2019.01.001-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003110_j.cad.2019.01.001-Figure3-1.png", + "caption": "Fig. 3. Initial, intermediate and final mesh configurations for a 120 degrees mesh morphing rotation. (a) First stage; (b) second stage; (c) third stage; and (d) last stage.", + "texts": [ + " All the examples have been executed using only one processor. In all the examples we use the relative element quality, qe defined in terms of the elemental distortion defined in Eq. (9) qe = 1 \u03b7e . The relative element quality takes values between zero for inverted elements to one for ideal elements. In all the examples we define the mesh quality as the minimum element quality. In this example we show a two-dimensional case in which we apply a rotation of 120 degrees to the inner circles of the mesh in Fig. 3(a). The mesh is composed of 2466 linear triangles and 1318 nodes. We apply the proposed augmented Lagrangian formulation to perform the mesh morphing, and we show several stages of the deformation in Figs. 3(b)\u20133(d). The displacement of the inner circles in the intermediate stages is automatically obtained by the method. Note that this displacement is neither a linear homotopy between the initial and the final positions, nor an incremental rotation between zero and 120 degrees. Themethod automatically detects the number of stages to fulfill the boundary constraint" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002730_s13369-018-3419-4-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002730_s13369-018-3419-4-Figure2-1.png", + "caption": "Fig. 2 Coordinate systems", + "texts": [ + " The superposition between the theoretical tooth surface and the deviation surface can yield the precise topological modified tooth surface of the pinion, and the position and normal vectors are expressed as: r1(u1, l1) = \u03b4(h, l)nr(u1, l1) + rr(u1, l1) (4) n1(u1, l1) = \u2202r1(u1, l1) \u2202u1 \u00d7 \u2202r1(u1, l1) \u2202l1 / \u2223\u2223\u2223\u2223\u2202r1(u1, l1)\u2202u1 \u00d7 \u2202r1(u1, l1) \u2202l1 \u2223\u2223\u2223\u2223 (5) where nr(u1, l1) is the unit normal vector of the theoretical tooth surface. The coordinate systems adopted by TCA and LTCA of the face gear drive are shown in Fig. 2. The movable coordinate system defined by S1 and S2 is rigidly connected to the spur cylinder pinion and the face gear, and \u03a61 and \u03a62 are the corresponding rotation angles. Two surfaces, 1 and 2, are meshed in the fixed coordinate system Sf . The fixed coordi- nate systems Sq, Sd and Se simulate the axial displacement error q, the offset error E and the shaft angle error \u03b3 , respectively. Parameter B is the difference between the radius of the pinion and the virtual shaper, \u03b3 is the shaft angle, \u03b3m is the complementary angle of \u03b3, L0 is the average of the inner and outer diameters of the face gear, \u03b3f is the actual shaft angle, and \u03b3f = \u03b3m + \u03b3 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003852_j.mechmachtheory.2019.103628-Figure16-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003852_j.mechmachtheory.2019.103628-Figure16-1.png", + "caption": "Fig. 16. The tooth grinding simulation platform.", + "texts": [ + " The final parameters of the grinding wheel are used to check the interference of the non-working side of the grinding wheel again as shown in Fig. 15 . It can be seen that there is no interference, and the minimum gap for concave side grinding is 0.449 mm, the minimum gap for convex side grinding is 0.217 mm. Both locate at the top of the grinding wheel. The distance increases gradually from the top of grinding wheel to the top of tooth. According to the motion principle of CNC hypoid grinding machine shown in Fig. 7 , a five-axis tooth grinding simulation platform for the FFHHG is constructed, as shown in Fig. 16 . The distance R 0 is 350 mm, L B and L M are 102.538 mm and 65 mm, respectively. Then the grinding wheel location of tooth grinding can be calculated according to the method described in Section 5 , which are shown in Figs. 17 and 18 . At the same time, the grinding wheel model (the green part) is built according to Table 4 , and the imported gear model (the orange part) is shown in Fig. 10. There is no feeding during the first grinding, and the grinding wheel should be tangent to the tooth surface theoretically" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002941_icelmach.2018.8506824-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002941_icelmach.2018.8506824-Figure7-1.png", + "caption": "Fig. 7. Air-flow distribution in cooling channels and temperature distribution by the CFD model for (a) the average and (b) the actual measured air flow speed considered in every cooling channel.", + "texts": [ + " This is attributed to the placement of the terminal box on the left side of the prototype\u2019s housing, which incurs the partial blockage of specific ducts. In particular, the channels that correspond to duct -5 (Fig. 4) are partially blocked. As a result, the airflow to the three channels associated with this duct is restricted, leading to reduced heat dissipation and slightly increased local temperature values. The measured non-symmetric air-flow distribution between the channels of the left and right side of the motor is depicted in Fig. 7(b). In the thermal model the time averaged flow speed of each duct was used as a boundary condition. It can be observed that the air flow speed does not vary linearly with rotational speed. Additionally, it can be observed that for the same rotating speed, the air-flow speed values between the different ducts vary. The maximum variation is approximately 50% and 30% for the cases of 1000rpm and 4800rpm respectively. For the latter case, the maximum difference exhibits an absolute value of 13m/s. For the representation of the air flow speed in the cooling channels, two approaches are investigated. The flow speeds are imposed on each individual channel based either on the measurements (uneven flow speed) or based on an average value calculated from the measurements (even flow speed). In the even flow speed case all ducts are modeled to have the same air flow speed. Figure 7 illustrates the variation of airflow speed in cooling channels for the implemented 3D CFD model of the machine, when (a) even and (b) actual measured air flow speeds are imposed. For the thermal models only the rotation in one direction was simulated. As the model was set up, this direction of rotation coincided with what was defined as counterclockwise rotation during the measurements. A simplified radiation model on the outer stator surface is introduced in order to assess the radiative heat transfer effects" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002089_0954406217718857-Figure11-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002089_0954406217718857-Figure11-1.png", + "caption": "Figure 11. Gear shaft deformation under the reference pressure and the corresponding influence matrix.", + "texts": [ + " In this way, the bearing influence matrix IM \u00f0b\u00de i;j corresponding to the film mesh is obtained as shown in Figure 10, and the overall bearing surface elastic deformation yields h b\u00f0 \u00de s \u00bc Xmp i Xnp 1 j pi,j pref IM b\u00f0 \u00de i,j \u00f032\u00de (2) Gear shaft surface elastic deformation The gear shaft surface elastic deformation refers to the deformation of the bearing location, which has an influence on the film geometry. It is related to the film pressure on the bearing location, the fluid pressure around the gear circumference (TSVs and tooth tips) and the meshing force. As shown in Figure 11, displacement constraints are applied to the gear\u2019s lateral sides, and a reference pressure (10MPa) is applied on the divided face of the bearing location, the TSV, and the tooth tip. After that, the corresponding influence matrix can be obtained in the same way as the bearing influence matrix. The influence matrix with respect to the meshing force can also be obtained by applying a reference force (1000N) on the meshing point. Hence, the gear shaft surface elastic deformation yields h s\u00f0 \u00de s \u00bc Xmp i Xnp 1 j pi,j pref IM s\u00f0 \u00de i,j \u00fe Xz1 k ptsv pref IM tsv\u00f0 \u00de k \u00fe Xz1 k ptip pref IM tip\u00f0 \u00de k \u00fe Fmj j Fref IM mesh\u00f0 \u00de \u00f033\u00de and the total solid deformation yields hs \u00bc h b\u00f0 \u00de s \u00fe h s\u00f0 \u00de s \u00f034\u00de Presented in Figure 12 is the simulation procedure implemented in MATLAB, starting with the input parameters including the geometric parameters and the initial values" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003894_ecce.2019.8913269-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003894_ecce.2019.8913269-Figure3-1.png", + "caption": "Fig. 3: Cross-section of the investigated induction motor casestudy.", + "texts": [ + " It should be mentioned that in addition to the rotor bar material conductivity modification by the slip parameter, the resistance of the connection between adjacent bars is also treated in the same manner as the rotor bar materials, and should accordingly be modified by the same slip factor. The permeabilities calculated through the VBR analysis in the frequency domain solver are then imported into the TSFE model to \u201cpush\u201d the solution forward towards the desired steady-state convergence. IV. STATIC ECCENTRICITY The case-study 75-KW, 380-volt, 50-Hz, 4-pole, induction motor with a cross-sectional view shown in Fig. 3, was simulated using ANSYS software FE package of \u201cversion 19\u201d on a state-of-the-art workstation (Intel-Xeon E5-2687W v4 3.00-GHz 256-GB-RAM). This induction motor has 50 rotor bars and 60 stator slots and an air-gap length of 0.8 mm. The FE mesh of this motor includes a total number of 65,108 triangular elements. A static eccentricity fault with 75% severity (0.6 mm displacement of stator axis in the positive \u201cX axis\u201d direction in the FE grid) was applied. This static eccentricity was applied both to the eddy current frequency domain solver as well as the TSFE model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001045_jsea.2010.312134-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001045_jsea.2010.312134-Figure6-1.png", + "caption": "Figure 6. The assembly model.", + "texts": [ + " It can be concluded that the proposed hybrid methodology with BP and NSGA II can solve tolerance synthesis problem effectively. The FEA integration (Figure 8-11), helps in determining deformation due to inertia effects like gravity, velocity,acceleration, etc., resulting in decrease in the critical assembly feature. The CAD integration (Figure 5), helps in determining contribution of various tolerances towards the critical assembly feature. The assembly model of the motor assembly is shown in Figure 6. The exploded view of the motor assembly is shown in Figure 7. In this research, the proposed approach provides better formulation of cost-tolerance relationships for empirical data. BP network architecture of configuration 4-6-1 generates a suitable model for cost-tolerance relationship of R2 value 0.9997, there by eliminating errors due to curve fitting in case of regression fitting. And it also generates more robust outcomes of tolerance synthesis. The proposed non conventional optimization technique obtains an optimal solution better than that of simulated annealing [6] and Response surface methodology (RSM) [1]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002631_0954410018774678-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002631_0954410018774678-Figure1-1.png", + "caption": "Figure 1. Illustration of related coordinate systems.", + "texts": [ + " In \u2018\u2018Controller design\u2019\u2019 section, based on FTSM combined with the new potential function, two finite-time antisaturation controllers are designed, and corresponding stabilities of the system under the controllers are proved via the Lyapunov theory. In \u2018\u2018Simulation analysis\u2019\u2019 section, numerical simulations are given to verify the effectiveness of the controllers. The conclusion of the thesis is presented in the final section. Problem formulation The coupling motion model of attitude and orbit The relative motion dynamics model. In order to establish the spacecraft attitude and orbit coupled model expediently, as shown in Figure 1, the chaser spacecraft body coordinate system Fcb(OcbXcbYcbZcb) is defined: The origin of coordinate Ocb is located at the center of mass of the chaser. Xcb is the docking axis of the chaser. Ycb is perpendicular to the longitudinal symmetry plane containing the docking axis. Zcb meets the right-hand rule with Xcb and Ycb. The definition of target spacecraft body coordinate system Ftb(OtbXtbYtbZtb) is similar to Fcb. In the earth centered inertial (ECI) frame, the dynamics of the chaser and the target are \u20acrc \u00bc rc r3c \u00fe fdc mc \u00fe Fc mc \u00f01\u00de \u20acrt \u00bc rt r3t \u00fe fdt mt \u00f02\u00de where is the gravitational constant and mc and mt are the masses of chaser and target" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003617_01691864.2019.1657947-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003617_01691864.2019.1657947-Figure2-1.png", + "caption": "Figure 2. The mechanical structure of the of the joint-type flexible endoscope.", + "texts": [ + " The balance of the paper is organized as the follows: Section 2 contains the force analysis and modeling of the endoscope. Section 3 explains the experimental setup. In Section 4, we compare the results of experiments and the predictions of the model. In Section 5, we discuss hysteresis for the bending shape and the accuracy of the model. Conclusions are set forth in Section 6. The bending section of an endoscope is driven by four cables. Guiding rings are distributed on the inner wall of the link with intervals of 90\u00b0 to provide the cable channels, as shown in Figure 2(a). The adjacent two links are connected by riveted joint. Driven by the cable tension, the tip bends towards the tension force, as shown in Figure 2(b). In order to simplify the model, we only focus on the bending of a single direction, and certain assumptions are made, as follows: A1: The friction force between the cable and backbone equals the maximum static friction force when the cable slides along the guiding rings. A2: Gravitational force is ignored because all the backbones are very small (the entire weight is less the 10 g). A3: The extra length of the cable added by stretching is ignored because the tension force is small. A coordinate system is defined at each joint, as shown in Figure 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003437_ceit.2018.8751810-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003437_ceit.2018.8751810-Figure2-1.png", + "caption": "Fig. 2. Inputs and outputs of AR.Drone", + "texts": [], + "surrounding_texts": [ + "Drones possess a challenging task in terms of the control. One of the reasons for this can be stated as that the performance of a controller generally depends on the accuracy of the mathematical model of the system. On the other hand, obtaining a highly accurate model of a drone is rather difficult. Most of the mathematical models widely used in literature are neglecting the effect of the drag forces resulting from the friction of the drone with the wind. Furthermore, the data coming from the sensors are inevitably noisy and involve some level of uncertainty, which can affect the performance of the controller adversely even if a sufficiently accurate model of the system is available. Under such circumstances, the use of model free approaches can be considered as a viable solution, as an approximate model of the system is generally sufficient for these kind of controllers. Among the model-free design approaches, artificial neural networks (ANNs) and fuzzy logic control can be called as the most widely used and extensively researched ones. In this study, a square shaped trajectory reference has been generated to examine the maneuverability of the platform. Besides neural network and fuzzy logic controllers, PID controllers for the roll and pitch of the drone have been also designed and applied to the trajectory tracking task to provide a mean of comparison. The deviations from the reference trajectory in x- and y-directions and their derivatives have been used as the input signals to be fed to the controllers. For the simulations and real-time experiments, the sampling time is set to 0.065s. The real time experiments have been carried out indoors to enable to have control over the environmental parameters." + ] + }, + { + "image_filename": "designv11_33_0003997_robio49542.2019.8961842-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003997_robio49542.2019.8961842-Figure2-1.png", + "caption": "Figure 2. Leg configuration and frame definition of our quadrupedal robot Pegasus. The 3DOF leg joints are defined as hip (roll), thigh (pich) and knee (pitch).", + "texts": [ + " Section \u2162 shows the analysis on energy-efficient stance postures based on experimental results. Section \u2163 introduces the CNN-based foothold classifier that can help our robot to realize energy-efficient locomotion. Section \u2164 shows the experimental result on a real quadrupedal robot platform. Section \u2165 discusses the conclusions. Our experimental platform is the quadrupedal robot Pegasus which weighs about 35.2 kg with 1.0 m long and 0.4 m wide. Each leg has three joints (hip, thigh and knee) powered by electric motors shown in Fig. 2. All the algorithms run on an onboard embedded computer (NVIDIA Jetson TX2) running Linux and ROS. To perceive the surrounding terrain information, a rotating lidar (Hokuyo UTM-30LX-EW) is mounted in the front of robot which rotates a full cycle at 1.5 s. The robot is expected to navigate to the goal position as efficiently as possible. To achieve this objective, we developed an entire system architecture (shown as Fig. 3) which includes the following modules: localization, perception, mapping, foothold selection, swing-leg trajectory planning and COG trajectory generator" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002159_gt2017-63208-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002159_gt2017-63208-Figure7-1.png", + "caption": "Fig. 7 Gear shrouds", + "texts": [ + " Each shaft is supported by ball bearings and sealed with a carbon mechanical seal. The oil supplies of the bearings and seals were performed using oil-jet nozzles placed in the spacers, as shown in Fig. 6. Oil was then scavenged from the lower part of the gearbox. Therefore, there was no contribution of oil to the loss occurring in the gear. The temperatures were measured at the gearbox housing, bearings, seals, and oil pipes. The flow rates were measured at each oil pipe of the gears, bearings, and the seals. Gear Shrouds, Oil Jet Nozzles, and Test Conditions. Figure 7 shows the tested gear shrouds. Shroud 1 has six openings on its periphery. Shroud 2 has two openings. In Shroud 2, which contained fewer openings than Shroud 1, we assumed that oil reacceleration loss increases as oil reflows into the gear mesh area, as shown in Fig. 7(b). An oil-jet nozzle was directed from the into-mesh to the gear mesh. The other oil-jet nozzle was directed from the out-of-mesh to the gear mesh. Table 2 shows the test conditions. The flow rate of the oil-jets to gears was changed. The flow rates of the oil-jets to bearings and seals were kept constant. In the following section, \u201coil-jet supply\u201d refers to the oil-jet supply to gears. Friction Loss Control. To minimize any errors in the measurement of friction loss, it is necessary to keep the temperature constant in the bearings, seals, and housing [20]", + " As a consequence, in fluid dynamic loss and windage loss, the experimental values approximately agree with the simulated values. Consideration on the correspondence between the lossclassification and the CFD. Figure 18 shows the simulated oil distribution (iso-surface of the oil volume fraction in the cells = 0.5). In Shroud 2, oil is observed below the smaller gear. It is considered that oil that is not discharged remains in the shroud because the opening area is small in Shroud 2. Possibly, the remaining oil causes the increase of oil reacceleration loss, as shown in Fig. 7(b). The validity of the loss-classification method suggested in this paper can be supported with the consistency between the increase of loss in the experiment and the flow visualization in the CFD. 8 Copyright \u00a9 2017 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 08/19/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use New loss-classification method in which loss is experimentally decomposed is suggested. Experiments were carried out to demonstrate the effectiveness of the method" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001422_pime_conf_1964_179_264_02-Figure21.2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001422_pime_conf_1964_179_264_02-Figure21.2-1.png", + "caption": "Fig. 21.2. Section of oilJiIm", + "texts": [ + " The film boundaries are taken to be at zero reference pressure, that is, at atmospheric pressure. Solution of a partial differential equation is required to establish the pressure distribution and the corresponding acceleration distribution with the assumed boundary conditions. The force on the journal is found from the resultant load of the pressure distribution. The steps in the calculation are summarized below. The calculation is made as far as convenient in non-dimensional terms. Consider a section of the film, as shown in Fig. 21.2, subtending a small angle S/l and of small axial width Sz. The circumferential length of the section is R Sj3. The film thickness h* at position /3 is small compared with R. Owing to the eccentricity of the journal, h* has changed to (h*+Sh*) at the other end of the section. Where the film is convergent, Sh* is negative. One of the assumptions made for calculation is that the oil is incompressible. If there had been flow of the oil across the faces of the section, the condition of continuity would have given the summation of volume flow as zero", + "1 Dimensional Symbols Any consistent system of absolute units is applicable. C D e h* L N n U W 8 81 + 8 2 + Diametral clearance, r = C/2 . J o u e l diameter, R = D/2 . = OA Eccentricity of journal. Film thickness at 8. Bearing length. Speed of rotation, w = 2nN. = e/r Eccenmcity ntio. Peripheral velocity of journal, U = mDN. Steady load applied vertically downwards. Angle from attitude line. Angular extent of pressure bearing film. Attitude angle. Horizontal and vertical axes x, y and attitude axes k, 1 as shown on Fig. 21.1. For local co-ordinate axes see Fig. 21.2. P,*, Py* x*, y* ax, ay FL Viscosity of lubricant. P Density of lubricant. P Pressure of lubricant. Disturbance force superimposed on steady load. Displacement of journal axis from steady-running Acceleration of journal axis. position. Non-dimensional Parameters S Re Sommerfeld duty parameter defined as - Reynolds number defined as 1 Ur = 5- NDC. \u201c P P U Vol 179 P1 3J at Gazi University on March 2, 2016pcp.sagepub.comDownloaded from 44 D. M. SMITH Non-dimensional Symbols APPENDIX 21.11 Sommerfeld (2) solved certain integrals which occur in the analysis of steady-running conditions in circular journal bearings, using the auxiliary angle y defined by INTEGRALS U S E D IN CALCULATION p,* P Y * Disturbance force P, = -9 P - -w y - w x* y* Displacement Y = --,y = - r n p Velocity -J - x a, Acceleration a,' = - = -, alv' = 2 = 3 w a u2r w2 war w w Coefficients (111," + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002839_j.mechmachtheory.2018.06.021-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002839_j.mechmachtheory.2018.06.021-Figure1-1.png", + "caption": "Fig. 1. The toroidal CVTs: (a) Full-toroidal CVT; (b) Half-toroidal CVT.", + "texts": [ + " All rights reserved. Toroidal continuously variable transmission (CVT), with the advantages of high torque capacity [1,2] and high power density [3] , was originally designed for the automobile transmission and successfully marketed in 1999 [4] . Nowadays, toroidal CVTs are also applied to kinetic energy recovery systems [5\u20137] and recommended to be used in wind power generators [8,9] , mechanical presses [10] , four-bar mechanisms [11] ,urban buses [12,13] and other off-highway vehicles [14,15] . As shown in Fig. 1 , a typical toroidal CVT mainly consists of an input disc, an output disc and serval rollers in the cavity formed by the discs. The CVT system relies essentially on the rolling traction (with limited slip) between the disc and the roller whose shapes result in nominal point contact [16] . However, metal-to-metal contact is prevented by the compacting traction oil between the two contact components. By tilting the rollers as shown in Fig. 1 , the speed ratio between the input and the output can be adjusted. A toroidal CVT can be classified into full-toroidal CVT and half-toroidal CVT. Full-toroidal CVT, proposed by Hunt in 1877, has a limited efficiency because of the spin loss induced by the spin motion. Spin motion is caused by the different speed distributions of the two contact areas and can be described as an angular velocity along the normal \u2217 Corresponding author at: School of Mechanical Engineering, Xihua University, No. 999 Jinzhou Road, Chengdu, Sichuan Province, 610039, PR China", + " The overall CVT mechanical efficiency can be rewritten as: \u03b7 = (1 \u2212 C r in )(1 \u2212 C r out ) t out t in , (55) where \u03b7 is the total efficiency, Cr in and Cr out are the creep coefficients at input and output contact points, respectively. t in and t out are input and output traction coefficients, respectively. t in and t out can be calculated by the following equations: t in = \u03bcin + \u03c7in sin ( \u03b8 + \u03b3 ) ; t out = \u03bcout \u2212 \u03c7out sin ( \u03b8 \u2212 \u03b3 ) (56) where \u03bcin and \u03bcout are the tangential force coefficients and \u03c7 in and \u03c7 out are the spin momentum coefficients at the input and output contact points, respectively. \u03b8 and \u03b3 (see Fig. 1 ) are half-cone angle and tilting angle of the CVTs, respectively. The calculation methods of the six coefficients ( Cr in , Cr out , \u03bcin , \u03bcout , \u03c7 in and \u03c7 out ) have be introduced in [26] . When the input torque is increased, the values of Cr in and Cr out become larger. Because, the larger torque will induce the larger shearing deformation of the traction oil. As stated in Section 2.1 , the larger deformation is the reason of the slip loss that is represented by Cr in and Cr out . The increased Cr in and Cr out also slightly influence the spin ratio" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001682_speedam.2010.5544954-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001682_speedam.2010.5544954-Figure1-1.png", + "caption": "Fig. 1 : Cross section of a three-phase prototype", + "texts": [ + " This simplified radial structure is then used for the study of axial heat transfer, in the r-z plane. The study focuses on a thermal machine closed, cooled by water at the external frame. An internal convection is ensured by shaft mounted fans on both sides of the rotor (Cf. Fig.4). II. INVESTIGATED MACHINE The synchronous machine studied in this article is a new structure of flux switching synchronous machine with a hybrid excitation. This particular structure uses the principle of both flux switching and flux concentration. We developed a prototype (see Fig. 1 and Fig. 2). This machine is composed of a stator that includes armature coils, permanent magnets and a wound inductor. The salient rotor is simply made of stacked, soft iron steel. The prototype is a three-phase machine containing twelve magnets, with each phase composed of four magnets and four concentrated coils. The rotor contains Nr teeth (with Nr=10), and the relation between the mechanical rotation frequency F and the electrical frequency f can be expressed as: f = Nr F. We can see from the picture presented in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003032_1350650118818345-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003032_1350650118818345-Figure1-1.png", + "caption": "Figure 1. The general journal bearing with two axial grooves.", + "texts": [ + " W is the static force acting on journal bearings and c is the radial clearance. kxxx R \u00fe kxyy R cxxx I cxyy I \u00bc fRx kxxx I \u00fe kxyy I \u00fe cxxx R \u00fe cxyy R \u00bc fIx kyxx R \u00fe kyyy R cyxx I cyyy I \u00bc fRy kyxx I \u00fe kyyy I \u00fe cyxx R \u00fe cyyy R \u00bc fIy 8>>>>< >>>>: \u00f04\u00de As previously discussed, another four complementary equations were needed to identify all eight linear dynamic coefficients. The complementary equations were deduced by taking the circular journal bearing with two grooves as an example. The structure of the bearing is shown in Figure 1. The oil film zone was separated into two parts by the supply grooves and the supply pressure was ignored in the following deduction process. At present, the Reynolds equation is the main method used to solve the lubrication problem of journal bearings. To maintain numerical stability in the analysis, the nondimensional form is usually used in practice, which is shown in the following equation @ @; H3 @P @; \u00fe d l 2 @ @l H3 @P @l \u00bc 3 @H @; \u00fe 6 _x sin; \u00fe _y cos ;\u00f0 \u00de \u00f05\u00de where P is the dimensionless oil film pressure; d and l denote the diameter and length of journal bearings; H is the oil film thickness, which is defined as H \u00bc 1\u00fe ycos; \u00fe xsin;; l is the dimensionless coordinate in the axial direction, which is defined as l \u00bc z= l=2\u00f0 \u00de; and ; and z denote the coordinates in the axial and circumferential directions", + " To avoid an ill-conditioned matrix problem, the equation was retransformed to the dimensionless form by equation (3) kxx \u00fe kxy \u00bc 0 kyx \u00fe kyy \u00bc \u00f019\u00de where \u00bc u 1 u21\u00fe v21 u1 u2 1\u00fe v2 1 \u00bc v 1 u21\u00fe v21 v1 u2 1\u00fe v2 1 \u00bc 0 1 1 u21\u00fe v21 0 1 1 u2 1\u00fe v2 1 8>< >: \u00f020\u00de Combining the dynamic equations (equation (4)) and the proposed complementary equations (equations (12) and (19)), the parameter identification matrix and observation vectors are A \u00bc = = 0 0 0 0 \" cos \" sin xR yR xI yI xI yI xR yR 2 6664 3 7775 f1 \u00bc 0 0 fRx f Ix T f2 \u00bc = 2 fRy f Iy T p1 \u00bc kxx kxy cxx cxy T p2 \u00bc kyx kyy cyx cyy T 8>>>>>< >>>>>: \u00f021\u00de where A is the identification matrix; f1 and f2 denote the observation vectors; p1 and p2 are the parameter vectors to be identified; and is defined as the maximum value between j j and . The identified parameters were calculated by the following p1 p2 \u00bc A 1 f1 f2 \u00f022\u00de A numerical experiment was used to verify the proposed method. The circular journal bearing with axial grooves, shown in Figure 1, was taken as the example for analysis. The parameters of the bearing system are shown in Table 1. The numerical experiment was calculated by solving the Reynolds equation under a dynamically loaded condition. The simulation was done in three steps. The static equilibrium position of the journal was calculated under a given static load and rotational speed in the first step. Then the whirl orbit of the journal was defined in step 2 by the following equation x \u00bc u0 \u00fe Ax sin!t y \u00bc v0 \u00fe Ay sin!t \u00f023\u00de where u0 and v0 are the static equilibrium positions of the journal in step 1; Ax and Ay denote the whirling radius in the horizontal and vertical directions, respectively; and " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001070_1.4002694-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001070_1.4002694-Figure2-1.png", + "caption": "Fig. 2 Some linear legs with different serial structure", + "texts": [ + " . ,u is a 6 6 auxiliary Hessian matrix versus T k. Their formulas are derived. Equations 1 \u2013 8 provide the foundation for the kinematics analysis of the linear legs with different structure and the dynamics analysis of the limited-DOF PMs with linear legs. 3 Angular Velocity/Acceleration of Linear Legs 3.1 Some Linear Legs With Different Serial Structure. A limited-DOF PM may include some different active/passive linear legs ri. Each of ri may be connected by various serial chains, as shown in Fig. 2 and Table 1. Each of U includes two crossed revolute joints Rij j=1,2 in leg ri. Let Rij be the unit vector of Rij, ij be the rotational angle of Rij, and ij be the angular velocity of Rij. The velocity vbi of leg ri at the connection point bi with m can be derived from Fig. 1 as follows: vbi = vri i + i ri i = vri i + i ri = v + ei 9 Here, i is the angular velocity of linear leg ri. Cross-multiplying both sides of Eq. 9 by i yields i vbi = i vri i + i ri i = i i ri i = ri i i \u00b7 i \u2212 ri i i \u00b7 i = ri i \u2212 ri i i \u00b7 i 10 In order to simplify the kinematics analysis of various legs, it is supposed that i ri in Refs. 12\u201314 . Thus, Eq. 10 leads to i \u00b7 i = 0, i = i vbi /ri 11 However, Eq. 11 is only suitable to some linear legs ri, which lower end is connected with the base B by a revolute joint R, and Transactions of the ASME 016 Terms of Use: http://www.asme.org/about-asme/terms-of-use R p r d a t e S R R c U U S S U i J Downloaded Fr ri is satisfied see Fig. 2 a . When the lower end of the linear leg ri is connected with B at oint Bi by joint U and the constraint conditions Ri2 ri ,Ri2 Ri1 are satisfied, i= i1+ i2 is satisfied. Here, i1 and i2 are the angular velocities of ri about Ri1 and Ri2, espectively see Fig. 2 c . Since ri is not always kept perpenicular with a plan Q including i, i1, and i2, i ri is not lways satisfied. Hence, in the light of some different linear legs, he derivations of the angular velocity i and the angular accelration i are analyzed. 3.2 i and i of Linear Legs With R at the Lower End. ome linear legs with R at the lower end such as RPS, RPU, and PU can be used for 3RPS PM or RPS+RPU+UPS PM in efs. 1,10 . These linear legs rotate about revolute joint R that onnects ri with B. The angular velocity i of ri is along R, and i T i=0 is satisfied see Fig. 2 a . From Eqs. 3 , 4 , 8 , and 11 , it leads to Table 1 The 12 types of linear active/passive legs Type of linear legs with different serial structures PS active leg SP active leg SP passive leg PU active leg RPS active leg UP passive leg PU active leg RPU active leg UPR passive leg PR active leg RRPR active leg UPU passive leg -universal joint, P-prismatic joint, S-spherical joint, R-revolute joint, bold symbol s active joint ournal of Mechanisms and Robotics om: http://mechanismsrobotics.asmedigitalcollection", + " 19 with respect to time, their angular acceleration i can be derived from Eqs. 8 and 19 as follows: FEBRUARY 2011, Vol. 3 / 011005-3 016 Terms of Use: http://www.asme.org/about-asme/terms-of-use t S a + c j + c s H E D a f E D t 0 Downloaded Fr i = D1 v + ei \u2212 vri i ri \u00b7 D1 \u2212 i ri \u00b7 D1 ri \u00b7 D1 = 1 ri \u00b7 D1 Ri1 \u00b7 i Ri2 \u00b7 i ri \u2212 v + ei \u00b7 D1 i + D1 a + ei + ei \u2212 ari i \u2212 vri i i 20 3.4 i and i of SPR or RRPR-Type Linear Legs. An acive SPR-type linear leg includes a serial chain spherical joint -active prismatic joint P-revolute joint R, and ri R is satisfied see Fig. 2 e . It can be used for a 3DOF 3SPR PM in Ref. 10 , 3DOF 2UPS+SPR+SP PM in Ref. 26 , and a 5DOF 4UPS SPR PM in Ref. 10 . A composite active RRPR-type linear leg ri includes a serial hain active revolute joint R1-revolute joint R2-active prismatic oint P-revolute joint R3 . It can be used for a 4DOF 2UPS RRPR 27 . In this linear leg ri, the active revolute joint R onnects the lower end of ri with base B at O and some contraints R1 Z ,R1 R2 ,R2 ri ,R3 R2 ,R3 y are satisfied. Their angular velocity i can be expressed as follows: i + RR = 21 ere, R is the rotational speed of R and R is the unit vector of R", + " 23 with respect to time, the ngular acceleration i of the SPR linear leg ri is derived as ollows: i = i v + i ei + ri i i T \u2212 vri i /ri = i i v + i a + i i ei + i ei + ei + ri i i T + vri i T + ri i i i T i i ion i of SPU-type linear leg is derived as follows: 11005-4 / Vol. 3, FEBRUARY 2011 om: http://mechanismsrobotics.asmedigitalcollection.asme.org/ on 01/29/2 + ri i i T \u2212 vri i /ri 24 3.5 i and i of SPU-Type Linear Leg. An active SPU-type linear leg includes a serial chain spherical-active prismaticuniversal joint, see Fig. 2 b . It can be used for 3DOF 3SPU +UP PM in Ref. 25 , 3DOF 2SPU+SPR+SP PM in Ref. 26 , 5DOF 4SPU+SPR PM in Ref. 10 , and 6DOF 6 SPU PM. In this linear leg ri, each of U includes two crossed revolute joints Rij j=1,2 , Ri1 is fixed on platform m and Ri1 Ri2 and Ri2 ri are satisfied. Thus, its angular velocity i can be expressed as follows: i = \u2212 i1Ri1 \u2212 i2Ri2 25 Cross-multiplying both sides of Eq. 25 by ri leads to ri \u2212 i1Ri1 ri \u2212 i2Ri2 ri = i ri = vbi \u2212 vri i = \u2212 \u0302i 2v + \u0302i 2e\u0302i 26 Dot-multiplying both sides of Eq", + " A central rotational-active RRPR-type linear leg r0 i=0 includes a serial chain active revolute joint R1-revolute joint R2-prismatic joint P-revolute joint R3 . It can be used for 4DOF 3UPS+UPR PM 28 . In this linear leg r0, the active revolute joint R connects the lower end of r0 with base B at O, and some constraints Transactions of the ASME 016 Terms of Use: http://www.asme.org/about-asme/terms-of-use l m + m t fi o i e t T c m r 4 T a w a b g t c d a L p a r l J Downloaded Fr 3.7 i and i of SP, SP, or UP Linear Legs. An active SP inear leg ri includes a serial chain spherical joint S-active prisatic joint P see Fig. 2 f . It can be used for a 4DOF 3SPU SP PM 29 . In this ri, the cylinder of P is fixed onto m, and ri is satisfied. A central passive SP linear leg r0 i=0 includes a serial chain spherical joint S-prismatic joint P . It has been used for a protoype of 4DOF 4SPU+SP PM 10 . In this r0, the cylinder of P is xed onto platform m at o, and m r0 is satisfied. A central passive UP linear leg r0 i=0 includes a serial chain universal joint U-prismatic joint P . It can be used for a prototype f 3DOF 3SPU+UP PM in Ref" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001768_jnn.2010.1722-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001768_jnn.2010.1722-Figure2-1.png", + "caption": "Fig. 2. Illustration of morphological variations forms of carbon nanotubes, (a) hollow-tube, (b) herringbone, and (c) bamboo-like. (d) Representation of edge-plane-like defects and basal plane in a single MWNT.", + "texts": [ + "40 41 While a SWNT can be described as a single tube of graphite, MWNTs are made of more complex structures and present several morphological variations.11 Among these, the most commonly reported are the \u201chollow tube,\u201d \u201cherringbone\u201d and \u201cbamboo-like\u201d forms, which are consecutively depicted in Figures 2(a\u2013c). Compton and co-workers have suggested and presented experimental evidence that similarly to a graphite electrode, the end of the tubes can be described as \u201cedge-plane-like\u201d while the tube walls themselves are \u201cbasal-plane-like\u201d (Fig. 2(d)).22 Actually, herringbone and bamboo-like forms are expected to possess a high density of these edgeplane-like defects, because in both cases the plane of the graphite sheet is at an angle to the axis of the tube.22 Comparing the electroreduction of ferricyanide and the oxidation of epinephrine and NADH in aqueous solutions using bamboo-like carbon-nanotube-modified electrodes with edge-plane pyrolytic graphite (EPPG) electrode, Compton and colleagues were able to demonstrate that the electrochemical activity of nanotube-modified electrodes is equivalent to that EPPG electrode and hence suggesting the site defects at the end or along the axis of the tube are the responsible for the electrochemical reactivity of bamboolike carbon nanotubes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001718_s11431-011-4499-5-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001718_s11431-011-4499-5-Figure7-1.png", + "caption": "Figure 7 The 2-PRU /PR(Pa)R mechanism.", + "texts": [ + " Link 3 has only a translating motion for its two rotation axes are parallel to plane O-xy, so 3,6 Im 6,3 Im ( 0 0 0). Loop II is formed by the link groups RPC and RRC, and the common constraint of RRC is gz IIm (0 0 0 0), X IIm = 6,3 Im (0 0 0)+ gz IIm (00 0 0)= X IIm (00 0 0)=2. IIF = X II P i iII mpn II 16 =6\u00d72(5\u00d72+4\u00d71)+2=0, F=FI +FII=3+0=3. The link group RRC in loop II is an Assur link group with X IIm =2 and mobility FII =0. We can obtain from eq. (10) that F=6\u00d77(5\u00d76+4\u00d73)+(1+2)=3. Example 5. Figure 7 illustrates the 2-PRU/PR(Pa)R mechanism [23] with three independent loops. In plane O-yz, link group 2-PRU\uff08axes C and D form a U-pair\uff09forms loop I denoted by ABCDEF by connection in series. Every link in the loop can neither translate along x-axis nor rotate around z-axis\uff0cthe constraint of the virtual loop is X Im (0 0x 0 0)=2, FI = In6 IP i ip1 + X Im =6\u00d775\u00d78+2=4. The constraint of the general pair 13,4 IG is 13,4 Im (0 0 x 0 0)=2, for link 4 can neither translate along x-axis nor rotate around z-axis too" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001517_ical.2011.6024750-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001517_ical.2011.6024750-Figure3-1.png", + "caption": "Fig. 3 The diagram illustrating the thruster biasing problem.", + "texts": [ + " The generalized inverse solution is given by: 1 1 1( )T T cf W T TW T \u03c4\u2212 \u2212 \u2212= (9) The generalized inverse is defined as: \u2020 1 1 1( )T TT W T TW T\u2212 \u2212 \u2212= (10) Note that if the cost factors on the diagonal are all equal (for example the identity matrix W = I ), then the equation reduces to the pseudo-inverse 1( )T TT T TT+ \u2212= (11) Thruster Biasing allows azimuth thrusters to counteract each other in groups so that the resulting effect of the biasing is zero. Each group can contain either two or three thrusters. The concept of thruster biasing comes from [12, 13], and it is illustrated in Fig. 3. It is useful in many cases, such as an azimuth thruster cannot give zero thrust, a higher power consumption is required (than is actually needed for positioning) or the weather is calm. It is also helpful to reduce the turning of azimuth thrusters when demand is changing. Thruster biasing does not limit the use of the thrusters since the counteraction will be reduced when the total demand increases. The scope of thruster biasing concept can be extended to groups, when a group composed by two adjacent thrusters is regarded as a single azimuth thruster, and this is called group biasing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003000_ece.2018.8554979-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003000_ece.2018.8554979-Figure4-1.png", + "caption": "Fig. 4. Star Connected Windings", + "texts": [ + " The design of 5 phase BLDC motor contains 12 poles and 15 slots for windings and operated at 500 rpm maintaining a voltage difference of 72 degree between phases. In Fig 3 the internal view of BLDC motor is shown. Rotor windings are divided into 5 sets representing 5 phases namely A B C D and E rendering each phase three set of rotor coils and connected in series with each other. The wiring scheme is done in star connection to increase the speed of motor. The wiring arrangement for five phase BLDC prtotype motor is shown in Fig 4. The overall switching sequence requires a complex inverter control scheme governed by 10 step commutation cycle, operating at a switching frequency of 50Hz. Hall sensors mounted on the rotor windings provide the current position of the rotor thus provide a feedback for the BLDC driver circuit to perform accurate commutation sequential switching. The motor is energized with a input voltage of 220V AC and mosfet circuitry provides 12V on the rotor windings. The motor is operated on the principle that at any commutation sequence two of rotor windings are positively energized, two of the rotor windings are negatively energized and the remaining one rotor windings is either positively or negatively energized as per commutation sequence requirements" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000345_6.2009-5890-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000345_6.2009-5890-Figure1-1.png", + "caption": "Figure 1. Modifications to the NextGen flight-test asset.", + "texts": [ + " The viability of this architecture is proven through batch simulation studies and hardware-in-the-loop testing. The MFX-Concept is used to demonstrate the advanced control technologies presented in this paper. The MFX-Concept is a 265 pound aircraft with a two degree-of-freedom morphing wing designed by NextGen Aeronautics. The aircraft is similar in size and shape to the MFX-2, the flight-test asset that was developed and flown by NextGen during a related research program. As can be seen in the schematic (Figure 1), the fuselage and the empennage of the MFX-2 are altered to resemble a more generic aircraft. Specifically, the fuselage is more cylindrical and the empennage is changed from a V-tail with ruddervators to a conventional tail with a rudder and an elevator. During hardware-in-the-loop testing, one ruddervator of the MFX-2 served as the rudder of the MFX-Concept and one ruddervator of the MFX-2 served as the elevator of the MFX-Concept. Because neither actuator is subjected to real flight loads, this simplification did not affect the validity of the results" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000777_s0025654411040017-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000777_s0025654411040017-Figure8-1.png", + "caption": "Fig. 8. Fig. 9.", + "texts": [], + "surrounding_texts": [ + "By setting \u03b4 = 0 in Eqs. (3.6) and (3.7) and by introducing new variables z1 and z2 by formulas z1 = z + a1\u03d5 and z2 = z \u2212 a2\u03d5, we obtain the equations z\u2022\u20221 + a11(c1z1 + \u03b31z \u2022 1) + a12(c2z2 + \u03b32z \u2022 2) = 0, z\u2022\u20222 + a21(c1z1 + \u03b31z \u2022 1) + a22(c2z2 + \u03b32z \u2022 2) = 0, (4.1) MECHANICS OF SOLIDS Vol. 46 No. 4 2011 a11 = 1 + a2 1 \u2212 a1f1, a12 = 1 \u2212 a1a2 \u2212 a1f2, a21 = 1 \u2212 a1a2 + a2f1, a22 = 1 + a2 2 + a2f2. (4.2) The dimensionless variables ak and fk (k = 1, 2) in formulas (4.2) are the ratios of the corresponding variables introduced in the preceding sections to the radius of inertia \u03c1; i.e., we assume that the radius of inertia \u03c1 is equal to unity. The variables z1 and z2 have the following physical meaning. These are small deviations of the points O1 and O2 from their equilibrium positions described in Sec. 3. Equations (4.1) and (4.2) and relations (1.1) with the use of the Lyapunov theorem permit solving the above stability and instability problem for any fixed v. Let us write out the characteristic polynomial of the linear system (4.1) as p4 + p3(a11\u03b31 + a22\u03b32) + p2(a11c1 + a22c2 + \u03b31\u03b32\u0394) + p(\u03b31c2 + \u03b32c1)\u0394 + c1c2\u0394 = 0, (4.3) where \u0394 = a11a22 \u2212 a12a21. Applying the Routh\u2013Hurwitz theorem to the polynomial (4.3), we arrive at the following result. Proposition 1. For the asymptotic stability of the equilibrium in the case of rectilinear and uniform motion of the vehicle, it is necessary and sufficient that the rolling friction coefficients f1 and f2 (and hence the velocity v of the vehicle translational motion determined by (1.1)) satisfy the inequalities a11\u03b31 + a22\u03b32 > 0, a11c1 + a22c2 + \u03b31\u03b32\u0394 > 0, \u0394 > 0, (4.4) (a11\u03b31 + a22\u03b32)[a11\u03b32c 2 1 + a22\u03b31c 2 2 + \u03b31\u03b32\u0394(\u03b31c2 + \u03b32c1)] \u2212 (\u03b31c2 + \u03b32c1)2\u0394 > 0, \u0394 = a11a22 \u2212 a12a21 = a(a + f2 \u2212 f1), a = a1 + a2. (4.5) The dimensionless parameters aij (i, j = 1, 2) are determined by formulas (4.2). Consider several particular cases where Proposition 1 applies. Proposition 2. If the system is completely symmetric, i.e., if \u03b3k = \u03b3 = 0, ck = c, fk = f, k = 1, 2, a1 = a2 = a0, (4.6) then inequalities (4.4) hold for each f ; i.e., the asymptotic stability of the vehicle equilibrium considered above is preserved for any wheel rolling friction coefficients and hence for any velocities of its translational motion. The proof is by a straightforward verification of inequalities (4.4) under conditions (4.6). In this case, it follows from (4.5) that \u0394 = a2 > 0 and the first two inequalities in (4.4) are satisfied, because they are reduced to the following ones: 2\u03b3(1 + a2 0) > 0, 2c(1 + a2 0) + \u03b32a2 > 0. The last inequality in (4.4) has the form c2(1 \u2212 a2 0) 2 + 4\u03b32c2a2 0(1 + a2 0) > 0. Thus, the stability conditions (4.4) are satisfied for any f \u2265 0, and the proof of Proposition 2 is complete. Proposition 3. Assume that all conditions (4.6) except for the last are satisfied; i.e., a1 =a2 (the case of incomplete symmetry; i.e., the car body center of mass is displaced). Then, for a1 < a2, the stability conditions (4.4) are satisfied for every f \u2265 0, and for a1 > a2, the stability is violated for f > f0 = 1 a1 \u2212 a2 [ 2 + a2 1 + a2 2 + \u03b5(a1 + a2)2 \u2212 \u221a \u03b52(a1 + a2)4 + 4(a1 + a2)2 ] > 0, \u03b5 = \u03b32 c . (4.7) Remark. The quantity f0 in inequality (4.7) can be arbitrarily small for some values of the system parameters. For example, as \u03b5 \u2192 0 (small damping or large rigidity of the wheel) and for a1a2 = 1 and a1 > 0, we have f0 = (a1 \u2212 1)(1 + a2 1)/(a1 + a2 1), which can be arbitrarily small as a1 \u2192 1. MECHANICS OF SOLIDS Vol. 46 No. 4 2011 The proof is by a straightforward verification. Under the conditions given in Proposition 3, the instability inequalities (4.4) become \u03b3[2 + a2 1 + a2 2 + (a2 \u2212 a1)f ] > 0, c[2 + a2 1 + a2 2 + (a2 \u2212 a1)f ] + \u03b32(a1 + a2)2 > 0, \u0394 = (a1 + a2)2 > 0, x2 + 2x\u03b5(a1 + a2)2 \u2212 4(a1 + a2)2 > 0, x = 2 + a2 1 + a2 2 + (a2 \u2212 a1)f. (4.8) For a2 > a1, these inequalities are necessarily satisfied for any f > 0. (For the first three inequalities in (4.8), this is obvious, and for the last, this follows from the inequality x > 2 + a2 1 + a2 2 > 2(a1 + a2).) But if a2 < a1, then the first inequality in (4.8) is violated for f > f1 = (2 + a2 1 + a2 2)/(a1 \u2212 a2), the second is violated for f > f2 > f1, the third is always satisfied, and the last is violated for f \u2208 [f0, f00], where f0 is given by the formula in (4.7) and f00 is given by the same formula with the plus sign at the radical. It is easily seen that 0 < f0 < f1 < f00, which implies that all inequalities in (4.8) are violated for f > f0. The proof of Proposition 3 is complete. Proposition 4. Let f2 = 0, \u03b30 = 0, and f1 = f > 0; i.e., assume that the fore (driving) wheel does not experience any rolling resistance and there is no dissipation in it. In this case, if a1a2 < 1, then the instability arises for f > a1, and if a1a2 > 1, then the instability arises for f > (a1a2 \u2212 1)/a2. Proposition 5. Let f1 = 0, \u03b31 = 0, and f2 = f > 0; i.e., assume that the rear (driven) wheel does not experience any rolling resistance and there is no dissipation in it. In this case, if a1a2 < 1, then the instability arises for f > (1 \u2212 a1a2)/a1, and if a1a2 > 1, then the system is stable for any f > 0. The proof of Propositions 4 and 5 is by a straightforward verification of inequalities (4.4) in Proposition 1 under the assumptions stated above. Let \u03b31 > 0, \u03b32 = 0, and \u03b32 > 0. Then the following assertion holds. Proposition 6. Under the above conditions, the asymptotic stability domain D1 in the plane of the parameters (f1, f2) is given by the following inequalities. If 0 < a1a2 < 1, then 0 < f2 < 1 \u2212 a1a2 a1 , 0 < f1, c1a1f1 \u2212 c2a2f1 < c0, f1 \u2212 f2 < a1 + a2. (4.9) If 1 < a1a2 < \u221e, then 0 < f1 < a1a2 \u2212 1 a2 , 0 < f2, c1a1f1 \u2212 c2a2f2 < c0, (4.10) c0 = c1(1 + a2 1) + c2(1 + a2 2). (4.11) The proof of Proposition 6 is by a straightforward verification of inequalities (4.4) under the restrictions given in the statement. Note that if one of the inequalities in conditions (4.10), (4.11) is not satisfied, then the vibrations under study are unstable. Now let \u03b31 = \u03b32 = 0; i.e., assume that there is no dissipation in the wheels. In this case, Eq. (4.3) is biquadratic, the stability of system (4.1) can be only nonasymptotic, and the conditions for such stability are equivalent to the conditions that this biquadratic equation has only pure imaginary roots. These considerations lead to the following result. Proposition 7. For \u03b31 = \u03b32 = 0, the stability domain D0 in the plane of parameters (f1, f2) is given by the following inequalities: f1 > 0, f2 > 0, c0 > c1a1f1 \u2212 c2a2f2, a1 + a2 > f1 \u2212 f2, [c0 \u2212 (c1a1f1 \u2212 c2a2f2)]2 > 4c1c2(a1 + a2)(a1 + a2 \u2212 f1 + f2). (4.12) The proof of Proposition 7 follows from well-known necessary and sufficient conditions for the existence of pure imaginary roots of a biquadratic equation. Proposition 8. The inclusion D1 \u2282 D0 holds. MECHANICS OF SOLIDS Vol. 46 No. 4 2011 The proof of Proposition 8 is by a straightforward substitution of the extreme values in the linear inequalities (4.9) or (4.10) into inequalities (4.12) and verification that the latter are satisfied. Remark. Proposition 8 means that this nonsymmetric addition of dissipation narrows down the stability domain in the plane of the parameters (f1, f2). Graphically (schematically), the domains D1 and D0 in various possible cases are shown in Figs. 2\u201315, where digit 1 corresponds to the line f1 \u2212 f2 = a1 + a2, digit 2 corresponds to the line c1a1f1 \u2212 c2a2f2 = c0 = c1(1 + a2 1) + c2(1 + a2 2), digit 3 corresponds to the line f2 = |1 \u2212 a1a2|/a1, digit 4 corresponds to the line f1 = |1 \u2212 a1a2|/a2, digit 5 corresponds to the line f1 \u2212 f2 = \u2212|1 \u2212 a1a2|(a1 + a2)/(a1a2), and digit 6 corresponds to the parabola (c0 \u2212 c1a1f1 + c2a2f2)2 \u2212 4c1c2(a1 + a2)(a1 + a2 \u2212 f1 + f2) = 0. Figures 2, 4, 6, 8, 10, 12, and 14 correspond to the case of \u03b31 = \u03b32 = 0, while Figs. 3, 5, 7, 9, 11, 13, and 15 correspond to the case of \u03b31 = 0, \u03b32 = 0 or \u03b31 = 0, \u03b32 = 0 for all possible distinct values of the parameters a1, a2, c1, and c2. Now consider the case of weighted wheels, m1 = 0 and m2 = 0. Since we assume that the car body mass is m = 1, the quantities m1 and m2 are the respective dimensionless ratios of the wheel masses to the car body mass. Simple calculations show that the form of the vibration equations (4.1) remains the same and the coefficients aij (i, j = 1, 2) are determined by the functions a11 = \u03c12 \u2212 (a1 + am2)(\u2212a1 + f1) \u03c12 0 , a12 = \u03c12 \u2212 (a1 + am2)(a2 + f2) \u03c12 0 , a21 = \u03c12 + (a2 + am1)(\u2212a1 + f1) \u03c12 0 , a22 = \u03c12 + (a2 + am1)(a2 + f2) \u03c12 0 , where \u03c12 0 = m0\u03c1 2 and m0 = 1 + m1 + m2. Then one can readily compute that \u0394 = a11a22 \u2212 a12a21 = a(a + f2 \u2212 f1)/\u03c12 0. Proposition 1 remains valid under the above variations in aij , i, j = 1, 2, and \u0394. In a similar way, one can restate Propositions 2\u20138. Equations (4.1) and (4.2) describing the vehicle vibrations under the action of the rolling friction forces admit the following analogies (interpretations) to well-known systems in mechanics. Each of the coefficient matrices multiplying zk and z\u2022k (k = 1, 2) in (4.1) with (4.2) taken into account consists of MECHANICS OF SOLIDS Vol. 46 No. 4 2011 two parts. The first part (it is said to be unperturbed) does not contain the friction coefficients f1 and f2. The second (perturbed) part depends on these coefficients linearly. As follows from the above derivation of the equations of motion, the origin of the perturbed part significantly depends on the following two external causes. One of them is the interaction between the deformable wheels and the rough road, which causes the rolling resistance torques Mk1 and Mk2; the second cause is the torque M1 from the engine (external energy source) to the driving wheels ensuring their (and the driven wheels) permanent rolling if there are resistance torques mentioned above. If these causes are absent, then the system motion occurs for f1 = f2 = 0, and Eqs. (4.1) contain only the first parts of the above-mentioned matrices for which the solutions have the property of (asymptotic) stability (a conservative system MECHANICS OF SOLIDS Vol. 46 No. 4 2011 under the action of some forces with complete or incomplete dissipation). But if f2 1 + f2 2 = 0, then these causes result in the appearance of additional terms in Eqs. (4.1), which, according to [6], can be interpreted as positional (circular) forces (for the corresponding terms with z1 and z2); gyroscopic and nonconservative (depending on the velocities) forces (for the corresponding terms with z\u20221 and z\u20222); conservative positional forces (for the corresponding terms with z1 and z2); and dissipative (Rayleigh) MECHANICS OF SOLIDS Vol. 46 No. 4 2011 MECHANICS OF SOLIDS Vol. 46 No. 4 2011 forces (for the corresponding terms with z\u20221 and z\u20222). In the general case, such a decomposition of forces was performed in [7]. Thus, the existence of the above-mentioned external causes leads to the appearance of additional forces of the described structure, whose destructive character is well known in mechanics of gyroscopic systems, structural mechanics (in particular, in mechanics of bridges), and other adjacent fields; it is described in the literature [8\u201310]. Even if there forces do not actually lead to the vehicle instability, they still significantly change the frequencies of its vibrations in the vertical plane, and this fact cannot be ignored when solving the corresponding problems. These facts were already considered earlier by the authors in the paper [11]." + ] + }, + { + "image_filename": "designv11_33_0002017_adv.2017.383-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002017_adv.2017.383-Figure1-1.png", + "caption": "Figure 1. (Left) A Janus sphere consists of a silica microsphere half-coated with platinum (Pt). The particle self-propels away from the Pt side in hydrogen peroxide. (Right) Depending upon the relative orientation of the Janus spheres in a dimer, we have different conformations.", + "texts": [ + " Moreover, at the micron and submicron length-scales, the effect of thermal fluctuations is sufficient to result in coupling between deterministic and stochastic dynamics, leading to emergent phenomena such as spiral diffusion [17, 18] for rotary swimmers. Here, we study the self-assembly of Janus spheres into dimers and the dynamics of the resulting microswimmer. A Janus sphere is comprised of a micron-sized dielectric silica sphere that is half-coated with a catalyst made of platinum (Pt). As shown in the left part of Fig. 1, this particle self-propels away from the platinum side in hydrogen peroxide by a self-phoresis mechanism [19, 20]. Upon moving and interacting, the Janus spheres may self-assemble into https:/www.cambridge.org/core/terms. https://doi.org/10.1557/adv.2017.383 Downloaded from https:/www.cambridge.org/core. Columbia University Libraries, on 03 Jun 2017 at 10:05:40, subject to the Cambridge Core terms of use, available at dimers, trimers, or form larger clusters [21]. As shown in Fig. 1, the relative orientation of the Janus spheres in a dimer leads to many possible conformations. In analogy with cis-trans isomerism, we call the dimers with Pt caps facing the same direction (with the plane of structural symmetry about the point of the spheres\u2019 contact) as cis-dimers, and dimers with the Pt caps facing opposite directions (with the point of structural symmetry about the point of contact) as trans-dimers. The majority of dimers rotate rather than exhibit rectilinear motion as a result of conformational asymmetry", + " This was accomplished by pipetting a small amount of the colloidal suspension onto the solid surface of the silicon wafer. Afterward, the wafers, which contained the microspheres, were placed into a vacuum environment and two layers of metal were deposited onto one-half of the particles: first 5 nm of Ti was deposited to ensure adhesion to the silica surface, which was followed by a 10 nm layer of Pt. Thus, the particles became Janus spheres with two distinct, nearly equal-area surfaces: one of silica and the other Pt, as shown schematically in Fig. 1. As the Pt is an efficient catalyst in the breakdown of hydrogen peroxide, the particles were activated by adding this \u201cfuel\u201d to the colloidal suspension. We observed the Janus spheres being propelled in a direction away from the catalyst site. The particles self-assembled into clusters and exhibited the effect of phase-separation in which a \u201csolid\u201d phase of close-packed particles coexisted with a \u201cgas\u201d phase of individual free particles [21]. The particle- and clustertrajectories were captured by video microscopy using a Zeiss Axiophot microscope with dry objectives with magnifying power 10\u00d7, 20\u00d7, and 50\u00d7", + " The red dashed line is the theoretical prediction according Eq. (6). The deviation of experimental trajectories from the theory is mainly due to the small number of starting frames used for ensemble averaging. We recently discussed how the relative orientations affect the dynamics of such self- assembled motors [25]. In that same study, we showed that certain conformations are more prevalent in proportion to other conformations, and these ratios are dependent upon hydrogen peroxide concentration. However, the cis-conformation (see Fig. 1) is more prevalent, in comparison to all other conformations of dimers, at all hydrogen peroxide concentrations, suggesting the occurrence of this conformation is independent of activity. Considering this observation, the likely explanation for the cis-conformation\u2019s prevalence, is that the metal caps of two adjacent spheres form a \u201cbridge\u201d between the two particles during the manufacturing process, and thus the two spheres remain rigidly fixed in this conformation regardless of activity. More interestingly, there is a strong positive correlation between hydrogen peroxide Figure 4", + " The solid spiral represents the experimental data and the dashed curve is the theoretical prediction. The dimensions are scaled by the instantaneous orbit of motion, which has a radius \ud835\udc45 = \ud835\udc63/\ud835\udf14. https:/www.cambridge.org/core/terms. https://doi.org/10.1557/adv.2017.383 Downloaded from https:/www.cambridge.org/core. Columbia University Libraries, on 03 Jun 2017 at 10:05:40, subject to the Cambridge Core terms of use, available at concentration, and thus activity, and the percentage of the trans-conformation (see Fig. 1). The assembly displays rotation about the point of contact of the two Janus spheres, which are oriented roughly in opposite directions [25, 26], and thus take on a staggered formation. Since the propulsion direction is away from the catalyst site, the dimer also spins with the metal caps trailing (see Fig. 1). Figure 5 shows the progression of the formation of such a trans-dimer. We suggest that the trans-conformation may be, in general, dynamically more stable in comparison to others or, the trans-conformation may simply be less susceptible to becoming assembled with another cluster. That is, from control experiments, with rare exception such as seen in the frames of Fig. 5., the dimers form before the addition of hydrogen peroxide. Furthermore, in our experiments, the trans-dimers were observed to be highly impervious to aggregating with other particles, while non-trans-conformations appeared to self-assemble with other clusters more readily" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000020_1.3610022-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000020_1.3610022-Figure3-1.png", + "caption": "Fig. 3 Three-loop planar kinematic chain", + "texts": [ + " The xiyizraxcs define a right-handed Cartesian coordinate system rigidly attached to link i, and the four parameters af) a,-, 0,-, and S{ define the position of link i + 1 relative to that of link 1. It has been shown that the relative positions of the two coordinate systems can be mathematically stated in terms of the following (4 X 4) transformation: A. = JL 0 a,- cos di cos di at sin 8{ sin 0< 0 Si 0 0 -sin di cos sin dt sin a,cos di cos cti \u2014 cos di sin a,- sin a,- cos cti (1) Around any closed loop of n links, the n + 1 coordinate system is the same as coordinate system 1, and this may be expressed as A1A2 \u2022 A. = I (2) For simplicity, the three-loop planar example of Fig. 3 is used as a basis for the following derivation. The mechanism is a planar analogy of a standard, link-type rear-end suspension and consists of three kinematic loops having two degrees of freedom. Choosing the pair variables associated with prismatic pairs along axes ZV and Z,% as inputs, the matrix loop equations may be written: Loop 1 AiA,/1?A3AsA., = I Loop 2 AiAiAsJsAs = I Loop 3 AiAiMA 9 A io A a = I (3) Loop 3 is an \"auxiliary\" reference loop added to relate the motion of the car body to the roadway (or solid rear axle)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001900_j.proeng.2017.02.331-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001900_j.proeng.2017.02.331-Figure7-1.png", + "caption": "Fig. 7. (a) Positions of the rotating ring and the sleeve\u2019s face depending on oscillation frequency; (b) Positions of the rotating ring and the rotor end with two sleeves in both positive and negative directions", + "texts": [ + " That can cause the destruction of the rotating ring and leads to failure of the seal unit. Fig. 6(b) shows the dependence of rotating ring and sleeve velocities on operation time. Starting from point 1 (point 1 corresponds to point 1 in Fig. 6(a)) rotating ring begins to decelerate relative to the rotor end; this eventually leads to separation at point 2. At point 3, the contact is reduced and the speed of the rotating ring and sleeve\u2019s face are identical. Separation time will increase if the oscillation frequency grows 1t < 2t < 3t (Fig. 7(a)). Rotating ring oscillation amplitude may exceed the amplitude of the rotor, however, due to the presence of the second sleeve (Fig. 4) the amplitude value is limited by the clearance ' . As a result, the time and amplitude of the rotating ring oscillation are limited (Fig. 7(b)). Thus, for each oscillation period, two collisions of the ring and the sleeves take place at the points 3 and 4 (points 1 and 2 are similar to points 1 and 2 in Fig. 6(a)). This phenomenon is not permitted and can lead to failure of the seal unit. Calculation results of the analytical three-mass model derived in the first and third parts of this article were compared with the results of the simulation by ANSYS Rigid Dynamic software. Initial data for calculation are presented in Table 1. 1. Joint solution of mathematical models (to calculate the static and dynamic characteristics of the gas film and the gap) has revealed that the elastic and damping components of the gas film dynamic response significantly exceed stiffness and damping properties of the secondary seals and springs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000976_ssp.188.400-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000976_ssp.188.400-Figure2-1.png", + "caption": "Fig. 2. Formiga P100 SLS machine. Working principle", + "texts": [ + " Each process involves several specific activities (Fig. 1.) without which the purpose of the paper cannot be accomplished. The samples subjected to the study were manufactured using SLS process, on EOS Formiga P100 machine. The rapid prototyping machine uses plastic powders as raw material to build 3D objects in an additive layer by layer process. In the prototyping process, the machine uses a 30 W CO2 laser for selective sintering of the powder particles. The working principle of the Formiga P100 machine is illustrated in the Fig.2. The powder polymer used to fabricate the samples was two times recycled Polyamide PA200 \u2013 EOSINT having grain size of 60 \u00b5m. In order to analyze the influence of the energy density on the sample morphology three jobs have been performed, each of them having different laser speed. Each job has produced five identical square samples, one positioned in the center of the building envelope and the others on the corners, according to the Fig. 3. The main numerical process parameters which have been used for prototyping jobs are presented in Table 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003959_iecon.2019.8926722-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003959_iecon.2019.8926722-Figure6-1.png", + "caption": "Fig. 6: Drawing of the weight reduction.", + "texts": [ + " Table II shows previously defined values for calculating (34) and (36). The values were determined in terms of strength and fabrication of the reduction gearbox. In Table II, ti is the thickness and Di is the density where i \u2208 {b, p1, p2}, and cij , i \u2208 {b, p1}, j \u2208 {p1, p2} is the distance between i and j. Thus, the thickness and the positional relationship between the balancer, the first and second planet gear cannot be changed. Therefore, in order to change mi and Zi, i \u2208 {b, p1, p2} in (34) and (36), weight reduction by removing material is performed. Fig. 6 shows the explanatory drawing of weight reduction. Here, Si, i \u2208 {b, p1, p2} is the weight reduction area. Fig. 6(a) highlights the areas where material is removed. If mi is divided into the weight of the thin parts m\u2032 i corresponding to the area shown in yellow and the weight of the remaining frame part mfi shown in red, it is given by mi = mfi +m\u2032 i = mfi + aiDiSi, i \u2208 {b, p1, p2} (37) where ai is the actual thickness corresponding to weight reduction area and mfi are values that also take into consideration the weight of the eccentric parts (screw, planet eccentric shaft, center eccentric shaft, bearing). Thus, (34) changes as mfp1 +m\u2032 p1 = mfb +m\u2032 b +mfp2 +m\u2032 p2. (38) In addition, using the parameters ti, cij and bi, i \u2208 {b, p1, p2} shown in Fig. 6(b), (36) changes as (e1 + tp1 2 )mfp1 + (e1 + bp1)m \u2032 p1 = tb 2 mfb + bbm \u2032 b + (e2 + tp2 2 )mfp2 + (e2 + bp2)m \u2032 p2 (39) where bi, i \u2208 {b, p1, p2} is position of the thickness ai, e1 = tb+cbp1 and e2 = e1+ tp1+cp1p2. Using the values shown in Table II, variables of ai and bi, which satisfy (38) and (39), are determined. The calculation results are shown in Table III. Here, since (38) and (39) could not be satisfied, the reduction gearbox was made using the value with the smallest error from the calculated values" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003285_s11837-019-03462-3-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003285_s11837-019-03462-3-Figure5-1.png", + "caption": "Fig. 5. (a) Mechanism of formation of induced Lorentz force, (b) main melt flows inside the weld pool, and (c) comparison of melt flow when welding with (black curves) and without (white curves) a magnetic field.", + "texts": [ + " In addition, the area of the HAZ first increased from 5.37 mm2 to 5.96 mm2 when the intensity was less than 60 mT, then decreased to 4.83 mm2. This indicates that the utilization of the laser energy was further increased by reducing the heat Xu, Rong, and Huang2298 conduction to the HAZ when the intensity was above 60 mT. The increased area of the FZ results from the magnetic field-induced Lorenz force, which effectively reduces the velocity and thermal convection in the flowing melt in the weld pool. Figure 5(a) shows the mechanism explaining the formation of the induced Lorentz force. It is hypothesized that the melt volume DV is infinitesimal and that its length along the direction of the magnetic field is Dl. The velocity of the melt is l t\u00f0 \u00de. A melt with flow direction perpendicular to the magnetic field can then induce an electric current density of j t\u00f0 \u00de \u00bc Dl l t\u00f0 \u00der B Dl \u00bc l t\u00f0 \u00deBr \u00f01\u00de where r is the electric conductivity and B is the intensity of the magnetic field. The interaction between the induced electric current and external magnetic field results in a Lorentz force of F t\u00f0 \u00de \u00bc BIDl \u00bc l t\u00f0 \u00deB2rDV \u00f02\u00de Note that the direction of the Lorentz force is contrary to the melt flow direction, thus it acts as a braking force. Schaefer et al.25 investigated the melt velocity distribution inside a weld pool by using an x-ray imaging system. The results provide a reference for the melt velocity distribution in the present study. The main melt flows inside the weld pool during laser welding are shown in Fig. 5(b). Only the melts with flow velocity components in the x\u2013y plane can interact with the external magnetic field and thereby induce a Lorentz force. Marangoni convection is dominant in the upper weld pool (UWP) because of the large-scale Marangoni vortices caused by surface tension forces, with a flow origin in the high-temperature region around the keyhole.18 The melt flow in the bottom weld pool (BWP) is first driven by the recoil pressure from the keyhole tip.26 To clarify the influence of the magnetic field on the melt flow behavior, the resulting reduction in the melt flow displacement is decisive", + " Note that the deceleration au was hypothesized to be 1 m/s2 in this study (density q = 4.43 9 103 kg/ m3, electric conductivity r = 5.875 9 105 S/m). Similarly, the other flow components (lux, lbx, and lby) have the same displacement variation. Therefore, when an external magnetic field is applied during laser welding, the melt displacements will be Magnetic-Field-Induced Partial-to-Full Penetration Evolution and Its Mechanism During Laser Welding 2299 reduced and new compressed flow paths are formed, as shown in Fig. 5(c). The effect of the magnetic field is to change the velocity and heat convection of the melt flows, giving rise to the formation of a weld pool of different size. These changes increased with increase of the magnetic field intensity, and the process is complex. To better understand the mechanisms, the evolution processes shown in Fig. 8 are simplified and it is hypothesized that the changes in the bottom weld pool occur before those in the upper weld pool. Figure 8(a) shows a three-dimensional (3D) view of the weld pool when welding at 0 mT" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002110_whc.2017.7989929-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002110_whc.2017.7989929-Figure4-1.png", + "caption": "Figure 4. Structure of fabric-jamming sheet", + "texts": [ + " The piled-plastic films can be made to fit on a spherical surface by employing a woven structure with strip-shaped films [25]. The variable-stiffness plates comprising plastic films capable of fitting on a spherical surface have been suggested [26]. However, the deformability of these mechanisms is limited by the size and nonstretchability of the films. In this paper, the structure and fundamental mechanical properties of the FJS are presented. A glove-type force display and a tabletop-type force display developed using the FJS are presented. Fig. 4 shows the structure of the FJS containing the piled-stretch fabrics. The stretch fabrics made of nonelastic fibers (such as polyester) are piled, and subsequently, covered with a rubber bag. The air inside the bag is evacuated through a fixed air tube. As shown in Fig. 5 (a), the fabrics are elongated by changing the shape of the meshes rather than elongating the fibers of the fabric. The entire sheet comprising piled fabrics and a rubber bag is as elastic as a rubber sheet. When the air inside the bag is evacuated, the atmospheric pressure helps in bonding the piled fabrics tightly" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003290_j.procs.2019.02.040-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003290_j.procs.2019.02.040-Figure1-1.png", + "caption": "Fig. 1. Coordinate systems of quadrotor.", + "texts": [ + " In [12] the problem of synthesis of quadrotor control is solved by one of the symbolic regression methods (variational analytical programming method [13]). The purpose of this study was to develop models that take into account the effect of cross-connections and the evaluation of this impact on the accuracy of the model. This work may be useful for specialists in various fields describing the motion of the quadrotor, as well as the synthesis of its control system. 2. The mathematical model of quadrotor motion in three-dimensional space with respect to perturbations and cross connections Fig. 1 represents the relative positions of the normal ground coordinate system 0 g g gO x y z and the associated coordinate system ,Oxyz the positive direction of reference of yaw angle , pitch angle and roll angle , thrust force iP and torque iM generated by the propellers 1,4,i the direction of rotation and angular velocities i of the propellers. The transition from the associated coordinate system to the normal ground coordinate system can be performed using the transition matrix [4]: cos cos cos cos sin sin sin sin cos sin cos sin sin cos cos sin cos cos sin cos sin sin sin cos sin sin sin cos cos R " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003222_ffe.12997-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003222_ffe.12997-Figure2-1.png", + "caption": "FIGURE 2 Pair of gear teeth in contact and the equivalent two cylinders model5,17", + "texts": [ + "5 In the following sections, the developed calculation model and its application to the meshing cylindrical involute gear teeth\u2014a representative case of rolling\u2010sliding contact loading\u2014is presented in more detail. Rolling\u2010sliding line contact of two convex bodies, such as the one between two meshing gear teeth flanks, can be observed as a combination of a simultaneous conformal contact (in the plane defined by gears' axes) and a nonconformal contact (in the plane normal to the gears' axes) of two equivalent cylinders whose radii R1 and R2 are equal to radii of curvature of teeth flanks \u03c11 and \u03c12 in the instantaneous point of contact (Figure 2). Due to the contact nonconformity, and with sufficiently small deformations assumption, width of the contact surface is rather small in comparison to the size, ie, radii of curvature of the undeformed bodies in contact. Consequently, relevant stresses and strains are highly concentrated to a small volume in the immediate vicinity of the contact region, and they diminish quickly with distance from the area of contact. Furthermore, the shape and size of contacting bodies, ie, gear teeth, further away from the contact, as well as the way they are supported, have negligible influence on stresses and strains and their distribution in the vicinity of contact" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003882_s11071-019-05343-5-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003882_s11071-019-05343-5-Figure1-1.png", + "caption": "Fig. 1 Schematic diagram of coordinate systems for ball bearing", + "texts": [ + " Finally, the frequency spectrum characteristics of the radiation noise regarding each component are analysed. Generally, in actual operations, the outer ring of the bearing is assembled into the bearingpedestal, but it can still vibrate. The inner ring rotates at a constant speed. It is assumed that the centre of mass and geometrical centre of each component are the same. The movement of the cage is guided by the inner ring. To better analyse the vibration and noise characteristics of the high-speed ceramic angular contact ball bearing, the coordinate systems of the bearing are set up as shown in Fig. 1. As shown in Fig. 1, the inertial coordinate system {O; X,Y, Z} is fixed, and coordinate origin O is fixed to the initial centre of the bearings. The X -axis represents the bearing rotation axis, which is parallel to the ground. The Y - and Z -axes represent the horizontal radial and vertical radial directions, respectively. The following notation is used to describe the components: ball (b), inner ring (i), outer ring (o), cage (c), cage pocket (p), and ordinal number ( j) for balls or cage pockets. In the coordinate system of the j th ball {Ob j ; Xb j ,Yb j , Zb j}, the origin Ob j represents the centre of mass of the ball, the Xb j axis is the axial direction of the bearing, the Yb j axis is the circumferential direction of the bearing, and the Zb j axis is the radial direction of the bearing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001437_iri.2011.6009590-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001437_iri.2011.6009590-Figure1-1.png", + "caption": "Figure 1. Formation representation in a 3D workspace, and illustration of the flight path and the orientation angles", + "texts": [ + " These laws allow to reach the goal, avoid collision, and move the formation is a stable configuration. \u2022 Consensus is achieved through algebraic collaboration functions of the states and actions of a group of networked agents. Most published work considers planar formations of unmanned air vehicles by ignoring the elevation ([11, 12, 13]). This formulation simplifies the problem considerably, but restricts the work space of the vehicles. In this paper no such assumption is made. The work space W \u2282 R 3 is attached to a global reference frame of coordinates \u03a9 as shown in figure 1. Point O is the origin of \u03a9. The ith vehicle in the formation is denoted by Vi, i = 1, ..., n; where n is the total number of vehicles in the formation. In the leader\u2013follower formulation, the global leader is denoted by V \u2217 i . Vehicle Vi has the following model x\u0307i = vi cos\u03d5i cos \u03b8i y\u0307i = vi cos\u03d5i sin \u03b8i z\u0307i = vi sin\u03d5i (1) where (xi, yi, zi) represent the coordinates of the reference point of Vi in \u03a9. vi is its linear velocity, \u03d5i is the flight path angle, and \u03b8i is the heading angle. The path of vehicle Vi is denoted by pi(t)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002336_iccas.2017.8204221-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002336_iccas.2017.8204221-Figure4-1.png", + "caption": "Fig. 4 Operating principle", + "texts": [ + " This results in 35% contraction with no load and over 5% for a 150 kgf load in case of 1.5 inch in diameter, 300mm in length and 130g in weight. Two ones are applied for each side and four ones are used in total Each paper must be divided into two parts. The first part includes the title, authors\u2019 name, abstract, and keywords. The second part is the main body of the paper. As shown in Fig. 3, the system consists of a body, a compressor, a solenoid valve, a switch that controls it, and a microcomputer. As shown in Fig. 4, wire is connected to both ends of the McKibben artificial muscle. Note that, in order to realize compact and simple mechanism, we apply one artificial muscle for each of right and left arm-lower back assistance independently. The compressive force generated by the McKibben artificial muscle rotates the shoulder pulley and the leg one through the wire. And then the wearer's arms and lumbar assistance for lifting loads are realized via the arm frame and the leg one connected to each of the shoulder pulley and the leg one" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003052_msf.941.988-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003052_msf.941.988-Figure1-1.png", + "caption": "Figure 1: A schematic drawing of the deposition of a wall-like structure.", + "texts": [ + " Object of the present study is the investigation of temperature profiles during wire-based LMD of wall-like AlMg alloy structures using three constant process parameter sets as well as the development of an approach to reduce occurring distortions. For the experiments an 8 kW continuous wave ytterbium fiber laser YLS-8000-S2-Y12 integrated with an optical head YW52 Precitec and implemented in a CNC-supported XYZmachining center (IXION Corporation PLC) was used. The aluminum alloy AlMg4.5MnZr (EN AW-5087) as wire material with a diameter of 1 mm and AlMg3 (EN AW-5457) as substrate material with a thickness of 3 mm were processed using three process parameter sets. The parameters are shown in Table 1, whereas Fig. 1 shows a schematic visualization of the unidirectional LMD process using local argon shielding. The time between the depositions of two layers was kept constant to 60 seconds. This ensured a sufficient cooling of the structure in such a way that the deposited track was completely solidified and able to remain its shape during the deposition a following layer. In order to analyze the temperature distribution during the LMD process at different positions, three thermocouples were employed for temperature assessment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001300_esars.2012.6387397-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001300_esars.2012.6387397-Figure1-1.png", + "caption": "Fig. 1: Examples of railway electrification systems supplied from an overhead catenary (simplified schematics): a) dc supply; b) ac supply.", + "texts": [ + " This is often realized by adding an inductor, forming an LC-filter together with the dc-link capacitance of the converter(s) (additional filters tuned for the given track circuit frequency may also be added). However, due to volume and economi cal constraints, the size of the filter inductor should be kept small. Considerable research efforts have therefore been put on developing converter modulation schemes that result in dc-link currents that provide signalling compatibility with different track circuits, see e.g. [8]. In ac supplied systems 978-1-4673-1372-8/12/$31.00 \u00a92012 IEEE (see Fig. 1 (b)), the line converter provides additional means for harmonic mitigation but also here, the size of the input filters used should be kept as small as possible. To fit the specific traction requirements, the level of rotor saliency and permanent magnet flux can be tailored by the motor designer [9]. However, depending on the specific winding and rotor configurations, the interaction of harmonics from the rotor magnets and stator winding layout can result in flux-linkage harmonics with significant magnitudes [10], [11]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001347_s10846-012-9725-2-Figure17-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001347_s10846-012-9725-2-Figure17-1.png", + "caption": "Fig. 17 Tested wing model; changing TC1 and RC2", + "texts": [ + " Changes in lift and drag due to alteration of the root and tip chords of the third segment are shown in Table 4. Finally, an outcome of the changes in the angle of attack of the third segment is given in Table 5. In order to determine the partial contribution of each wing segment to the overall lift, drag, moment of inertia and mass, some more complex wing shapes were tested. Due to the basic laws of aerodynamics, a lower contribution of the first and second segment in the total lift and drag, as well as in other physical values is expected. For this reason, the tested wing was modelled as depicted in Fig. 17. Changes in the lift and drag, due to the altering of TC1 and RC2 are given in Table 6. The shapes given in Table 6, with an invariable parameter TC1, ensure a robustness of the wing. Due to its smaller influence on the lift and drag, and the larger influence on the total mass and moment of inertia, for the large range of different materials, according to their density, these shapes might be extremely useful. Altering the conjunction position of the first and second segment, while keeping their sum unchanged, has an insignificant impact in the total lift and drag" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003300_cyber.2018.8688228-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003300_cyber.2018.8688228-Figure2-1.png", + "caption": "Fig. 2. Other configurations for MLR systems", + "texts": [ + "00 \u00a9 2018 IEEE 913 To give details about such development, the rest of this survey is organized as follows. In section II, dynamics modeing for MLR systems is briefly outlined. In section III and IV, studies of multi-lift rotorcraft systems are presented in detail, which consist of coordinated control strategies as well as experimental studies. In section V, the main challenging problems in this field and some development tendencies are discussed. In section VI, some concluding remarks as well as future work are given. Some studies put emphasis on configurations of MLR systems. Fig. 1 and Fig. 2 show the most common configurations in this research field [7], [29], [34], [35]. To make the mathematical modeling of the MLR systems, most of previous work made the following assumptions: 1) The tether is massless and the tether force is always nonnegative. 2) The rotorcrafts are rigid and symmetrical. 3) The slung load is regarded as a mass point. 4) Aerodynamics of the load/tether system is neglected. 5) The suspension point coincides with the rotorcrafts\u2019 center of gravity. Under the above assumption, a majority of relative modeling work was absorbed in twin-lift helicopter systems and considered the coupling in the lateral/vertical plane", + " Three small size helicopters were arranged as an equilateral triangle on the ground, with a distance of 8m between the helicopters. The other experiment was carried out in May 2009, which was similar to the first one. Additionally, researchers from Norwegian University of Science and Technology (NTNU) further verified their strategies by employing two flight experiments [28], where three hexacopter UAVs cooperatively transported an unknown payload of up to 2.2 kg. Both the work proposed by DLR and NTNU adopt the same configuration like Fig. 2(a). In 2017, researchers from ETH Zurich [26] used two AscTec Firefly hexacopters weighing 1600g to carried out experiments for cooperative transportation of a bulky object,a 1.2 m carton tube weighing 370 g. By the same token, scholars from University of Zurich [27] used the same configuration as shown in Fig. 2(c). They utilized two quadrotors weighing 800g to carry a 1m long aluminum rod which weighed 263g and was attached by tethers of 0.4m. However, both experiments from ETH Zurich and University of Zurich\u2019s flight experiments were indoor experiments without respect to some uncertainties as well as external disturbances. Remark 4: The flight experiments about MLR systems are rarely mentioned except relative work mentioned above. It is indeed of great importance to carry out flight experiments to verify the proposed control methods", + " It can be seen from references [4], [35] [19], [20], section II and Remark 1 that some uncertainties could be considered and some assumptions may do harm to systems\u2019 control. So how to derive a reasonable model for a MLR system is of great importance. (2) Reasonable configuration and structure The configurations of MLR systems are diverse, such as with support bars and without support bars. The length of tethers, the suspension position and the mass and shape of slung load are all extremely vital for the research of MLR systems. It is clear from Fig.1, Fig.2 and reference [29], [34], [35] that proper configuration and structure are beneficial to enhance the performance of MLR systems, yet there is no flight test comparing different configurations and structures. Designing proper configuration and structure for MLR systems may be quite interesting. (3) Perception and communication The perception and communication of MLR systems are indeed of great significance. It can be seen from this survey and the references mentioned in this paper that few teams have considered the perception or communication of MLR systems, which means that this topic may be of great interest" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000239_sii.2010.5708334-Figure14-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000239_sii.2010.5708334-Figure14-1.png", + "caption": "Fig. 14. An overview of a quad rotor tail-sitter UAV.", + "texts": [ + " This computer calculates control input based on - 258 - SI International 2010 each sensor data, and sends command signals to 4 motors. Flight data and other information are recorded on a micro-SD card for postexperiment analysis. The main computer receives commands from an R/C transmitter through an R/C receiver, but is not used in control calculation. A R/C transmitter is mainly used by operator to switch to manual control mode if problems occur. An overview of the proposed quad rotor tail-sitter UAV is shown in Fig. 14. The airframe length as the basis for landing condition is 0.8 [m], the width is 0.99 [m] and the overall height is 0.28 [m]. Calculated specification of the quad rotor tail-sitter UAV is shown in Table III. Motor mounting parts increased total weight about 0.2 [kg]. Hence, flight range and flight time decreased than simulation results. In this paper, we developed a new quad rotor tail-sitter VTOL UAV attempt to improve energy efficiency. Three dimensional UAV simulator was developed to evaluate the advantage in energy efficiency of the proposed UAV" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002109_1350650117723484-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002109_1350650117723484-Figure4-1.png", + "caption": "Figure 4. Forces acting on a ball.", + "texts": [ + " Besides, the thermal effect is taken into account through updating bearing parameters using temperature results from the inverse heat transfer model of the spindle. The bearing characteristics, with or without thermal effects, can be obtained and compared in detail. The proposed approach can analyze an angular contact ball bearing not only as an individual bearing but also under the influence of its surroundings during actual conditions. High-speed ball bearing model and equilibrium forces on a ball in Figure 4 are given by Koj oj 1:5 sin oj \u00fe lojMgj D cos oj Kij ij 1:5 sin ij lijMgj D cos ij \u00bc 0 \u00f03\u00de Koj oj 1:5 cos oj \u00fe lojMgj D sin oj Kij ij 1:5 cos ij lijMgj D sin ij Fcj \u00bc 0 \u00f04\u00de with A1j \u00bc BD sin p \u00fe z \u00fe xRi sin \u2019j \u00fe yRi cos \u2019j \u00f05\u00de A2j \u00bc BD cos p \u00fe x cos \u2019j \u00fe y sin \u2019j \u00f06\u00de Ri \u00bc 0:5dm \u00fe \u00f0 fi 0:5\u00deD cos p \u00f07\u00de BD \u00bc \u00f0 fi \u00fe fo 1\u00deD \u00f08\u00de cos oj \u00bc X2j \u00f0 fo 0:5\u00deD\u00fe oj sin oj \u00bc X1j \u00f0 fo 0:5\u00deD\u00fe oj cos ij \u00bc A2j X2j \u00f0 fi 0:5\u00deD\u00fe ij sin ij \u00bc A1j X1j \u00f0 fi 0:5\u00deD\u00fe ij 8>>>>>>< >>>>>>: \u00f09\u00de The gyroscopic moments Mg and centrifugal force Fc in equations (3) and (4) are calculated by Fcj \u00bc 1 2 mdm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001992_s40435-017-0333-7-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001992_s40435-017-0333-7-Figure1-1.png", + "caption": "Fig. 1 Reference axes", + "texts": [ + " \u2022 Control effectiveness has no dependency on sideslip angle. Other assumptions are also considered in this paper: \u2022 The gravity field is uniform and there are no changes in gravity with altitude \u2022 HL-20 is assumed to have maximum engine thrust equal to 80 KN \u2022 Thrust is assumed to act along the X body axis and through the center of gravity For deriving the equations of motions, body reference axes which is the most common axes system fixed on the aircraft, has been used. The general form of aircraft and system axes are illustrated in Fig. 1, where the origin o is assumed to be coincident with the center of gravity. Aircraft dynamics in translational and rotational motions are presented by the following nonlinear rigid-body equations of motion [28,29]. u\u0307 = rv-qw \u2212 gE sin \u03b8 + T + XA + XC m (1) v\u0307 = pw-ru+gE sin \u03c6 cos \u03b8 + YA + YC m (2) w\u0307 = qu-pv+gE cos\u03c6 cos \u03b8 + ZA + ZC m (3) p\u0307 = IY qr \u2212 IZqr + L A + LC IX (4) q\u0307 = IZ pr \u2212 IX pr + MA + MA IY (5) r\u0307 = IX pq \u2212 IY pq + NA + NC IZ (6) \u03c6\u0307 = p + q sin \u03c6 tan \u03b8 + r cos\u03c6 tan \u03b8 (7) \u03b8\u0307 = q cos\u03c6 \u2212 r sin \u03c6 (8) \u03c8\u0307 = q sin \u03c6 sec \u03b8 + r cos\u03c6 sec \u03b8 (9) x\u0307E = u cos\u03c8 cos \u03b8 + v(cos\u03c8 sin \u03b8 sin \u03c6 \u2212 sin\u03c8 cos\u03c6) +w(cos\u03c8 sin \u03b8 cos\u03c6 + sin\u03c8 sin \u03c6) (10) y\u0307E = u(sin\u03c8 cos \u03b8) + v(sin\u03c8 sin \u03b8 sin \u03c6 + cos\u03c8 cos\u03c6) +w(sin\u03c8 sin \u03b8 cos\u03c6 \u2212 cos\u03c8 sin \u03c6) (11) z\u0307E = \u2212u sin \u03b8 + v(cos \u03b8 sin \u03c6) + w(cos \u03b8 cos\u03c6) (12) In Eqs", + " Equations (10)\u2013(12) are navigation equations which relate the aircraft body axis velocity to earth axes velocity, where xE , yE , zE are the earth axes coordinates of the system. Aerodynamic forces and moments acting on the aircraft are highly depended on relative motion of the aircraft to the atmosphere. This relative motion is expressed by the total air speed (VT ), attack angle (\u03b1), sideslip angle (\u03b2) and rotational rate of the aircraft with respect to air. These parameters have been shown in Fig. 1. In real conditions, atmosphere is rarely calm and aircrafts usually are confronting with different kinds of AD like winds, gust and turbulence. Therefore, in non-stationary atmosphere which is more realistic, attack angle, sideslip angle and total speed are calculated as follow [30]: \u03b1 = tan\u22121 ( w \u2212 wa u \u2212 ua ) (13) \u03b2 = sin\u22121 ( v \u2212 va VT ) (14) VT = \u221a (u \u2212 ua)2 + (v \u2212 va)2 + (w \u2212 wa)2 (15) In Eqs. (13)\u2013(15), subscript a is related to the speed of the air in body axes coordinate. In NASA technical memorandum for HL_20 [23], a conventional \u201ccoefficient build-up method\u201d has been used in formulation of aerodynamic model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000930_j.saa.2012.04.060-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000930_j.saa.2012.04.060-Figure2-1.png", + "caption": "Fig. 2. Central composite design\u2019s (a) response surface and (b) Pareto chart with estimated effect (absolute value).", + "texts": [ + " The chosen levels for concentration of sodium hydroxde, with the codified level in parenthesis, were: 0.60 mol L\u22121 + \u221a 2), 0.55 mol L\u22121 (+1), 0.45 mol L\u22121 (0), 0.35 mol L\u22121 (\u22121) and .30 mol L\u22121 (\u2212\u221a 2). For the UV exposition time, the chosen levls were: 74 min (+ \u221a 2), 70 min (+1), 60 min (0), 50 min (\u22121) and 6 min (\u2212\u221a 2). Replicates (n = 5) were made only at the central point 0, 0). The response surface, where fluorescence intensity (IF) is odeled by Eq. (1) (using codified coefficients), is shown in Fig. 2a. he results in the Pareto\u2019s Chart (Fig. 2b) indicated a significant uadratic contribution for both variables, which were responsible or the curvatures of the response surface. The linear interaction etween both factors was found insignificant. By using Eq. (1), the est experimental values were 60 min of UV irradiation time for TBZ solution prepared in NaOH 0.45 mol L\u22121. These parameters ere found to be very robust (small variations in a NaOH concenration range from 0.40 to 0.50 mol L\u22121 and an irradiation time from 5 to 65 min) caused no significant signal variation, as indicated in he response surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003896_j.triboint.2019.106096-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003896_j.triboint.2019.106096-Figure3-1.png", + "caption": "Fig. 3. Schematic sketch of the acceleration process used in the two-disc test rig for the recreation of thermally induced WEL in oil-lubricated contacts using bearing inner rings made of 100CrMn6.", + "texts": [ + " The oil level and oil temperature can be adjusted using the corresponding heating and lubrication system. All tests are conducted with an oil temperature of 70 \ufffdC and an oil level up to the axis of shaft 1. In Ref. [25] and based in the findings of [26,33,34] it was proposed that the probability of a smearing/scuffing damage is higher under certain bearing operating conditions such as fast acceleration or deceleration processes and sliding under normal loading. Therefore, a three-step acceleration process was defined in Ref. [29] in order to reproduce WEL (schematic sketch in Fig. 3). In the first step, the servo-motor operating in speed control accelerates the bottom shaft (Shaft 1) and thus the inner ring NU2208 \u2013 \u201cdriver\u201d \u2013 up to a pre-defined angular speed \u03c91. Subsequently, the upper shaft (Shaft 2) is pulled down until both rings are in contact under a defined pre-load F1 of 0.2 kN. The selected pre-load is lower than the load needed to exceed the breaking torque MB, which is applied by the servo-motor operating in torque control to the upper shaft (Shaft 2), resulting in a low loaded sliding contact", + "5 Nm during the plotted test) of the upper shaft and is thus the load necessary to set the follower in motion. The acceleration process is concluded when the follower speed matches the driver speed. The acceleration time is accordingly defined as the time the follower needs to reach the speed of the driver. As shown in Fig. 6-bottom the maximum reached area-related power is 88 W mm2. The reader should note that due to their different circumferential speeds the frictional energy density E/A applied to driver and the follower (according to Fig. 3) have different magnitudes. F. Guti\ufffderrez Guzm\ufffdan et al. Tribology International 143 (2020) 106096 Thus, due to the inverse relationship between circumferential speed and contact time it is clear that the follower, which starts from standstill, experiences a higher energy input. The calculation of the frictional energy density is, therefore, conducted only for the follower. The spatially resolved distribution of the frictional energy density E/A for the acceleration process shown in Fig. 6 is shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000758_s11465-012-0317-4-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000758_s11465-012-0317-4-Figure1-1.png", + "caption": "Fig. 1 Twist and wrench in three-dimensional space", + "texts": [ + " When the n linear independent screws are obtained, we could gain the other 6 \u2013 n reciprocal screws by Eq. (6), vice versa. We call a screw a twist if it represents an instantaneous motion of a rigid body, and a wrench if it denotes a system of forces and couples acting on a rigid body. The physical meaning of reciprocal can be described by the principle of virtual work. If a wrench acts on a rigid body in such a way that it produces no work while the body is undergoing an infinitesimal twist, the two screws are said to be reciprocal screws [28], shown in Fig. 1. The virtual work performed between the wrench $r \u00bc $\u0302 r and the twist $ \u00bc _q$\u0302 is given as \u03b4W \u00bc _q\u00bds\u22c5\u00f0rr sr \u00fe hrsr\u00de \u00fe sr\u22c5\u00f0r s\u00fe hs\u00de \u00bc _q\u00bd\u00f0h\u00fe hr\u00de\u00f0s\u22c5sr\u00de \u00fe rr\u22c5\u00f0r s\u00de \u00fe s\u22c5\u00f0rr sr\u00de , (6) From the geometry of the lines associated with the two screws shown in Fig. 1, the following relationship are obtained s\u22c5sr \u00bc cos\u03b1, (7) sr\u22c5\u00f0r s\u00de \u00fe s\u22c5\u00f0rr sr\u00de \u00bc \u2013 a\u22c5\u00f0s sr\u00de \u00bc \u2013 asin\u03b1, (8) where a is a vector along the common perpendicular leading from the screw axis of $ to $r, and \u03b1 is the twist angle between the axes of $ and $r, measured from $ to $r, about the common perpendicular according to the righthand rule. Substituting Eqs. (7) and (8) into Eq. (6), we obtain \u03b4W \u00bc _q\u00bd\u00f0h\u00fe hr\u00decos\u03b1 \u2013 asin\u03b1 : (9) By definition, the virtual work produced by the two reciprocal screws is equal to zero" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002165_j.jmapro.2017.08.008-Figure11-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002165_j.jmapro.2017.08.008-Figure11-1.png", + "caption": "Fig. 11. (a) Example 2, (b) Support needed points (all points are into coherent region), (c) Grain generation from support needed points, (d) Proposed support structure, and (e) Fabricated part, (f) Demonstrating support removability.", + "texts": [ + " For comparison, the support is also generated using the commercial catalyst software on the same model and printed as shown in Figure 10(a). The generated support following catalyst algorithm is sticking on the model material and difficult to remove as shown in Figure 10(b). The proposed partial contact support technique makes the support removal easy and no artifact is left on the object surface as shown in Figure 10(d). The implementation sequence of the proposed methodology on Example 2 is shown in Fig. 11(a\u2013d). Since all support needed points has normal along \u2212z axis or 1800, it does not have normal vector range discussed in section 2.2. This object has a volume of 25564 cubic mm with a build height of 84 mm. Following the proposed technique, it generates 3769 cubic mm of support volume and takes total 464 min to fabricate the whole object along with support as shown in Fig. 11(e). To demonstrate the support removability, a aker s s n m C o s upport require portion of this object is printed with model and upport material as shown in Fig. 11(f). The support structure is ot sticking on the model material and sufficiently supporting the odel material with smoother finish as shown in Fig. 11(e and f). omparing with the commercial machine, the proposed methodlogy shows 13% improvement of total build time (TBT) where 83% avings of support material as shown in Table 2. Fig. 12 shows Bot (b) MeshMixer (c) Proposition. the comparison of the support structure between commercial support generators and proposed methodology. Numerical comparison with respect to support volume, number of support contour and support build time between commercial support generators and proposed methodology is shown in Table 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000225_004051756703700306-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000225_004051756703700306-Figure5-1.png", + "caption": "Fig. 5. Relationship between ( fiber length )\u2019 and deflection for two mohair fibers Mi and Ms.", + "texts": [ + ", the slope of the line, has a least-squares value [16] of 7.69 \u00b1 0.09 X 10-5 g/micro-amp, with a C.V. of 1.2%.. A typical fiher deflection-applied force (or current) curve for mohair is given in Figure 3, showing a linear relationship and thus satisfying Equation 1. The graph of the deflection A vs 13 gave a straight line, as may be seen for the two kemp fihers illustrated in Figure 4, which also substantiates Equation 1. These results are similar to those for mohair, as shown by the examples in Figure 5. The fact \u2019 that one of the curves for kemp does not pass through the origin is probably due to the fiber not emerging at Yale University Library on May 16, 2015trj.sagepub.comDownloaded from 207 The results for mohair are given in Tables I and I I, while kemp results appear in Tables III and IV. There is a large fiber-to-fiber variation in Eb for kemp fibers, although the values do not vary significantly for a single fiber. The ratio ls/0 is independent of the moment of inertia, although not completely independent of the radius, so that differences seem to be due to the elasticity of the fiber" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003469_j.engfailanal.2019.06.024-Figure16-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003469_j.engfailanal.2019.06.024-Figure16-1.png", + "caption": "Fig. 16. The deformation and Mises stress of the cage.", + "texts": [], + "surrounding_texts": [ + "The maximum total deformation and stress of needle roller and cage are obtained by simulating the same size of the original invalid bearing by FE model, fillet radius of needle roller is set to 0.25mm. As shown in Figs. 15 and 16, the maximum von Mises stress is at one end of the needle, almost the same as the actual failure figure of the needle roller. Stress concentration occurs at the edge of cage, which is also consistent with the failure position of the cage. From the results we can know the maximum Mises stress of the needle is 1332MPa, which isalmost equal to the ultimate contact stress [p]= 1500MPa. The maximum Mises stress of the cage is 292.46MPa, which is less than yield stress of cage material \u03c3b= 325 MPa.\nTable 2 Relative error of the maximum stress.\nElement size of cage 2.0mm 1.0 mm 0.8mm 0.5 mm 0.3mm\nError of needle roller (%) \u221223.663 \u22126.979 \u22122.183 \u22120.202 0 Error of cage (%) \u221241.224 \u221221.251 \u221212.9018 \u22122.153 0 Element size of cage 1.0mm 0.8 mm 0.5mm 0.3 mm 0.2mm Error of needle roller (%) \u221212.156 \u22125.853 \u22121.969 \u22120.505 0 Error of cage (%) \u221233.029 \u221222.950 \u221213.318 \u22122.801 0\nan gl\ne/ \u00b0", + "According to the deformation diagram of the needle roller, the maximum inclination angle of the needle roller is produced:\n= \u2212 \u03b2 \u03b3 \u03b3\nL arctan\nb\nmax min\n(20)\nwhere is the maximum inclination angle of the needle roller,\u00b0; \u03b3max, the maximum deformation at one end of needle roller, m; \u03b3min, the minimum deformation at the other end of needle roller, m.\nIt can be obtained \u03b2=0.0524\u00b0. According to the relative fatigue life of the roller bearing at different inclined angles [32\u201333], if the fatigue life of the bearing without inclined angle is set to 1, the relative fatigue life of needle roller bearings is only 0.02 while\u03b2=0.05\u00b0.\nFrom the above analysis we can know the wear of the right end of the needle roller is obviously larger than the left side, which is consistent with the actual failure form of roller needle and the theoretical mechanical law.The needle and cage edge produce a stress concentration, that gradually wear the roller needle and cage. The inclined angle seriously affects the use of needle roller bearings and accelerates the failure of the needle and cage, which eventually leads to the failure of the cantilever bearing.\n6.2.2. Influence of fillet radius on needle roller Another problem with the failure needle roller is that there is no fillet or the fillet radiusare too smallon the both ends of needle, the initial fillet of needle roller is shown in Fig. 17. Based on the FE model, the failure of the cantilever bearing with different fillet radius (r=0mm, 0.25mm, 0.5 mm, 0.75mm, 1mm, 1.25mm, 1.5 mm) was studied. The simulation results of rolling needle and cage are shown in Fig. 18 and Fig. 19, while under different fillet radius of needle rollers.\nFigs. 18 and 19 show that it caneffectively reduce the maximumvon Mises stress, total deformation and inclination \u03b2 while appropriately increasing the fillet radius. For example, the maximum stress of the needle roller and the cage are 684.36MPa, 98.878Mpa, and the inclination is 0.0272\u00b0while the fillet radius is 0.75mm. The maximum stress, deformation and inclination reduced by 48.62%, 66.19% and 48.09% respectively while compared to fillet radius is 0.25mm. The results show that the fillet radius of needle roller have obvious influence on the maximum stress, deformationand inclination of the needle roller and the cage. By increasing the fillet angle to about 0.75mm, the maximum stress and deformation can be reduced.", + "A new modelfor analyzing the influence of the structural parameters of main shaft on the offset force P required for eccentric ring, the deflection angle \u03b8 at the lower end of the main shaft, the counterforce R1 of self-aligning bearing and the counterforce R2 of cantilever bearing is introduced.The effects of the distance between components and the stiffness of the shell are considered. A FE model of main shaft with cantilever bearing and self-aligning bearing is established to analyze the causes of failure of cantilever bearing. The theoretical model and the FE model are verified by experiments. The main conclusions of this paper are summarized as follows:\n1. The stiffness and inner diameter of the shell have little effect on the target parameters (\u03b8, R1,R2). All those target parameters increase with the increase of the stiffness and decrease with the increasement of the inner diameter of the shell, but the increase and decrease amplitude are smaller. 2. The target parameters aredecrease with the increase of the distance between eccentric ring and bearings,deflection angle satisfies the minimum angle of the tool within the range of L2's changes in the analysis, but L1 should not be less than 0.64m since the deflection angle must be bigger than the \u03b8min.While the distance between the cantilever bearing and the self-aligning bearing is invariant, except that the counterforce R2 decreases with the increase of L1/L,other target parameters are reduced first and then increased with the increase of L1/L, and the L1/L should less than 0.24 or bigger than 0.76 to meet the requirement of the \u03b8min. 3. Base on the original structure parameters of wellbore trajectory control tool of high build-up rate, the radial force at the cantilever bearing can be reduced by decreasing the inner diameter of shell, increasing the distance between the eccentric ring and the self-" + ] + }, + { + "image_filename": "designv11_33_0002846_1.g003430-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002846_1.g003430-Figure1-1.png", + "caption": "Fig. 1 Detection and collision regions for the ith vehicle and its virtual agent.", + "texts": [ + " Accordingly, for any pair of vehicles we define an Avoidance Region as \u03a9r ij fq: q \u2208 RnN; kqi \u2212 qjk \u2264 rg (4) where q qT1 ; : : : ; qTN T .We say that a collision between the ith and jth vehicles occurs if for some time t, q t \u2208 \u03a9r ij. Similarly, for each pair of vehicles we define a Detection Region as DR ij fq: q \u2208 RnN; kqi \u2212 qjk \u2264 Rg (5) The latter implies that the ith and jth vehicles can communicate or detect their positions at time t if q t \u2208 DR ij. The Avoidance and Detection Regions for the ith vehicle are depicted in Fig. 1. To achieve our control objective, namely, trajectory tracking with collision avoidance, we propose a control framework for each vehicle comprising two parts: an online planner and a local trajectory D ow nl oa de d by U N IV E R SI T Y O F G L A SG O W o n Se pt em be r 9, 2 01 8 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /1 .G 00 34 30 tracking control system. An illustration of the control framework is depicted in Fig. 2. The goal of the ith vehicle\u2019s online planner is to trace a collision-free trajectory that converges to the vehicle\u2019s desired trajectory", + " The reader can verify that \u2202VT i \u2215\u2202pi \u22c5 \u03b7Ti 0, which implies that \u03b7i is perpendicular to \u2202VT i \u2215\u2202pi. In addition, note that \u03b7i 0whenever the ith virtual agent is safely away from other vehicles or whenever it reaches the desired configuration. Analogous to the previous definitions ofAvoidance (4) andDetection Regions (5) for the ith and jth vehicles, we define the Avoidance and Detection Regions for their virtual systems as \u03a9rp ij fp:p \u2208 RnN; kpi \u2212 pjk \u2264 rpg and D Rp ij fp:p \u2208 RnN; kpi \u2212 pjk \u2264 Rpg, respectively, where p pT 1 ; : : : ;p T N T . See Fig. 1 for an illustration. Remark 1: To simplify analysis, this Note uses a single integrator model for the virtual system. Other vehicle models (such as double integrators) could be similarly implemented. The role of the virtual agent is to set a safe path for the real vehicle to follow and assuming a double integrator model as in Eq. (1) may potentially generate trajectories that are more akin to the real system. Notwithstanding, note that the trajectories of the single integrator model are smooth enough given that the control inputs for the virtual agents are continuous (except for a set of Lebesgue measure zero)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002165_j.jmapro.2017.08.008-Figure12-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002165_j.jmapro.2017.08.008-Figure12-1.png", + "caption": "Fig. 12. Support generated by: (a) M", + "texts": [ + " Following the proposed technique, it generates 3769 cubic mm of support volume and takes total 464 min to fabricate the whole object along with support as shown in Fig. 11(e). To demonstrate the support removability, a aker s s n m C o s upport require portion of this object is printed with model and upport material as shown in Fig. 11(f). The support structure is ot sticking on the model material and sufficiently supporting the odel material with smoother finish as shown in Fig. 11(e and f). omparing with the commercial machine, the proposed methodlogy shows 13% improvement of total build time (TBT) where 83% avings of support material as shown in Table 2. Fig. 12 shows Bot (b) MeshMixer (c) Proposition. the comparison of the support structure between commercial support generators and proposed methodology. Numerical comparison with respect to support volume, number of support contour and support build time between commercial support generators and proposed methodology is shown in Table 3. Due to the presence of contour plurality [17], the number of contour is increased in proposed methodology, however both build time and support volume Table 3 Comparison between commercial support generators and proposed methodology" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003435_978-3-030-22747-0_49-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003435_978-3-030-22747-0_49-Figure6-1.png", + "caption": "Fig. 6. Selected nodes in a mesh composed of SAE-AISI 1524 Carbon Steel Substrate (200 6 200 mm) and layers of alloy IN-738 (12 2 50 mm each) (a) Mesh used for processes AM1 and AM2 (b) Mesh used for processes AM3 and AM4. The selected nodes are located in the central XY cross-sectional plane (i.e., from the top of cross section along Y axis). Element size 2 1 2 mm. Element type DC3D8.", + "texts": [ + " To predict the evolution of temperature distribution in the entire weldment (substrate, two and three cast IN-738LC alloy layers) for the entire welding and cooling cycle of the process, a 3D transient nonlinear heat flow analysis was performed. To observe the heat transference among layers, all welding layers had gluing contacts. Figure 5 shows the experimental set-up, specimen dimensions, and x, y and z directions for AM weld tests. The initial temperature T0 was set to 20 \u00b0C. The designed mesh is shown in Fig. 6a for processes AM1 and AM2, and in Fig. 6b for processes AM3 and AM4. Numerically predicted thermal gradients, isotherms and thermal cycles produced by the AM process were calculated with the developed thermal model. For the first part of this work, the following results corresponding to the processes AM1 and AM2 were obtained in a mesh representing two wire layers (composed of four finite element rows) plus the substrate (Fig. 6a). The welding conditions used in the simulations are shown on Table 2. Figure 7 shows thermal contourns and thermal cycles calculated at the reported nodes. Compared with the AM2 process, higher peak temperatures are observed in the AM1 process. The reason is because in the AM1 process, a higher heat input is deposited to the workpiece. The following results are obtained for the second part of this work. For instance, Fig. 8 shows thermal cycles calculated at the locations specified in Fig. 6b for process AM3. Even though the heat input is keep constant, it is apparent in Fig. 8 that as the weld velocity and heat distribution parameter decreases, the peak temperature increases. For these particular welding conditions, the effect of the heat distribution parameter is more representative. Figure 9 shows thermal cycles calculated at the reported nodes located in the substrate and in the three layers for the process AM4 at weld velocities of 12.5, 6.25 and 3.125 mm/s. The heat source is applied over the top surface of each layer (50 mm length) during a lapse of time that depends on the welding speed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000336_6.2010-2790-Figure35-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000336_6.2010-2790-Figure35-1.png", + "caption": "Figure 35. Saddle modes of damselfly and dragonfly wings. The saddle mode of a damselfly forewing (left) and dragonfly hindwing (right) show striking resemblance to the other insects considered in our study.", + "texts": [ + " This treatment added an evenly distributed mass load to each wing (as much as 50% for the damselfly), so we know that the actual wing modes were higher than observed. The affect on modal ratios was not known so we elect to not report their values here. However, because an evenly applied powder coat has the effect of only scaling the mass matrix while leaving stiffness undisturbed, the modeshapes themselves should be unaffected, rendering the modeshapes of the treated wings valid representations of the untreated wings. We show the first four modeshapes for the bumblebee in Figure 34 and only the saddle mode for the damselfly and dragonfly in Figure 35. The flap and feather modes were both evident in the damselfly and dragonfly but the bisaddle mode was not. Since their wings were especially transparent, even the powder coat allowed for a fair amount of laser light to pass through the wing and reduce overall quality (coherence) of response data. We suspect that the bisaddle mode may indeed exist, but may have been hiding in \u201cthe noise\u201d of the data. So, the natural question we ask from the underlying similarity is: Why? Why should these very diverse creatures, having wildly different scale, wing shape (planform), venation pattern, and even wingbeat kinematics be so similar in their underlying eigenstructure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001232_tmag.2012.2198203-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001232_tmag.2012.2198203-Figure1-1.png", + "caption": "Fig. 1. Basic structure of two-DOF resonant actuator.", + "texts": [ + " Recently, we have also been developing various kinds of resonant actuators that have multi-degrees of freedom [1]\u2013[5] and a coupled analysis method for analyzing its dynamic characteristics by combining the magnetic field with the electric circuit, control method, and its motion in our finite-element method (3-D FEM) code [6]\u2013[12]. However multi-axis force control for these resonant actuators has not been proposed. In this paper, we propose a two degree-of-freedom (two-DOF) resonant actuator that can be independently driven in the two axes under vector control. The dynamic characteristics of the actuator are clarified by 3-D FEM analysis and the measured results of a prototype. The basic structure of the two-DOF resonant actuator is shown in Fig. 1. This actuator mainly consists of a mover, a stator and resonance springs in the and directions that support the mover. The mover is composed of permanent magnets (NbFeB, T) and a back yoke. The stator is composed of an E-type laminated yoke with three phase excitation coils (45 turns). This actuator is assumed to move with a range of mm in the -direction and mm in the -direction, respectively. This movable range assumes use with a compact appliance such as electric shaver, electric toothbrush and etc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000598_icems.2011.6073667-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000598_icems.2011.6073667-Figure1-1.png", + "caption": "Fig. 1. Prototype of a concentrated-winding exterior-rotor surface-mounted NdFeB-PM machine integrated in a flywheel.", + "texts": [ + "ndex Terms\u2014Slot and pole combination, PM machine, fractional slot concentrated windings, transient FEM, MMF. I. INTRODUCTION In this paper, permanent magnet (PM) machines with fractional slot concentrated windings are used as power supply for domestic load in a small ship application. They have a rated speed of 3150 rpm and an output power about 3kW. The PM machine is integrated in a flywheel of a diesel engine. Its prototype is shown in Fig.1. For applications with low rotor speed of few hundred rotations per minute such as electric bike, wind turbine applications, etc., some authors have studied the influence of combination of slot number and pole number on performances of PM machines: in [1], authors focused on single layer PM machines for bikes; in [2], authors focused on radial force and vibration; in [3], authors concentrated on electromagnetic torque. For electric/hybrid car [4], influence of slot and pole combination on the loss and inductance was studied" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000242_17543371jset39-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000242_17543371jset39-Figure1-1.png", + "caption": "Fig. 1 Ergometer seat bracket and force transducers; the positions of the transducers are shown by the white circles", + "texts": [ + " This study also aimed to demonstrate that, once calibrated, the apparatus could be used to record pertinent information on rowing technique. The rowing ergometer used during the calibration procedure was a modified Concept 2 model D indoor rowing machine (Concept Inc., Morrisville, Vermont, USA). Four Entran ELPM uniaxial load cells (Entran, Lexington, Kentucky, USA) were mounted on the ergometer slide rail to measure the magnitude of vertical forces and the point of force application on the seat (Fig. 1). Each transducer had a range of 1250N and the manufacturer\u2019s calibration stated a hysteresis of \u2013 0.15 per cent and a non-linearity of \u2013 0.15 per cent. The development of the instrumentation used a redundant approach and provided a bracket to which the seat could be attached in the usual way. In this case the redundancy of the design refers to the fact that the system could have used only three load cells. Shear forces were not measured. These modifications were performed as part of another research project [14] and aspects relating to Proc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000188_arso.2010.5680046-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000188_arso.2010.5680046-Figure3-1.png", + "caption": "Fig. 3 Rear view display and control input device", + "texts": [ + " The proposed backward motion control scheme can be summarized as following three steps: STEP1) A driver decides the direction of a passive trailer by securing the rear view display of a passive trailer, and enters the control input to DAS. STEP2) DAS transforms [v1 , 1, 1] into [v0 , 0] by solving inverse kinematics. [v0 , 0] is converted to [v0 , ] STEP3) ECU enters [v0, ] into car-like mobile robot, and the velocity of the car-like mobile robot [v0 , 0 ] generate the velocity of a passive trailer [v1 , 1 ] STEP1 a driver seating on driver s seat and securing the rear view which is displayed by rear view camera, and intuitively decide the driving direction of a passive trailer as presented in Fig. 3. A driver can enter the control input into the DAS using simple control device as joy-pad. The control input is consist of linear velocity and angular velocity of a passive trailer [v1 , 1] which generates constant turning radius. STEP2 The controller transform the angle of revolute joint [ 1] provieded by potentiometer and control input entered by a driver into desired velocity of the car-like mobile robot [v0 0] by eq. (4). Desired velocity of the car-like mobile robot [v0 0] is converted to control input of car-like mobile robot [v0 , ]", + " The maximum linear reversing velocity of a trailer system is 0.07m/sec. This velocity is determined by the reversing velocity and length of the commercially available automobile. In C1), a driver control a trailer system by keypad. The left and right buttons control the steering angle of a trailer system, and the up and down button control the forward, backward motion respectively. In C2), C3), a driver control a trailer system by the simple control pad offers the five control inputs. It is presented in Fig. 3. If drivers want to control a trailer system to the forward motion, a driver can change control mode to C1). Three types of drivers D1)-D3) perform the experiment three times using C1)-C3). Fig. 12(a)-(e) shows the representational results of experiment in condition C1)-C3). Initial pose and final pose is marked green-circle and red-triangle respectively. red-dash -line is trace of a passive trailer. If a driver controls a trailer system in forward motion, it is indicated by blue-line . In Fig. 8(a), a driver repeated forward motion control 31 times to correct the direction of a passive trailer" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002110_whc.2017.7989929-Figure14-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002110_whc.2017.7989929-Figure14-1.png", + "caption": "Figure 14. Schematic of glove-type force display using fabric-jamming sheets", + "texts": [ + " The reaction force obtained by pushing the center of the FJS with a 15-mm-diameter rod was measured, as shown in Fig. 12. Fig. 13 shows the results, which indicate that the reaction force can be controlled via the vacuum pressure inside the FJS. When an FJS with 96 piled fabrics (as opposed to 32 piled fabrics) was used in the force display, the reaction forces at a displacement of 10 mm were 26.8 and 1.5 N for inner pressures of 80 kPa and atmospheric pressure, respectively. A glove-type force display was developed using FJSs with 32 piled fabrics. Fig. 14 shows the schematic structure of the force display. The piled fabrics (each with a width of 15 mm and a length of 20 mm), which were serially connected using plastic plates, were sealed in an inner latex glove and were connected with an air tube. The inner glove was fixed with the outer latex glove where the piled fabrics were located on the digital joints. The stiffness and disturbed finger motion increased by evacuating the air inside the inner glove. Fig. 15 shows the reaction force of the index finger at different vacuum pressures when bending the finger, which is measured as a pressure on the finger pad using a pressure sensor (AMI3037-SB, AMI Techno co" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003738_1.b37415-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003738_1.b37415-Figure3-1.png", + "caption": "Fig. 3 DFIM structure.", + "texts": [ + " The DFIM (also called a woundrotor induction machine) has historically received limited use as an induction motor with variable rotor resistance, which was obtained by connecting the rotor windings to stationary, variable resistors. Developments in power electronics have made new modes of operation possible. The DFIM has been used especially as a variablespeed generator in wind turbine applications [12]. The advantage of the DFIM as a generator is that it can be designed and exploited in such away that the power electronics required to control the generator are relegated to the rotor windings, with a reduction of the size of the power converter. Figure 3 shows the structure of aDFIM.On the left is a side view of the machine and on the right is a front view (i.e., from the shaft) showing three-phase stator windings and three-phase rotor windings (in a simplified representation). The line-neutral voltages applied to the stator windings are labeled va;s, vb;s, and vc;s, whereas the line-neutral voltages applied to the rotor are labeled va;r, vb;r, and vc;r. The rotor voltages are applied through slip rings, as shown to the left of Fig. 3. The angle of the rotor winding a; r with respect to the stator winding a; s defines the rotor angular position, denoted \u03b8m. The angular velocity is denoted \u03c9m d\u03b8m\u2215dt. The shaft is connected to the load, in the case of a motor, or a prime mover in the case of a generator. In steady state, the stator voltages are sinusoidal variables with angular frequency \u03c9s, and can be represented by a single complex variable Vs (typically called a phasor) such that: va;s Re Vse j\u03b8s (1) vb;s Re Vse j\u03b8s\u2212j2\u03c0\u22153 (2) vc;s Re Vse j\u03b8s j2\u03c0\u22153 (3) where \u03b8s \u03c9st", + " The total active (PS) and reactive (QS) powers absorbed by the stator windings are given by: D ow nl oa de d by U N IV E R SI T Y O F N E W M E X IC O o n O ct ob er 2 6, 2 01 9 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /1 .B 37 41 5 Ps jQs 3 2 VsI s (4) where I S is the complex conjugate of the phasor for the stator current. The rotor voltages are sinusoidal variables described by phasors similarly to Eqs. (1\u20133), but with phasor Vr and angle \u03b8r \u03b8s \u2212 np\u03b8m, where nP is the number of pole pairs of themachine (nP is equal to 1 on Fig. 3). The angular electrical frequency of the rotor voltages is: \u03c9r d\u03b8r dt \u03c9s \u2212 np\u03c9m (5) One defines sn, the normalized or per-unit slip, as: sn \u03c9r \u03c9s (6) Note that the rotor electrical frequency, \u03c9r, and the slip can be negative, which corresponds to an angle \u03b8r for the rotor voltage rotating in the counterclockwise direction, that is, three-phase rotor voltages in reverse or backward sequence. Defining a rotor current phasor Ir, the total active and reactive powers absorbed by the rotor are defined as for the stator (4)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003552_1350650119866037-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003552_1350650119866037-Figure2-1.png", + "caption": "Figure 2. Schematic diagram of simplified contact between rough cylinders.", + "texts": [ + " As shown in Figure 1, z is the height of the asperity and d is the distance between the plane after contact deformation and the average height line of the asperities, also known as the contact distance ! \u00bc z d \u00f01\u00de The tooth contact can be equivalent to the contact of two cylinders by the meshing point, so the line contact between rough cylindrical surfaces is considered. According to the Hertz contact theory and the idea of Greenwood and Williamson,21 the contact of two rough cylindrical surfaces can be simplified into the contact between a smooth curved surface and a nominally flat rough plane. The contact diagram is shown in Figure 2. So, the contact distance between the two surfaces is u(x) u\u00f0x\u00de \u00bc d0 \u00fe x2 2R 2 E Z \u00fe1 1 p\u00f0s\u00de ln s x s ds \u00f02\u00de Here, d0 is the contact distance at the center point between two nominal curved surfaces, R is the combined radius of curvature at the meshing point, p is the contact pressure, and x is the distance between the contact point and the center of contact area in the direction perpendicular to the axial direction of the cylinder. When the length of the cylinder is in unit length, the relation between contact pressure p(x) and contact distance u(x) is p\u00f0x\u00de \u00bc 4 3 E 1=2 Z u\u00f0x\u00de\u00fe" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002291_ccsse.2017.8087894-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002291_ccsse.2017.8087894-Figure2-1.png", + "caption": "Figure 2. Data link module.", + "texts": [ + " The ground control station carries out wireless remote control to the UAV, so it requires higher real-time and high reliability of communication. The system uses two sets of data link communication, data link 1 main transmission demonstration start, end and emergency processing instructions to set the initial state of the experiment, through the data link 1 can guarantee the safe operation of the system. Data link 2 to form a mobile ad hoc network, the ground control station 2 for the nodes in the network, UAV formation through the ad hoc network can be command information and status information interaction, as shown in Figure 2 for the data link module. The data link 1 is built by the CUAV Supter radio. The digital transmission uses the industry level USB to TTL, the frequency is 915MHz. The reliability of its functions has been recognized by the industry and widely used in UAV communications, so the system chooses the data to build the data link. Data link 2 is based on the mesh networking technology developed YL-800 module, the frequency of 433MHz. In the course of the movement, the UAV needs to know the status information of other UAVs in the formation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002928_fuzz-ieee.2018.8491512-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002928_fuzz-ieee.2018.8491512-Figure1-1.png", + "caption": "Fig. 1. Tridimensional representation of the 3 DOF helicopter.", + "texts": [ + " This paper is organized as follows: the section II starts with the dynamic model of the helicopter and description of the parameters of the mathematical model, in the section III it is presented the design of adaptive fuzzy PID and in the section IV the validation of the control strategy, comparing the experimental data against the data obtained from the simulation, given a variable reference signal for the travel and elevation angles. II. MATHEMATICAL MODEL The free-body diagram of the 3 DOF helicopter is illustrated in Fig. 1. Using Euler-Lagrange formulation, the complete helicopter dynamics can be described by differential equations for the elevation angle (\u03f5), the pitch angle (\u03c1) and the travel angle (\u03c4). In the Fig. 1, and are the thrust of each one of the brushless-type DC motors located at the prototype. The helicopter rotates around the perpendicular axis (elevation angle). The forces acting on the system are the sum of the thrusts caused by the propellers powered by the two 978-1-5090-6020-7/18/$31.00 \u00a92018 IEEE brushless motors and the momentums that the counterweight and the main beam weight exercise. The differential equation that governs this movement is = ( + ) cos \u2212 \u2212 cos sin+ ( ) \u2212 (1) where ( ) = \u2212 ( \u2212 \u210e) sin cos + cos +( + cos ) cos + ( sin +\u210e) sin \u2212 cos (2) The main torque around the pitch axis is due to the difference of speed between the two motors, besides, when \u03c1\u22600, the gravitational force of the assembly also produces torque around the pitch axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003609_0954407019867491-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003609_0954407019867491-Figure6-1.png", + "caption": "Figure 6. The vehicle and the body-fixed coordinate frame in planar motion. Vx and Vy represent the longitudinal and lateral velocities, r is the yaw rate, b is the vehicle sideslip angle, and a and b represent the distance between the center of gravity and the front and rear axles, respectively.", + "texts": [ + " The other four measurements, on the other hand, namely Vfrom_fl,fr,rl,rr represent the vehicle speed, which are obtained using the speeds across wheel plane for the four corners. The expressions for each corner can be written as follows Vfrom fl= Vfl Vysin d\u00f0 \u00de r asin d\u00f0 \u00de+ t 2 cos d\u00f0 \u00de cos d\u00f0 \u00de Vfrom fr= Vfr Vysin d\u00f0 \u00de r asin d\u00f0 \u00de t 2 cos d\u00f0 \u00de cos d\u00f0 \u00de Vfrom rl =Vrl t 2 r Vfrom rr=Vrr+ t 2 r \u00f013\u00de where d is the steering wheel angle and Vy is the lateral speed, which is estimated separately (explained in VSC Simulations section). Note that Figure 6 representing the planar vehicle motion is used to obtain equation (13). Vfl, Vfr, Vrl, and Vrr in equation (13) are expressions of tire slip and wheel speeds Vi = rivi si +1 \u00f014\u00de where s is obtained using the slip-slope method; making use of the wheel dynamics equation _v= Ths Tb rFx Iw \u00f015\u00de where Ths is the half-shaft torque (for Architecture 3, it is the electric motor torque directly), Tb is the brake torque, obtained from pressure measurement, Fx is the longitudinal force, and Iw is the wheel inertia", + "m)/Vx; that is road adhesion limits, where m is the estimated adhesion coefficient, kus is the understeer coefficient, and is usually selected to be 0 \\ kus 1 to have a control that will yield a slightly understeer stable vehicle. Bayar et al.24 showed that by tuning the understeer coefficient appropriately, excessive sideslip angles can be avoided. Maximum vehicle sideslip angle: the maximum value of the angle that determines the stability and steerability of the vehicle that is intended to be kept in a certain band rather than regulating directly. It is the angle b shown in Figure 14, as well as in Figure 6. Deviation from the initial vehicle speed: the maximum deviation from the initial vehicle speed during the maneuver after activation of VSC. Maximum tire longitudinal slip: an indication of how much tire force is utilized to generate the required corrective yaw moment. Side-shaft angle of twist: the shafts\u2019 angle of twist (the shafts between the differential and the wheels for Architecture 1, the shafts between the axle motor and the wheels for Architecture 2) for the first and second drivetrain architectures", + " The first one is using the accelerometer and yaw rate signals, together with the estimated longitudinal speed _b= ay Vx r b ax Vx b tan b\u00f0 \u00der 1+b2 \u00f021\u00de The second one is first estimating the lateral forces, by properly filtering the expression below (recall Fx_estimated is obtained through equation (17)) Fy estimated =Fx estimated Csa Cxs \u00f022\u00de where a is the tire slip angle, Cs and Cx are the tire cornering and longitudinal stiffness values, respectively. Once Fy values are computed, lateral speed can be obtained through the following equation, which can be derived from Figure 6, neglecting roll motion of the vehicle Vy = \u00f0 P Fy M Vxr dt \u00f023\u00de This approach can be found in Tseng et al.\u2019s38 study as well. Next is the decision of fusing the two methods with each other, depending on tire slip condition. When slip is low, the weighting of equation (21) is increased, whereas during heavy braking which causes tire slip to increase, weighting of the above expression is increased. The reason of this is that in order for equation (22) to give accurate results, tire slip should not be at low values" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002793_ilt-07-2017-0208-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002793_ilt-07-2017-0208-Figure2-1.png", + "caption": "Figure 2 Two viewpoints to calculate the face contact loads in a dynamic model", + "texts": [ + " Fclosing is the total closing load at equilibrium consisting of the pressure drop and the spring support load. Msi (i.e. Msi = ksgg si) is a constant moment caused by the static stator tilt g si that is because of manufacturing tolerance and assembly imperfection. In equation (5), Msi is arbitrarily set to be only about the x axis. According to Section 2, to achieve a solution that includes face contact, the direct numerical dynamic model of a gas face seal should firstly calculate the contact pressure pc on each node based on its own elastic deformation (i.e. d 2 as shown in Figure 2) by the use of a certain parallel-plane asperity contact model (Greenwood and Williamson, 1966; Chang et al., 1987; Majumdar and Bhushan, 1991; Morag and Etsion, 2007). Then, it continues to integrate pc over the sealing area As referring to equation (4b) to obtain Fc, Mcx and Mcy. This computationally cumbersome approach has been adopted by Harp and Salant (1998), Green (2002) andHu et al. (2016b). However, the face-contact situation of a gas face seal is similar to the typical axial rub-impact of rotor dynamics (Yuan et al., 2007, 2008). As shown in Figure 2, in an axial rubimpact model of rotor dynamics, the rub-impact stiffness kri is obtained by experiments. The total normal contact load Fc is calculated from a single spot (i.e. Fc = krid 1). Then, the moments can be calculated according to phase angles. The elastic deformation d 1 is measured from the mean surface height zma. The rub-impact viewpoint in rotor dynamics is efficient but is out of the surface topography. Therefore, the following issue is how to introduce the axial rub-impact model of rotor dynamics into a gas-face-seal system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001082_jzus.a1001368-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001082_jzus.a1001368-Figure1-1.png", + "caption": "Fig. 1 shows the connective components of the UAV helicopter system used to collect flight data. This system contains three main parts: a hobby helicopter, an onboard system and a ground station. The R/C receiver receives the command signal from the", + "texts": [], + "surrounding_texts": [ + "For this study, initial experiments were conducted on a Hirobo Sceadu 50 radio-controlled (R/C) helicopter, which was first used as an experimental platform by Chen et al. (2006) for sensors configuration. Due to its payload capability and manoeuvrability, the upgrade of autopilot systems, sensors and communication devices is easy to perform with a UAV helicopter. The Hirobo Sceadu helicopter ac- tuation is performed by five onboard servo actuators. To execute all the swash plate movements (collective, aileron, and elevator), three of these servos are connected to the swash plate via the cyclic collective pitch mixing (CCPM) method. The centre piece of the helicopter onboard system includes: (1) a PC/104, used as data storage memory; (2) a inertial measurement unit (IMU), connected directly to the flight computer through a special serial port; (3) a diamond GPIO-MM counter/ timer card; and, (4) a hobby helicopter. Table 1 presents the specifications of this helicopter. transmitter and sends pulse width modulation (PWM) signals to the actuators. The PWM signals generated in R/C receiver are captured by the counter/timer device. The IMU sends the airframe signals which are accelerations (ax, ay, az), angular rates (p, q, r), and Euler angles (\u00d8, \u03b8, \u03c6). Signal data from both the counter/timer device and the IMU are recorded and stored in the compact flash (CF) card. Using a human operator to pilot the craft, a number of experimental flights were conducted to collect flight data. During the flight data test, the IMU sends the airframe data of the vehicle back to the PC/104, while the counter/timer captures the PWM signals generated by the R/C receiver. Fig. 2 shows the overall system identification procedure. The identification block consists of setting the hardware used to collect the flight data. To select the optimal hovering flight data, the helicopter was in stand-still hovering mode for several trials. Hovering flight data were selected based on the assumption that, in these conditions, the system is linear. Collected data, first, requires pre-processing to screen out noises. To perform this task, a moving average filter was used in this study. The NLARX model is selected as a mathematical black box model to be identified for an autonomous helicopter. To avoid excess computation at the overhead, we divided the model into three principal components: the longitudinal, lateral, and heave sub-models. Each of these sub-models is identified separately. The data contains 800 input-output data samples in near-hovering mode, generated at a sampling rate of 0.03 s. This data was then split into two subsets for estimating and validating the dynamics model, respectively. The control input vector contains the duty cycles of the five actuators, which are set up according to the swash plate layout. The output vector contains nine variables: Euler angles (\u00b0), angular rates (rad/s), and translation velocities (m/s) along the x, y, z axes." + ] + }, + { + "image_filename": "designv11_33_0000896_iros.2010.5650713-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000896_iros.2010.5650713-Figure1-1.png", + "caption": "Fig. 1. The robot platform for visual route following navigation: (a) A picture of the sensor equipped robot, (b) Top view illustration of the robot.", + "texts": [ + " Many visual route following navigation algorithms have been researched in outdoor environments, the reason is that outdoor navigation is a challenging work because the robot has to operate in various environment conditions. Detail and successful navigation results are presented in [7], however it does not consider obstacles in the environments. There is a slight mention about the obstacle avoidance problem in [8], but it also remains the deep A 978-1-4244-6676-4/10/$25.00 \u00a92010 IEEE 1799 implementation about obstacle avoidance as a future work. In this paper, the visual route following navigation system is constructed to have a steady performance whether it is in indoor or outdoor. As shown in Fig. 1, we equipped a four-wheeled skid steering mobile robot with three web cameras and a laser range finder. The two monocular web cameras are installed at both sides of the robot, respectively, and the rest camera is installed at the front of the robot. We extract path lines of the route using the side cameras and we obtain information about the path in front of the robot using the forward camera. Using the laser range finder which is attached behind the forward camera, we detect unknown obstacles up to 4m", + " Since, the robot continuously turns with \u03b8steer maintaining its velocity v, we assign the command which consists of \u03b8steer and v. The robot has a rotational velocity using assigned \u03b8steer: r steerw k \u03b8= \u22c5 (14) where wr is the robot\u2019s rotational velocity and k is the experimentally determined system gain that depends on the system\u2019s processing speed and the robot dynamics. Obstacle avoidance tests were conducted using a four-wheeled skid steering mobile robot, a Korean-made vehicle, with three monocular cameras and a laser range finder as shown in Fig. 1 (a). Three 40\u00b0 field of view Logitech QuickCam Sphere AF web cameras and a URG-04-LX-UG01 laser scanner by Hokuyo were mounted to the robot. We implemented the algorithm using MS Visual Studio program. Running time was measured using single-threaded execution on a 2.1GHz Core 2 Duo. Fig. 6 (b) represents the real experimental environments. The robot was moving with a constant velocity of 0.4m/s. Here, we used the obstacle as in Fig. 6 (a) and its radius r is 25cm. We set the width of the route WR as 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000509_rnc.2931-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000509_rnc.2931-Figure1-1.png", + "caption": "Figure 1. Positioning of encoder indices.", + "texts": [ + " The answer to this question will be presented in the following section. In classical discrete-time adaptive control, the dynamics Q kC1 in the closed-loop system (5) is derived by subtracting the parameter adaptation law (9) from the time invariance relationship .sk/D .sk 1/ for all k, which however does not hold for a varying .sk/. On the other hand, for the periodic parameter, we have .sk/D .sk L/ and that LD lN s where N is the number of encoder indices per revolution (or encoder resolution), as shown in Figure 1, s is the angle between each encoder index, and l is a positive constant such that lN is the period in terms of number of encoder indices. As such, we can modify the standard adaptation law (9), originally designed to Copyright \u00a9 2012 John Wiley & Sons, Ltd. DOI: 10.1002/rnc Int. J. Robust Nonlinear Control 2014; 24:1177\u20131188 update the parameter estimate between two consecutive time instances, namely from k 1 to k, into a new periodic adaptation law that updates the parameter estimate periodically after a time interval that is set based on the encoder indices counted from jj j lN to jj j and j 2 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002146_s1063785017070136-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002146_s1063785017070136-Figure1-1.png", + "caption": "Fig. 1. Measurement setup. Powder layer 1 on the substrate, melt track 2, illumination 3, laser radiation 4, laser head and scanning system 5, positioning system 6, highspeed video camera 7, macro objective 8, and optical filter 9 are indicated. The arrow denotes the direction of motion of the laser beam.", + "texts": [ + " The results of determination of the temperature field for a nickel-based alloy powder (Inconel 718) are presented. With this aim in view, the experimental setup for layer-by-layer laser alloying was fitted with a Bonito CL-400B/200 video camera mounted at an angle of ~60\u00b0 to the horizontal. The CMOS sensor had a spatial resolution of 7 \u03bcm with up to 256 shades of gray resolved in the visible range. An objective with a focal distance of 0.135 m was used in the 1 : 1 macro photography mode. The diagram of the experimental setup is presented in Fig. 1. It is known that the emission of a heated metal surface includes spectral luminescence lines of impurities (salts and metal oxides). A 6-mmthick BGG-22 colored glass filter secured in front of the camera objective was used to suppress the intensity 628 TECHNICAL PHYSICS LETTERS Vol. 43 No. 7 2017 ZAVALOV et al. of luminescence lines that reduce the accuracy of measurement of the temperature of thermal emission. Figure 2 shows the dependence of the relative output signal level of an element of the camera image sensor on the brightness temperature" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003466_j.mechmachtheory.2019.06.017-Figure32-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003466_j.mechmachtheory.2019.06.017-Figure32-1.png", + "caption": "Fig. 32. Contact stress on the face-gear surface at a contact position in case (III).", + "texts": [ + " The stress analysis is performed by applying a general-purpose finite element analysis (FEA) application ABAQUS?c? c . The boundary conditions of the face-gear drive finite element models are set as the same in [11] . The FEM model is composed of 296,400 nodes and 232,288 elements. The Poisson\u2019s ratio and Young\u2019s modulus of the material are set to be 0.29 and 2.068 \u00d7 10 5 MPa, respectively. A torque of 1600 Nm is loaded on the coupling point of the tapered involute pinion. The contact stress on the face-gear surfaces in case (III) is presented in Fig. 32 . The distribution of the contact points is orientated longitudinally, showing good accordance with the TCA results. The major axis of contact ellipse is longer in comparison with the face-gear drive in case (I). The variations of the contact and bending stresses are illustrated in Fig. 33 . The maximum contact stress on the facegear surface decreases in case (II) \u223c(IV) compared with the case (I). Because the high contact stress function occurs in case (I) disappear, and edge contact is avoided" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001589_1.4003148-Figure12-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001589_1.4003148-Figure12-1.png", + "caption": "Fig. 12 The pressure distributions of fluid film of a Rev-EGJB and of a Rev-HGJB under the operating conditions in Table 2 \u201e\u03b5=0.3\u2026", + "texts": [ + " In sumary, this paper concludes that a greater load capacity of the ev-EGJB than that of the Rev-HGJB can be gained no matter here the grooves are located in the bearing or journal. Furtherore, adopting the elliptical grooves to increase the load capacity s beneficial at a high length-diameter ratio. ilm. To examine how elliptical grooves benefit load characteris- ics, this study investigates the differences in the pressure distriutions of fluid film on a Rev-HGJB and Rev-EGJB with the ame operating conditions, as shown in Table 2. Figure 12 dislays the pressure distributions of the fluid film. Although the eak pressure in the Rev-HGJB is higher than that in the RevGJB, a larger portion of high pressure is achieved in the ressure-generated region of the Rev-EGJB than that of the Rev- able 2 Parameters of the Rev-EGJB \u201egrooved member otation\u2026 learance c 6 m ength L 0.008 m adius r 0.002 m luid viscosity 0.00124 Pa s o. of grooves 8 roove elliptical axis ratio for Rev-EGJB groove angle 0 deg for Rev-HGJB 1.732 roove depth ratio 1.0 roove width ratio 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000303_ichr.2010.5686287-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000303_ichr.2010.5686287-Figure2-1.png", + "caption": "Fig. 2. Kinematic model of the robot. All joint angles are shown in the zero position, and positive rotation is defined in a right-hand sense. The origin is located at the center of the shoulder.", + "texts": [ + " The second relation shows that the desired secondary impedance task in (2) is implemented with a minimum-error projection into the collective null space. The free DOFs from the object are thus shared amongst the manipulators for the secondary task. We developed a fully dynamic simulation to test the control law. The simulation consisted of two manipulator arms and a spatial object. The manipulators each modeled a humanoid arm with seven DOF and three links: an upper arm, lower arm, and hand. The kinematics of those arms are shown in Fig. 2. The contact constraints between the end-effectors and the object were enforced through springdamper forces. Each body in the simulation had a symmetric mass distribution, with the center of mass located at the center of the link. The physical properties are listed in Table I. For the six-DOF object, orientation was represented using xyz Euler angles. The orientation error was subsequently converted to an axis-angle representation for the sake of \u2206y in the control law [9]. The experiment implemented a two-hand grasp with the position-only task definition. The robot was asked to hold the position of the object fixed while achieving a desired configuration in the joint space. Since the object orientation was left free, the experiment was expected to rotate the object while minimizing the errors in the secondary impedance. The robot started out in the initial position shown in Fig. 2\u2014but with the lower arms pointing out horizontally (q4 = q11 = 90\u25e6). This placed the object at an initial orientation of (90, 0, 0) degs and an initial position of (0, 0.5,\u22120.4) m. In this position, we provided a step input to the joint space impedance commanding the right shoulder to swing forward (q1 = 30\u25e6). This command tended to pull the left elbow into the torso, and so we added a second command to keep the left elbow out (q9 = \u221210\u25e6). These commands were implemented through the stiffness of the secondary impedance, where the only non-zero elements of Kj were the diagonal elements corresponding to q1 and q9" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003222_ffe.12997-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003222_ffe.12997-Figure9-1.png", + "caption": "FIGURE 9 Visual representation of the discretized tooth flank with the corresponding Hertzian loading profile5,33", + "texts": [ + " Number of points into which tooth flank line is discretized is at the same time number of loading substeps each meshing cycle is divided into. Depth of the surface layer and the number of layers into which it is subdivided can be defined as well. Movement of the loading along the flank is performed automatically with the normal and the tangential load profiles being calculated separately for every individual loading substep throughout the meshing cycle. An example of the discretized tooth flank and the Hertzian load profile acting in a single point of the contact transformed onto actual geometry of the gear tooth is shown in Figure 9. In this particular sample, part of the flank line below kinematic diameter is divided into 21 segments of equal length and part of the flank above it is divided into 20 equal segments. The surface area is divided into 10 layers, and the subsurface area is divided into eight layers. For actual calculations, flank line and depth can be discretized in many more segments, ie, layers (flank line up to 1000 and more), depending on the size of the gear teeth and needed calculation precision. For the determination of fatigue lives, ie, the number of load reversals until crack initiation of components and elements operating in rolling\u2010sliding contact conditions, different model and criteria have been proposed in the literature with varying success" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000252_004051756503500607-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000252_004051756503500607-Figure4-1.png", + "caption": "Fig. 4. I\u2019rym\u00b7~~cl l\u2019quiB.dl\u2019nl ht\u2019l1dillg mod..!.", + "texts": [ + " As more cord length were extended, the additional increase in bending moment near the support would further reduce the size of the central portion or core of the cord in which there would be no relative fiber motion and a still further reduction in bending stiffness would be reitlize(l. Thus, it is proposed that an equivalent bending- l11(xlt:l, which has the capability both physically and analytically of describing- approximately the initial and straight line portion of the bending curves in Figure 2, could he represented as shown in Figure 4. In this equivalent mode), the cord is assumed to be circular in cross section. Over the length l&dquo;, the fibers are assuroe<1 to undergo no relative IllOtlOII and the cord behavl\u2019s as a solid circular rod. For values of > 1o, the central core (cross hatched on the figure) begins to diminish in size and continues to taper off toward the supported end. \u2019I\u2019he cross-sectional shape of this central infinite friction core will 1e assumed elliptical, with the major axis along the axis of bending. ", + " It will now be assumed that, at the core surface location y = ~t for any x > lo, the difference in these two strain com- ponents will be equal to their difference at x = lo, where relative fiber motion is first considered to occur. That is - Making use of Equations (7) and (8) and putting the strain components at x ia similar form gives By simplificatioo and rearranenment, Equation (10) can be reduced to the form This, then, gives the required relation Il = R(r) for / > x > lo and a deflection analysis of the equivalent model can now be performed. Applying strain energy principles and the. unit load method of deflection analysis [6], the end deflection of the cantilever beam in Figure 4 can t>e expressed in the form I . where .\u00a1 or The factor 0 which appears in Equation (12) is inserted to account for the twist of the structure. 1\u2019latt et al. [2~ have shown that the effect of twist in the bending of a yarn can be considered equiva!ent!y as a reduction in \u2019the extension modulus of the in R The main property which results from the orthogonal projection is: \u03b8\u0303 T p (t + 1)\u03b8\u0303p(t + 1) \u2264 \u03b8\u0303 T (t + 1)\u03b8\u0302(t + 1) (10.60) where: \u03b8\u0303p(t + 1) = \u03b8\u0302p(t + 1) \u2212 \u03b8 (10.61) \u03b8\u0303 (t + 1) = \u03b8\u0302 (t + 1) \u2212 \u03b8 (10.62) This means that the projected vector \u03b8\u0302p(t) is \u201ccloser\u201d to \u03b8 than \u03b8\u0302 (t), or at least at the same distance (in the sense of the Euclidian norm)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001381_kem.490.237-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001381_kem.490.237-Figure7-1.png", + "caption": "Fig. 7. Generating system of the pinion \u2013 cutting with helical motion", + "texts": [], + "surrounding_texts": [ + "Due to simultaneous cutting of the both sides of the tooth space (pinion and gear), the bearing contact bias occurs after assembling of a transmission made in that way. Such an effect is undesirable since it causes loud work of the transmission and non-uniform transfer of the motion which leads to the faster fatigue wear. In order to remove the bias the helical motion is applied during pinion finishing. It is realized by the axial offset of the fixed headstock of the workpiece connected with the generating gear rotation. The generating system of the pinion is the technological hypoid gear obtained thanks to the hypoid offset of the workpiece\u2019s axis by the E value with respect to the cradle\u2019s axis. The cradle with inclined toolhead (tilt) creates the bevel generating gear. For the pinion cutting the head\u2019s axis inclination is applied in order to compensate the difference between the pressure angle and tool profile so in the initial position the c S system is rotated with respect to the d S system by the angle of j around d X axis. The head\u2019s axis inclination angle (in degrees and minutes) and the value of the hypoid technological offset E is determined on the basis of the gear geometry analysis and calculation cards. Positive values mean down-shift by the pinion with left inclination line direction of a tooth or up-shift by the pinion with right inclination line direction of a tooth. The negative value means upshift by the pinion with left inclination line direction of a tooth or down-shift by the pinion with right inclination line direction of a tooth. Mathematical model of tooth flank surface Mathematical notation of the tooth flank is apparent from the equation of the surface of action of the tool, kinematics, and accepted treatment technological system. The following discussion presents the side of the tooth surface obtained with the use of a technological system with a bevel generating gear - this is the case more generally in comparison with a ring generating gear. While processing the envelope, the equation of the tooth flank, which is the bounding surface of the utility, is determined from the system of equations [1, 2, 6]. This system includes the equation of the family of tool surfaces and the equation of meshing, resulting from the method for determining the kinematic envelope: ( ) ( ) , , , , 0 t t t t t t s s \u03b8 \u03c8 \u03b8 \u03c8 \u22c5 = 1 t1 1 1 r n v (1) where: ( ), ,t t ts \u03b8 \u03c81r - determining the vector function of the family tool surfaces system bounded with treating pinion ( 1S ), 1 n - the unit normal vector defined in 1S , ( ), ,t t ts \u03b8 \u03c8t1 1v - the relative velocity vector defined in 1S . Based on the defined technological model, the family of tool surfaces is determined as follows: ( ) ( ) ( ), , ,t t t t t ts s\u03b8 \u03c8 \u03c8 \u03b8= \u22c51 1t tr M r (2) where: ( ),t ts \u03b8tr - vector equation of the tool surface referred to the system associated with the tool, tS , ,t ts \u03b8 - curvilinear coordinates surface form, ( )t\u03c81tM - the transformation matrix being the product of a transformation matrix representing the rotations and translations of homogeneous coordinate systems included in the technological gear model t\u03c8 - parameter of motion (in this case - angle of the cradle rotation). Vector equation of the tool surface as a function of curvilinear coordinates ,t ts \u03b8 shows the relationship (3), which involves the processing of the active side of the tooth. ( ) ( ) ( ) cos sin , sin sin cos t wk t wk t t t wk t wk t wk r s s r s s \u03b8 \u03b1 \u03b8 \u03b8 \u03b1 \u03b1 + = + \u2212 tr (3) where: wk\u03b1 - the angle of the external blades, wkr - the radius of the cutterhead. Based on the model of technological gear it is designated a family of the tool surfaces according to equation (2), for which it determines conversion matrix equation (4). ( ) ( )( ) ( )1t t t\u03c8 \u03c8 \u03c8 \u03c8= \u22c5 \u22c5 \u22c5 \u22c5 \u22c5 \u22c5 \u22c51t 1w wr rh hm mk kc cd dgM M M M M M M M M (4) where: ( )t\u03c81tM - conversion matrix, ij M - the elementary transformation matrices representing the rotations and translations of homogeneous coordinate systems, ij - subscript indicating the direction of the transformation from system jS to system iS . Technological gear model is also used to determine (in the chosen system, for example 1(S or )mS a unit normal vector and relative velocity. The normal vector to the surface of the tool sets at any of the predefined layouts. In this way we can get the components of meshing equation. Solving equations (1) by eliminating one of the variables such as: ts from the meshing equation and then substituting into the equation of the family of tool surfaces ( )1 , ,t t ts \u03b8 \u03c8r we can obtain the tooth flank surface equation in two-parametric form (5). ( ) ( )( ), , , ,t t t t t t ts\u03b8 \u03c8 \u03b8 \u03c8 \u03b8 \u03c8=1 1r r (5) where: ( ),t t\u03b8 \u03c81r - the equation of the pinion tooth surfaces in the two-parametric form, ( ),t t ts \u03b8 \u03c8 - the variable ts in the function of other parameters. Model of technological gear is designed to create the flank surface of the gear and pinion teeth, which will be used for the analysis of meshing for constructional spiral bevel gear. In order to obtain a tooth surface as the family of tool surfaces it should be designated in the system rigidly bounded with cutting pinion S1 , which represents the envelope to family of surfaces of the cut gear tooth ( )\u03a3 1 . Family of tool surfaces is shown in the following way: ( ( )) ( ( ( ) ( )) )\u03c8 \u03c8 \u03c8\u2190 \u2190 \u2190 \u2190 \u2190 \u2190 \u2190 \u2190 \u2190= \u22c5 \u22c5 \u22c5 \u2212 + \u22c5 \u22c5 \u22c5 \u2212 \u2212(t) (t) 1 1 w 1 t w r r h h m m k t c d d t t k c r hr L L L T L L L r T T (6) In this equation the vector record (t) 1 r , (t) m r , (t) gr , concerns a family tool surfaces ( )\u03a3 t (as evidenced by the superscript ( t )), set respectively in the cutting gear system S1 , the basic system (stationary) and the tool system. Other signs are bounded with the transformations of coordinate systems used in the model of technological gear. The applied rotation and translation matrixes of the coordinate system are as follows: cos i 0 sin i 0 1 0 sin i 0 cos i , \u2190 = \u2212 d g L (7) cos j sin j 0 sin j cos j 0 0 0 1 , \u2190 = \u2212 c d L (8) U cos q U sin q 0 , \u2190 \u2212 \u22c5 = \u2212 \u22c5 k c T (9) cos sin 0 ( ) sin cos 0 0 0 1 , \u03c8 \u03c8 \u03c8 \u03c8 \u03c8\u2190 \u2212 = t t m k t t t L (10) 1 B1 1 0 A X p ( ) ,\u03c8 \u2190 = \u2212 + h m t" + ] + }, + { + "image_filename": "designv11_33_0001215_s00521-012-1224-7-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001215_s00521-012-1224-7-Figure1-1.png", + "caption": "Fig. 1 The transelevator", + "texts": [ + " As the third step, the Lagrangian is obtained as follows: L \u00bc Xn i\u00bc1 \u00f0Ki Vi\u00de \u00f03\u00de As the fourth step, the following equation is used to obtain each dynamic equation for the mathematical model: d dt oL oq i oL oqi \u00bc si \u00f04\u00de Remark 1 The method to obtain the mathematical model of this paper is different to others because they use Jacobians, inertia tensors, or Christoffel symbols while in this paper none of these methods is used. 3.1 Structural mathematical model A transelevator is a crane similar to a Cartesian robotic arm which moves inside of a warehouse. This transelevator is used to move loads from one place to another one inside of the warehouse, for example, from the floor to the deposit, from the deposit to the floor, from one deposit to another one, or from a trailer to a deposit. Figure 1 shows the transelevator. From the Fig. 1, the parameters X1; Y1; and Z1 of the Eq. (1) are: X1 \u00bc x1 Y1 \u00bc 0 Z1 \u00bc 0 \u00f05\u00de where x1 is the length to the mass center of the link 1. From the Fig. 1, the parameter x1 is bounded as follows 0 B x1 B xmax. Using (1) and (5), the kinetic energy for the joint 1 is: K1 \u00bc 1 2 m1x 2 1 \u00f06\u00de From the Fig. 1, the parameters x2, y2, and z2 of the Eq. (1) are: X2 \u00bc x1 Y2 \u00bc 0 Z2 \u00bc z2 \u00f07\u00de where z2 is the length to the mass center of the link 2. From the Fig. 1 , the parameter z2 is bounded as follows 0 B z2 B zmax. Using (1) and (7), the kinetic energy for the joint 2 is: K2 \u00bc 1 2 m2x 2 1 \u00fe 1 2 m2z 2 2 \u00f08\u00de From the Fig. 1, the parameters X3; Y3 and Z3 of the Eq. (1) are: X3 \u00bc x1 Y3 \u00bc y3 Z3 \u00bc z2 \u00f09\u00de where y3 is the length to the mass center of the link 3. From the Fig. 1, the parameter y3 is bounded as follows 0 B y3 B ymax. Using (1) and (9), the kinetic energy for the joint 3 is: K3 \u00bc 1 2 m3x 2 1 \u00fe 1 2 m3z 2 2 \u00fe 1 2 m3y 2 3 \u00f010\u00de From the Fig. 1, the potential energy for the three links is as follows: V1 \u00bc 0 V2 \u00bc m2gz2 V3 \u00bc m3gz2 \u00f011\u00de Substituting the Eqs. (6), (8), (10), and (11) in (3) gives the Lagrangian as follows: L \u00bc 1 2 m13x 2 1 \u00fe 1 2 m23z 2 2 \u00fe 1 2 m3y 2 3 m23gz2 \u00f012\u00de where m13 = m1 ? m2 ? m3 and m23 = m2 ? m3. Using the Euler Lagrange method of Eq. (4) in (12) for the joint 1 gives: m13x 1 \u00bc s1 \u00f013\u00de where s1 is the force needed to move the link 1. Using the Euler Lagrange method of Eq. (4) in (12) for the joint 2 gives: m23z 2 \u00fe m23g \u00bc s2 \u00f014\u00de where s2 is the force needed to move the link 2", + " (27), (29) (30), (31), (33), (34), (35), (37), and (38) yields: _X1 \u00bc X2 \u00f040\u00de _X2 \u00bc km1 m13 X3 \u00f041\u00de _X3 \u00bc kb1h1 max L1xmax X2 Rm1 \u00fe Rc1\u00f0 \u00de L1 X3 \u00fe 1 L1 u1 \u00f042\u00de _X4 \u00bc X5 \u00f043\u00de _X5 \u00bc km2 m23 X6 g \u00f044\u00de _X6 \u00bc kb2h2 max L2zmax X5 Rm2 \u00fe Rc2\u00f0 \u00de L2 X6 \u00fe 1 L2 u2 \u00f045\u00de _X7 \u00bc X8 \u00f046\u00de _X8 \u00bc km3 m3 X9 \u00f047\u00de _X9 \u00bc kb3h3 max L2ymax X8 Rm3 \u00fe Rc3\u00f0 \u00de L3 X9 \u00fe 1 L3 u3 \u00f048\u00de Y1 \u00bc Vo1 max h1 max X1 \u00f049\u00de Y2 \u00bc Vo2 max h2 max X4 \u00f050\u00de Y3 \u00bc Vo3 max h3 max X7 \u00f051\u00de Equations (40)\u2013(51) describe the dynamic behavior of the transelevator crane with sensor and actuator. To validate the mathematical model, the simulation results will be compared with the experimental results. The simulations are obtained using the Matlab, the experimental results are obtained using the prototype of Fig. 1. Figure 5 shows the sensor and actuator which are connected to the prototype of Fig. 1 to obtain the experimental results. The mathematical model of the transelevator with sensor and actuator is described by Eqs. (40)\u2013(51). Table 1 shows the parameters used in the simulation results and the experimental results. In this section, the root mean square error (RMSE) for the inputs and outputs [15] is used to find the effectiveness of the proposed mathematical model, it is given as: RMSE \u00bc 1 T ZT 0 e2ds 0 @ 1 A 1 2 \u00f052\u00de where e2 = eu 2 = (U1r - U1)2 ? (U2r - U2)2 ? (U3r - U3)2, or e2 = ey 2 = (Y1r - Y1)2 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001163_andescon.2010.5633569-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001163_andescon.2010.5633569-Figure1-1.png", + "caption": "Fig. 1. Above: Illustration of a home made mobile robot \u201cRoverto\u201d; Below: Car kinematic parameters with asynchronous in-wheel motors.", + "texts": [ + ". INTRODUCTION In this paper we provide a solution for inverse and direct kinematics of a four-wheel drive skid-steering (4WDSS) mobile robot with asynchronous speed control in each wheel (see fig.1). Steering is based on controlling similarly to differential drive wheeled vehicles but with an asynchronous modality. Since all wheels rolling direction is aligned with the longitudinal axis of the vehicle, turning requires wheel slippage. The proposed mathematical framework is a velocitybased model in continuous time that provides an explicit functional form to calculate the robot\u2019s rotational heading angle rate based on the robotic platform geometry only, unlike other common approaches [12], [13], [14]", + " On the classical differential drive control model, only two different wheel speeds are considered for the calculations. When there are more than two wheels, the wheels of each side are synchronized, so there are yet only two different speeds, the left side, and the right side speeds. The instantaneous linear velocity vt is defined by the averaged tangential wheels speed vt = r\u03d5\u0307i t where the nominal wheels radius r is assumed to be equal for all wheels. Each wheel speed is indexed by i = {1, \u00b7 \u00b7 \u00b7 , 4}, ordered as depicted in fig.1. vt = r 4 \u2211 i \u03d5\u0307i t (1) We might approach to a differential driven locomotion system by defining both robot speed-sides with opposite signs, positive when clockwise, and minus when counter clockwise. Thus, the robot differential velocity is formulated as, v\u0302t = r(\u03d5\u03071 t + \u03d5\u03072 t \u2212 \u03d5\u03073 t \u2212 \u03d5\u03074 t ) (2) Rotational and linear speeds are given in terms of the wheels angular speeds and are directly controlled by the system. The robot\u2019s rotational speed is the rate of change of the heading angle \u03b8t with respect to an inertial frame of reference, the model quantifies how fast the angle changes it terms of time expressed by \u03c9t = d\u03b8/dt" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003636_s10846-019-01084-0-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003636_s10846-019-01084-0-Figure4-1.png", + "caption": "Fig. 4 The sample scenario of a mobile robot", + "texts": [ + " s = { 0 otherwise 1 s1 = 1 & s2 = 1 & s3 = 1 & s4 = 1 (17) A temporary target vector can be defined by the switching signal, which represents either the rotation vector or the target vector, as in Eq. 18. pi = sEi + (1 \u2212 s)vgi (18) The position of the real-time target point of the i-th robot can be obtained based on the temporary target vector, as in Eq. 19. pdi = [xdi, ydi]T = pi + \u0394pi (19) where \u0394pi = [\u0394pxi, \u0394pyi]T . The robots should move toward the real-time target points. We can get the desired angle of the robots based on the position of the real-time target points, as in Eq. 20. \u03b8di = atan2(\u0394pyi, \u0394pxi) (20) The sample scenario of a mobile robot is shown in Fig. 4. Firstly, the virtual center point of the robot is calculated, i.e., the point p\u0304o in Fig. 4. Secondly, we determine whether the robot needs to switch to the obstacle-avoidance mode. In this scenario, it is obvious that the robot needs to avoid the obstacle. Thirdly, the desired target point is obtained, i.e., the point pdo in Fig. 4, including direction and distance, need to be calculated based on the position of the robot and the virtual center point. At the end, the desired linear and angular velocity are obtained by calculation based on the desired target point. The desired linear and angular velocity can be derived as: vdi = x\u0307di cos(\u03b8di) + y\u0307di sin(\u03b8di) (21) wdi = \u03b8\u0307di (22) Through the Eqs. 21 and 22, we get: vdi = ( \u02d9\u0394pxi + x\u0307) cos(\u03b8di) + ( \u02d9\u0394pyi + y\u0307) sin(\u03b8di) (23) wdi = \u2212 \u0394pyi \u0394p2 xi + \u0394p2 yi \u02d9\u0394pxi + \u0394pxi \u0394p2 xi + \u0394p2 yi \u02d9\u0394pyi = \u2212 \u02d9\u0394pxi sin(\u03b8di) \u2212 \u02d9\u0394pyi cos(\u03b8di)\u221a \u0394p2 xi + \u0394p2 yi (24) Note that \u03b8ei is an angle difference between the direction and the desired angle of the i-th robot, as in Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003469_j.engfailanal.2019.06.024-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003469_j.engfailanal.2019.06.024-Figure4-1.png", + "caption": "Fig. 4. Bearing failure.", + "texts": [ + " When the electromagnetic clutch is electrified, it meshes with the deceleration device, and the power is transmitted to the eccentric mechanism after the deceleration, thus realizing the rotation of the inner and outer eccentric rings [5]. In order to test the performance of the wellbore trajectory control tool, our team developed several experiments on the dynamic performance test bench of downhole tools, as shown in Fig. 3. The cantilever bearing of borehole trajectory control devices failure twice during the experiments, as shown in Fig. 4. After disassembling the faulty bearing, we found that both the rolling body and the cage of the bearing were damaged. It can be seen that the cages are badly worn and deformed, and it has a certain degree of fracture, at the same time, there are obvious deformation caused by extrusion damage at one end of the cage. The other end of the cage has no obvious wear marks, and the wear and extrusion condition are obviously better than the first end. There are large number of needle rollers and metal crumbs on the inner and outer surfaces of the bearing, the needle is seriously detachedand is partly fractured with needle rollers" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000050_s0034-4877(09)00012-3-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000050_s0034-4877(09)00012-3-Figure1-1.png", + "caption": "Fig. 1. The Appel-Hamel dynamical system in an elevator.", + "texts": [ + " t qa qa j=I Like the holonomic constraint case, this equation depicts the noholonomic motion in a potential field. If the constraint is homogeneous in the relative velocities, i.e. f-.... afJ q'/ = f-.... afJ [q' _ v(O)(t)] = 0 ~a'/ a ~a' a a ' a=l qa a=l qa (20) then the virtual displacement equation (11) can be replaced by n af)L -.8qa = O. (21) a=1 aqa This is just the Chetaev condition. Thus we conclude that the Chetaev condition is only applicable to the constraint which is homogeneous in relative velocities. As an example, we consider the Appell-Hamel dynamical system in an elevator as depicted in Fig. 1: A block of mass m smoothly confined in a vertical tube is attached to the end of a thread that passes over two small pulleys, and then is wound around a drum of radius b. The drum is rigidly attached to a wheel of radius a which can roll on the horizontal plane of an elevator without slip. The legs of the frame support the pulleys and keep the wheel slide frictionless on the x'y'-plane. The mass of the large wheel is M whereas the masses of the pulleys and the frame are negligible. Let x, y, z be the coordinates of the block m in the rest frame, and e the angle of the wheel plane with respect to the x'-axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002503_1350650117753915-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002503_1350650117753915-Figure1-1.png", + "caption": "Figure 1. FRB structure diagram.", + "texts": [ + " Hydrodynamic lubrication, computational fluid dynamic, air entrainment, floating ring bearing Date received: 16 June 2017; accepted: 10 December 2017 Currently, turbochargers are popular in automobiles for increasing the power of internal combustion engines. The turbocharger shaft has traditionally been radially supported with a pair of floating ring bearings (FRBs). This bearing has been used for many years due to its robustness and reliability. In addition, its simplicity and ease of manufacturing, installation, and maintenance reduce system costs. Figure 1 illustrates the basic structure of an FRB. The FRB uses floating rings to create two fluid films in series, one between the shaft and floating ring and the other between the floating ring and main housing. In comparison with conventional single oil film bearings, FRBs have lower friction losses. Furthermore, FRBs show improved damping behavior based on the double oil film. The advantages of FRBs have attracted researchers to investigate its characteristics and to improve their performance.1\u20135 Of all the issues studied, two major problems concerned researchers the most: the decrease of ring-to-shaft speed ratio at high shaft speed and the characteristics of subsynchronous vibration" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003049_1.g003824-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003049_1.g003824-Figure2-1.png", + "caption": "Fig. 2 Local north\u2013east\u2013down reference frame.", + "texts": [ + " The geometric, inertial, propulsive, and aerodynamic properties of this aircraft are modeled using a combination of computational and experimental methods. In particular, the stability and control derivatives are initially estimated using the vortex lattice method and subsequently updated using system identification flight experiments [43]. Because this aircraft is assumed to be rigid, the pertinent states are the Euler angles (\u03d5, \u03b8, \u03c8), the angular velocity in the body axes (p, q, r), the airspeed in the body axes (u, v, w), and the position of the aircraft in a local north\u2013east\u2013down frame (pN , pE, pD) (see Fig. 2). The nonlinear equations of motion of rigid, fixed-wing aircraft are documented in several textbooks [44,45] and are thus not repeated here. To design controllers, these equations are linearized about a steady, wings-level, constant altitude flight condition, at a cruise airspeed of 15.4 m\u2215s. The throttle \u03b4t is normalized to the interval [0,1]. The left \u03b4l and the right \u03b4r elevons can each attain a physical deflection range of [\u221230, 20] deg, where positive values correspond to trailing-edge down deflections" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003893_0954406219888240-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003893_0954406219888240-Figure1-1.png", + "caption": "Figure 1. 3D model of the ideal spherical joint.", + "texts": [], + "surrounding_texts": [ + "The three-dimensional (3D) model of the ideal spherical joint and the combined spherical joint is shown in Figures 1 and 2, respectively. Usually, the workspace shape of the ideal spherical joint is a cone that can be simply determined by a constant angle, but it is not accurate for the combined spherical joint. If the spherical joint is assembled by a universal joint and a rotating unit, the workspace of the combined spherical joint is related to two Euler angles (yaw angle and pitch angle) of the universal joint and a roll angle of the rotating unit. The roll angle of the combined spherical joint is free, but the yaw angle and pitch angle are constrained by the non-interference domain of universal joint. Thus, the workspace of the universal joint is important for the combined spherical joint, which will be discussed first. Although the workspace analysis of this class of universal joints has been discussed,11 they are too complex to use. Inspired by the cross-sectional area method, a concise projection method is presented to study the joint workspace in this paper, so that only two parameters can determine the non-interference domain of this class universal joint." + ] + }, + { + "image_filename": "designv11_33_0001138_j.mechmachtheory.2011.07.015-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001138_j.mechmachtheory.2011.07.015-Figure5-1.png", + "caption": "Fig. 5. Coordinate systems between ZN worm-type hob cutter and RA worm gear.", + "texts": [ + " The results of this study can be applied to further investigations on tooth contact analysis, kinematic errors, contact paths and contact patterns of the semi RA, full RA and standard proportional tooth worm gear drives. Fig. 1 shows three types of worm gear drive meshing, semi RA, full RA and standard proportional tooth worm gear drives, with double-depth teeth at the same standard center distance. Fig. 1(a) shows the worm gear drive with standard proportional tooth Nomenclatures bx,bn width of hob cutter at varying pitch line, respectively (Fig. 3) C1 center distance of hob cutter and RA worm gear (Fig. 5) dx distance measured from the middle of hob cutter tooth height to the varying pitch line (Figs. 2 and 3(b)) h straight-lined edge height of hob cutter cutting blade (Fig. 3(b)) ht whole cutting blade height of hob cutter (Fig. 3(b)) Lij coordinate transformation matrix transforming from coordinate system Sj to Si (Eqs. (26), (27), (29), and (31)) l1 surface parameter of hob cutter (Fig. 3(b)) Mij homogeneous coordinate transformation matrix transforming from coordinate system Sj to Si (Eqs. (8), (9), (18), and (19)) mn normal modulus (Fig", + " 3) Si, Sj reference and rotational coordinate systems (i= f, g, p and j=c, 1, 2, 3 ) T1, T2 number of teeth of hob cutter and RA worm gear, respectively tc, tt transverse chordal thicknesses at pitch circle and throat circle of RA worm gear, respectively (Fig. 11) V12 (1) relative velocity vector of hob cutter and RA worm gear expressed in coordinate system S1 (Eqs. (22) and (33)) Vi (1) velocity vectors of hob cutter and RA worm gear (i =1, 2) expressed in coordinate system S1 (Eqs. (22)) \u03b11 pressure angle of hob cutter (Fig. 3(b)) \u03b31 cross angle of hob cutter in generating RA worm gear (Fig. 5) \u03b81 rotation angle of hob cutter in screw surface generation (Fig. 4) \u03bb1 lead angle of hob cutter (Fig. 4) \u03d51, \u03d52 rotational angles of hob cutter and RA worm gear, respectively (Fig. 5) \u03c91, \u03c92 angular velocities of hob cutter and RA worm gear, respectively (Eqs. (24), (25), (28) and (33)) \u03c9i (j) angular velocity vectors, expressed in coordinate system Sj (j=1, p) of hob cutter and RA worm gear (i=1, g) (Eqs. (23)\u2013(26) and (28)) meshing system. As contact progresses from point A (starting point of tooth engagement) to point P (pitch point), there is an approach action, while recess action occurs from point P to point C (final point of tooth engagement). It is noted that two parts of line of action for approach action and recess action are of equal contributions", + " (8), the circular tip surface equation of the ZN worm-type hob cutter R1 (c) can also be obtained as follows: R c\u00f0 \u00de 1 = cos \u03b81 rt + bn 2 tan\u03b11 + h 2 + rc cos \u03b1c\u2212 sin \u03b11\u00f0 \u00de \u2213 sin \u03bb1 sin \u03b81 bn 2 + h tan\u03b11 2 + rc cos \u03b11\u2212sin \u03b1c\u00f0 \u00de \u2212 sin \u03b81 rt + bn 2 tan \u03b11 + h 2 + rc cos \u03b1c\u2212sin \u03b11\u00f0 \u00de \u2213 sin \u03bb1 cos \u03b81 bn 2 + h tan\u03b11 2 + rc cos \u03b11\u2212sin \u03b1c\u00f0 \u00de cos\u03bb1 bn 2 + htan\u03b11 2 + rc cos\u03b11\u2212sin\u03b1c\u00f0 \u00de \u2212p1\u03b811 2 6666666666666666664 3 7777777777777777775 : \u00f011\u00de The normal vector of the hob cutter straight-lined edge surface, represented in the coordinate systemS1, can be obtained by: N1 = \u2202R1 \u2202l1 \u00d7 \u2202R1 \u2202\u03b81 ; \u00f012\u00de where \u2202R1 \u2202l1 = cos \u03b11 cos \u03b81 \u2213 sin \u03b11 sin \u03bb1 sin \u03b81 \u2212cos \u03b11 sin \u03b81 \u2213 sin \u03b11 sin \u03bb1 cos \u03b81 sin \u03b11 cos \u03bb1 2 4 3 5; \u00f013\u00de and \u2202R1 \u2202\u03b81 = \u2212 rt + l1 cos \u03b11\u00f0 \u00desin\u03b81 \u2213 l1 sin \u03bb1 sin \u03b11 cos\u03b81 \u2212 rt + l1 cos \u03b11\u00f0 \u00decos \u03b81 l1 sin \u03bb1 sin \u03b11 sin\u03b81 \u2212p1 2 4 3 5: \u00f014\u00de Similarly, the normal vector of the hob cutter circular tip surface, also represented in the coordinate system S1, can be obtained by: N c\u00f0 \u00de 1 = \u2202R c\u00f0 \u00de 1 \u2202\u03b1c \u00d7 \u2202R c\u00f0 \u00de 1 \u2202\u03b81 ; \u00f015\u00de where \u2202R c\u00f0 \u00de 1 \u2202\u03b1c = rc \u2212sin \u03b1c cos \u03b81 sin \u03bb1 cos \u03b1c sin \u03b81\u00f0 \u00de rc sin \u03b1c sin \u03b81 sin \u03bb1 cos \u03b1c cos \u03b81\u00f0 \u00de \u2212rc cos \u03bb1 cos \u03b1c 2 4 3 5; \u00f016\u00de and \u2202R c\u00f0 \u00de 1 \u2202\u03b81 = \u2212sin \u03b81 rt + bn 2 tan \u03b11 + h 2 + rc cos \u03b1c \u2212 sin \u03b11\u00f0 \u00de \u2213 sin \u03bb1cos \u03b81 bn 2 + h tan \u03b11 2 + rc cos \u03b11\u2212 sin \u03b1c\u00f0 \u00de \u2212 cos \u03b81 rt + bn 2 tan \u03b11 + h 2 + rc cos \u03b1c\u2212 sin \u03b11\u00f0 \u00de sin \u03bb1 sin \u03b81 bn 2 + h tan\u03b11 2 + rc cos \u03b11\u2212sin \u03b1c\u00f0 \u00de \u2212p1 2 6666666666666666664 3 7777777777777777775 : \u00f017\u00de Fig. 5 shows the schematic generating mechanism of hob cutter and RA worm gear. Coordinate systems Sp(Xp, Yp, Zp)and Sg(Xg, Yg, Zg) are reference coordinate systems for the hob cutter and generated RA worm gear, respectively. Coordinate system S1(X1, Y1, Z1) of the hob cutter rotates through an angle \u03d51 counterclockwise with respect to the reference coordinate system Sp. Similarly, coordinate system S2(X2, Y2, Z2) of the generated RA worm gear rotates through an angle \u03d52 counterclockwise with respect to the reference coordinate system Sg", + " the relative velocity of the generated RA worm gear with respect to the hob cutter is perpendicular to their common normal vector N1 at any cutting instant. Therefore, the equation of meshing of the hob cutter and generated RA worm gear can be expressed as follows [12]: N1\u2022V 1\u00f0 \u00de 12 = N1\u2022 V 1\u00f0 \u00de 1 \u2212V 1\u00f0 \u00de 2 = 0; \u00f022\u00de where V1 (1) and V2 (1) denote the velocities of the hob cutter and generated RA worm gear, respectively, and superscript \u201c (1) \u201d indicates the velocities are represented in coordinate system S1. According to Fig. 5, the relative velocity of the generated RAworm gearwith respect to the hob cutter represented in coordinate system S1, can be obtained by: V 1\u00f0 \u00de 12 = \u03c9 1\u00f0 \u00de 1 \u2212\u03c9 1\u00f0 \u00de 2 \u00d7 R1\u2212O1O2 1\u00f0 \u00de \u00d7 \u03c9 1\u00f0 \u00de 2 ; \u00f023\u00de where \u03c91 (1) and \u03c92 (1) are the angular velocities of the hob cutter and the generated RA worm gear, respectively, and they can be expressed in their respectively coordinate systems S1 and Sg as follows: \u03c9 1\u00f0 \u00de 1 = 0 0 1\u00bd T\u03c91; \u00f024\u00de and \u03c9 g\u00f0 \u00de 2 = 0 0 1\u00bd T\u03c92: \u00f025\u00de Angular velocity \u03c92 (g)can be also represented in coordinate system S1, by applying the following coordinate transformation matrix equation: \u03c9 1\u00f0 \u00de 2 = L1pLpg\u03c9 g\u00f0 \u00de 2 = L1g\u03c9 g\u00f0 \u00de 2 ; \u00f026\u00de where L1g = cos 1 cos \u03b31 sin 1 \u2212 sin \u03b31 sin 1 \u2212 sin 1 cos \u03b31 cos 1 \u2212 sin \u03b31 cos 1 0 sin \u03b31 cos \u03b31 2 4 3 5: \u00f027\u00de Substituting Eqs. (25) and (27) into Eq. (26), the angular velocity of the generated RA worm gear, represented in coordinate system S1, can be obtained by: \u03c9 1\u00f0 \u00de 2 = \u03c91 \u2212m21 sin \u03b31 sin 1 \u2212m21 sin \u03b31 cos 1 m21 cos \u03b31 2 4 3 5; \u00f028\u00de where m21=\u03d52/\u03d51 is the angular velocity ratio of generated RA worm gear to hob cutter. Similarly, according to Fig. 5, vector O1O2 can be obtained and expressed in coordinate system S1 by O1O2 1\u00f0 \u00de = L1pO1O2 p\u00f0 \u00de ; \u00f029\u00de where O1O2 p\u00f0 \u00de = C1 0 0\u00bd T ; \u00f030\u00de and L1p = cos 1 sin 1 0 \u2212sin 1 cos 1 0 0 0 1 2 4 3 5: \u00f031\u00de Substituting Eqs. (30) and (31) into Eq. (29), vector O1O2 is represented in coordinate system S1 as follows: O1O2 1\u00f0 \u00de = cos 1 C1 \u2212 sin 1C1 0 2 4 3 5: \u00f032\u00de Again, substituting Eqs. (10), (11), (24), (28), and (32) into Eq. (23) yields: V 1\u00f0 \u00de 12 = \u03c91 m21 cos \u03b31\u2212 1\u00f0 \u00deY1 + m21 sin \u03b31 cos 1 Z1 + cos \u03b31 sin 1 C1\u00f0 \u00de \u2212 m21 cos \u03b31\u2212 1\u00f0 \u00deX1 + m21 \u2212 sin \u03b31 sin 1 Z1 + cos \u03b31 cos 1 C1\u00f0 \u00de m21 sin \u03b31 \u2212 cos 1 X1 + sin 1 Y1 + C1\u00f0 \u00de 2 4 3 5: \u00f033\u00de Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001346_j.proeng.2012.04.091-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001346_j.proeng.2012.04.091-Figure1-1.png", + "caption": "Fig. 1. Recurve bow and terminology", + "texts": [ + " A \u201cleave-one-out\u201d cross validation procedure revealed consistent estimate of the model (all corrected R\u00b2 were in a range of 0.66 and 0.75 with p<0.03). It has been shown that a highly precise timing in arrow release in terms of small CVs of clicker time is important for high mean scoring. \u00a9 2012 Published by Elsevier Ltd. Keywords: Recurve archery; arrow release; measurement system; clicker time; motor program * Corresponding author. Tel.: +43-1-4277-48886; fax: +43-1-4277-48889. E-mail address: mario.heller@univie.ac.at. The process of shooting with a recurve bow (see Fig. 1) can be described as follows (see Fig. 2): The archer draws the bow, pulls the arrow to the clicker, fixes in this position and aims. Then he/she pulls the arrow through the clicker (the so called \u201cfinal pull\u201d) and releases the shot. From a biomechanical point of view, the archer has to cope with the breakdown of the static balance of forces between the external tension and his/her muscular forces at the time of shooting [1]. The final pull and the release of the shot have been of interest in some earlier studies" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000945_ssd.2010.5585533-Figure17-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000945_ssd.2010.5585533-Figure17-1.png", + "caption": "Figure 17. 3-D flux path trough the south claw, the mag netic ring and the collector.", + "texts": [ + " These are described as follows: \u2022 axially in the stator yoke (see figure 12), \u2022 radially down through the collector and crossing the air gap down to the magnetic ring (see figure 12), \u2022 radially then axially in the magnetic ring (see figure 12), \u2022 axially-radially in the claws (see figure 13), \u2022 radially up in air gap facing the stator laminations (see figure 14), \u2022 radially up in the stator teeth and circumferentially in the stator core back (see figure 15), \u2022 radially down in stator teeth facing the two adjacent claws (see figure 16), \u2022 radially-axially in the adjacent claws (see figure 16), \u2022 axially then radially in the magnetic ring (see figure 17), \u2022 radially up through the air gap and the mag netic collector on the other side of the machine (see figure 17), \u2022 axially in the other side of the stator yoke (see figure 18). 3.2.2. Leakage Flux Paths Not all the flux produced by the excitation winding and the armature contributes to the EMF generation. Different leakage fluxes have been distinguished in the CPAES. The two main ones are \u2022 the leakage flux linking adjacent claws, \u2022 tow dimensional flux paths which flow through the magnetic circuit as follows: - axially in the stator yoke, - radially down through the magnetic collector and air gap el, - radially then axially in the magnetic rings holding the claws, - axially in the claw, - radially in air gap e2 and in the stator teeth" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000305_robot.2009.5152486-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000305_robot.2009.5152486-Figure1-1.png", + "caption": "Fig. 1. A typical Stewart platform for active vibration isolation.", + "texts": [ + " The satisfactory vibration isolation performance can be achieved in all six DOFs. The rest of this paper is organized as follows. Section II introduces the dynamics of the Stewart platform driven by the linear voice coil motors for vibration isolation. Section III addresses the decoupling algorithm. Section IV discusses the controller design procedure and stability analysis. Simulation results are given in Section V to validate the effectiveness of the controller. Section VI concludes this paper. VIBRATION ISOLATION A typical Stewart platform, as shown in Fig. 1, is described in detail in [4]. Three Cartesian reference frames {P}, {B} 978-1-4244-2789-5/09/$25.00 \u00a92009 IEEE 1788 and {U} are used here. {P} is a reference frame rigidly attached to the upper plate. {B} is a reference frame rigidly attached to the lower plate. {U} is a universal inertial reference frame. In this paper, based on the researches [6] and [11], the dynamic model of the Stewart platform for vibration isolation is derived by the Newton-Euler method as M\u03c7\u0308+C\u03c7\u0307+B\u03c7\u0307+K\u03c7+\u2206s = \u03c4+ws, (1) where \u03c7\u2208R6 is the pose vector of the payload in the universal frame {U}, \u03c7\u0307 and \u03c7\u0308 are the first and second derivatives of \u03c7, respectively; M, B, K\u2208R6\u00d76 denote the inertia matrix, damping matrix and stiffness matrix, respectively; C\u03c7\u0307 \u2208 R6 corresponds to the centripetal and Coriolis forces vector; \u2206s \u2208R6 is the term of unmatched uncertainties of the model, including parameter perturbation, unmodeled dynamics and frictions, etc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003864_s11012-019-01081-5-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003864_s11012-019-01081-5-Figure2-1.png", + "caption": "Fig. 2 Kinematic diagram of the generic kinematic chain [21]", + "texts": [ + "1 Bases of the generic kinematic chain In [21], the authors presented a systematic method to obtain several types of kinematic chains derived from one unique chain called the generic kinematic chain. It was demonstrated that this generic chain can be easily particularized to obtain various parallel manipulators commonly used in practice. Although the bases of the generic kinematic chain are explained in depth in [21], to provide an understanding of the analyses presented afterwards, in the current paper the most significant steps are summarized. The structure of the generic chain can be described as follows (see Fig. 2): \u2022 It is composed of only prismatic and revolute joints. These two types of kinematic joints are alternatively located along the serial chain. \u2022 The axes of two consecutive joints will be either parallel or perpendicular. \u2022 To achieve the required generality, the generic chain has six R and six P joints. \u2022 By restricting some of the geometric parameters of this generic chain, different configurations of 6-DOF kinematic chains are obtained. With respect to position and velocity problems, the approach initially defines the loop equation of the generic chain shown in Fig. 2, and from this, the velocity equation is obtained. The loop equation can be written as: rP \u00bc rA \u00fe a\u00fe b\u00fe c\u00fe d \u00fe e\u00fe f \u00fe rGP \u00f01\u00de By expressing each vector associated with a variable-length link as the product of its module (length) times the unit vector of the corresponding P joints, the following expression is obtained: rP \u00bc rA \u00fe a ua \u00fe b ub \u00fe c uc \u00fe d ud \u00fe e ue \u00fe f uf \u00fe rGP \u00f02\u00de Now, by taking the derivative with respect to time, the velocity equation is obtained: vP \u00bc _a ua \u00fe _b ub \u00fe _c uc \u00fe _d ud \u00fe _e ue \u00fe _f uf \u00fe xb b\u00fe xc c\u00fe xd d \u00fe xe e\u00fe xf f \u00fe x rGP \u00f03\u00de where vP is the linear velocity of the coupler point P. In Eq. (3), the angular velocity of each element and that of the moving platform are given by: xb \u00bc _a ua xc \u00bc xb \u00fe _b ub xd \u00bc xc \u00fe _c uc xe \u00bc xd \u00fe _u uu xf \u00bc xe \u00fe _h uh x \u00bc xf \u00fe _w uw \u00f04\u00de 3.2 From the generic chain to the 3-CPCR parallel manipulator The CPCR chain, shown in Fig. 3a, is easily derived by simply particularizing the following parameters of the generic chain depicted in Fig. 2: d \u00bc constant e \u00bc f \u00bc 0 u \u00bc h \u00bc w \u00bc 0 \u00f05\u00de Once the parameters of the CPCR chain have been determined, the whole manipulator can be modeled as shown in Fig. 3b by taking into account that the two first P and R joints form a C joint and that similarly the third P and the second R joints form another C joint. Based on the kinematic diagram in Fig. 3a, the loop equation is: rP \u00bc rA \u00fe a\u00fe b\u00fe c\u00fe d \u00fe rEP \u00f06\u00de Then, by differentiating with respect to time, the velocity equation becomes: vP \u00bc _a ua \u00fe _b ub \u00fe _c uc \u00fe xb b\u00fe xc c\u00fe xd d \u00fe x rEP \u00f07\u00de where xb \u00bc _a ua xc \u00bc xb xd \u00bc xc \u00fe _b ub x \u00bc xd \u00fe _c uc \u00f08\u00de Subsequently, the components of the unit vectors associated with the translational motion ua, ub, uc and the rotational motion ua, ub, uc are obtained" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000151_iemdc.2009.5075258-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000151_iemdc.2009.5075258-Figure2-1.png", + "caption": "Fig. 2. Mesh of the 3D geometry", + "texts": [ + " A single layer is forbidden by the most software products as a sliding surface technology is used to model the rotation. This technology requires two successive slip air layers. For more detailed calculations are multiple layers required. The width of six micrometers implies elements between a half and a micrometer side length in the air gap to achieve a good quality two layer mesh. A second aspect is the the dimension in axial direction. It\u2019s huge compared with the the small air gap. We tried a lot of different combinations of structured and automatic mesh. Fig. 2a shows an example how to minimize the elements. On the side of the air gap additional layers of iron have been introduced. This allows us to control the expansion of the mesh. We could significantly reduce the number of volume elements in the mesh. Optimizing the mesh with towards minimal number of elements (without surrounding air and air gap) we get a valid mesh with about 950.000 volume elements. But if we add the air gap to the mesh generator, building the mesh is a heavy task. The reason is the combination of structured and automatic boundary faces", + " It takes several days to mesh the complete geometry with about three million elements. Therefore we wanted to use the more progressive advancing front algorithm. This ones can handle combinations of structured and automatic mesh more easily. But the mesh definition becomes more difficult if one wants to save elements. Especially the use of various axial mesh densities for the extrusive mesh leads to ill connected elements at the boundary faces between different mesh. The algorithm is much faster but not as stable. An alternative is to take the entire machine in a 2D plane (see fig 2b), mesh it and then extrude it in one piece with extrusive mesh generator. For the generator itself it is more easy if you use more then one axial extrusion. This results in smaller volumes, which were easier to mesh (see fig 4). The big amount of very small elements in the air gap together with an average PC leads to inadmissible large meshing and computation durations. Therefore, the 3D model for the calculation of dynamic processes and optimizations is inappropriate. There can be no more than a reference calculation with this model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003921_b978-0-12-803581-8.11722-9-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003921_b978-0-12-803581-8.11722-9-Figure3-1.png", + "caption": "Fig. 3 Disk viscous brake with ER fluid: 1 \u2013 input shaft, 2 \u2013 housing, 3 \u2013 sealing ring, 4 \u2013 bearing, 5 \u2013 slip ring for supplying high voltage, 6 \u2013 insulator, 7 \u2013 disk of driving part, 8 \u2013 disk of immobile housing, 9 \u2013 bolt.", + "texts": [ + " The principle of functioning of hydraulic clutches and brakes with ER fluid is based on using a high voltage U to change intensity E of the electric field occurring between the electrodes connected to high voltage power supply\u2019s terminals, and at the same time, changing the shear stress t in the fluid during its flow through the gap between electrodes. Wherein, in viscous clutches and brakes, an increase in shear stress t causes an increase of the transferred torque M, while in hydrodynamic clutches an increase in shear stress t causes a decrease in the transferred torque M, due to a difference in types of ER fluid\u2019s flow in these two clutch types. Disk Viscous Brake With ER Fluid Disk viscous brake with ER fluid presented in Fig. 3 is attached to a horizontal frame with bolts (Olszak et al., 2019a). The outer diameter of the brake is 160 mm, and its length 60 mm. The brake consists of 6 disks attached to the input shaft, creating the propelling part, and 5 disks fixed to the immobile housing. Between the disks, spacer rings were placed to determine the width of the working gap. The \u201c\u00fe \u201d terminal of the high voltage power supply unit is connected to the clutch\u2019s shaft with a brush and a slip ring, while the \u201c\u2013\u201d terminal is connected directly to the immobile housing of the brake" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002503_1350650117753915-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002503_1350650117753915-Figure8-1.png", + "caption": "Figure 8. Transient air volume fraction distribution.", + "texts": [ + " Although this simplification might hide the fluctuations of the flow field caused by the rotation of the floating ring, this method is adopted considering the feasibility in practice. In addition, this simplification is considered a good approximation when considering vibration problems. The detailed results are analyzed in the following section. The transient calculation results demonstrate that the impact of the motion of the floating ring and shaft is significant. Figure 7 shows the pressure distribution at different time steps at 10 kr/min shaft speed, whereas Figure 8 shows the 10 kr/min shaft speed air entrainment condition. Given the transient effect, the position of the high pressure zone fluctuates with time. Air volume fraction also fluctuates with the fluctuation of the low pressure zone and with the effect of rotation. DFT is a signal processing method that converts time domain signals into frequency domain. This method is adopted to analyze the vibration frequency of the rotor using transient CFD results. The sampling interval is set to be the time step of the transient simulation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003001_access.2018.2883332-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003001_access.2018.2883332-Figure2-1.png", + "caption": "FIGURE 2. Gear and pinion subassembly with a relief chamber on the outside of the bearing block: \u00ac bearing blocks, spur gear pairs, \u00ae trapped volume and kidney ports, and \u00af relief chamber.", + "texts": [ + " As the fluid travels through the pump, the fluid in the inlet will be transported by a pair of rotating gear and pinon to the outlet in \u2018\u2018pockets\u2019\u2019. As the pockets of fluid travel to the meshing region of the two gears, an isolated space will be formed by the adjacent teeth and the hydraulic fluid inside is said to be trapped. Themaximum andminimum trapped volume in Fig. 1 suggest high pressure pulsation for fluid with relatively low compressibility. To reduce the magnitude of the pressure spike in the trapped volume, the author proposes a relief chamber, as shown in Fig. 2, a cavity machined on the opposite side of the bearing block that supports and seals the side of gears, connected to both trapped volume. As oppose to the fixed orifice position design in [7], the connection of the trapped volume to damper is achieved through two kidney ports, which not only consistently absorbs the cyclic pressure spikes from a single trapped volume, but can also take advantage 77510 VOLUME 6, 2018 of the compression/expansion phase offset between the two trapped volume for partial ripple cancellation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002713_978-3-319-74459-9_12-Figure12.3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002713_978-3-319-74459-9_12-Figure12.3-1.png", + "caption": "Fig. 12.3 Construction and operation of MDC", + "texts": [ + " Basically, the operation principle of MDCs is similar to electrosmosis and electrodialysis (Cao et al. 2009; Qu et al. 2012). In MDC, the anodic and cathodic compartments are separated by a desalination chamber and the electrochemical potential of the microorganisms is utilized to drive the transport of ions. The anode and desalination chambers are partitioned by an anion exchange membrane whereas the cathode and the desalination chamber are partitioned by a cation exchange membrane. The schematic diagram showing the principle of MDC is shown in Fig. 12.3. Microorganisms are utilized for the oxidation of electron donors in the anode compartment and reduction of electron acceptors in the cathode compartment, as in the case of MFCs. The potential gradient developed in this process is used for driving the transport of dissolved Na+ and Cl ions through the selective ion exchange membranes toward the cathode and anode, respectively. The larger the potential difference between the anode and the cathode, the higher is the rate of desalination. Hence, for any bioelectrochemical systems, the ideal choice of bioanode should have more negative anodic potential, and the biocathode should have more positive cathodic potential" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003881_012010-Figure12-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003881_012010-Figure12-1.png", + "caption": "Figure 12. The punch in the selected industrial case (Figure 5): (a) topology optimized using LS-TaSC in LS-DYNA and in compliance (stiffness) with the conventional punch, and (b) the 3D-printed topology optimized punch and the 3D-printed (with honeycomb inner structure) conventionally designed version of the same punch. Both versions are 3D-printed in maraging steel (DIN1.2709).", + "texts": [ + " The total costs (comprising material, machine, salary, and logistics costs) increase, as displayed in Table 5, from 26,000 to 31,000 SEK, in case the 3D-printing inclusive process is selected. The total cost of the 3D-printing inclusive process is based on the assumption that the depreciation period for the 3D printing machine is 5 years. A 10 years long depreciation period for the 3D printing machine reduces the total costs to 29000 SEK. The punch in the industrial case (Figures 5 and 11) was topology optimized using LS-TaSC in LSDYNA. This topology optimized punch was 3D-printed in maraging steel (DIN1.2709). Figure 12 displays the simulation results, the 3D-printed topology optimized punch, and the 3D-printed (with International Deep Drawing Research Group 38th Annual Conference IOP Conf. Series: Materials Science and Engineering 651 (2019) 012010 IOP Publishing doi:10.1088/1757-899X/651/1/012010 honeycomb inner structure) conventionally designed version of the same punch. Both of these punches have so far run more than 120 thousand strokes without any problems. Compared to a 3D-printed solid punch, topology optimization and a honeycomb inner structure improved the material usage (and thereby reduced the weight) and printing time by ca 45% & ca 34% respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001598_speedam.2010.5544842-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001598_speedam.2010.5544842-Figure1-1.png", + "caption": "Fig. 1: Flux density for 3-D FEM in (a) healthy motor and (b) motor with 40% eccentricity.", + "texts": [ + " Due to inherent eccentricity, more than 10% eccentricity is considered in this paper and caused by collision of the rotor and stator poles, the relative eccentricity of more than 40% is not considered in this study. IV. NUMERICAL ANALYSIS As mentioned before, to study the effects of dynamic eccentricity on the switched reluctance behavior, the motor is simulated utilizing 3-D finite element analysis. Distribution of flux density in the healthy motor and the motor with 40% dynamic eccentricity utilizing 3-D FE analysis are shown in Fig. 1. The reduction of air gap length and consequently reduction of its related magnetic reluctance causes an increase in the flux density. As shown, the magnitude of flux density obtained in the core of coil one in phase A in faulty motor, shows considerable increase when compared to the healthy motor. Fig. 2 shows the flux-linkage of coil one in phase A, utilizing 3-D FEM and the variation of rotor position in healthy motor as well as the motor with various dynamic eccentricities. As shown above the flux-linkage peaks at about 44 degrees, correspond to the rotor pole located completely aligned with the related stator pole" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003832_s11665-019-04449-6-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003832_s11665-019-04449-6-Figure1-1.png", + "caption": "Fig. 1 Schematic of specimen orientation with respect to material block", + "texts": [ + " The objective of this work is to study the anisotropy of fracture toughness and strength in mill annealed and beta annealed Ti-6Al-4V alloy in 10-cm (4-in.)-thick sections. All the samples in different orientations have the same thickness. This paper aimed to find the fracture toughness in the different orientations. These orientations were the L\u2013T, T\u2013L, and S\u2013L. L is the length parallel to the rolling direction, T is the transverse, and S is the thickness or short transverse direction. The designation has two letters; the first letter represents the direction of the load application and the second represents the direction of the crack growth. Figure 1 shows how specimen was extracted from the material block relative to the rolling direction. Tables 1 and 2 show the breakdown of the specimen tested. There were a total of 21 tensile tests and 23 fracture toughness tests. The material used for this study was Ti-6Al-4V received from RTI International (now Arconic). Provided by the supplier, the composition of the Ti-6Al-4V plate was 0.004 wt.% C, 0.006 wt.% N, 0.19 wt.% Fe, 6.48 wt.% Al, 4.04 wt.% V, 0.19 wt.% O, less than 50 ppm of Y, and with the remaining wt" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002168_ifsa-scis.2017.8023339-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002168_ifsa-scis.2017.8023339-Figure8-1.png", + "caption": "Fig. 8. Membership Functions (E2) Fig. 9. Membership Functions (E4)", + "texts": [], + "surrounding_texts": [ + "At first in this research, we measured various EMG data using several electrodes and examined the EMG electrode position which is effective for the instruction recognition of moving direction. In this section, we report the result to confir the effectiveness of a proposed method by comparing the moving time of path course and the questionnaire result in the control experiment of an omni-directional wheelchair by using the proposed method and the EMG data only." + ] + }, + { + "image_filename": "designv11_33_0001139_1.4004814-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001139_1.4004814-Figure1-1.png", + "caption": "Fig. 1 Physical model for ultrafast phase change", + "texts": [ + " In this paper, the irradiation of femtosecond pulsed laser on a single nanosized particle is simulated numerically. For the rapid melting and resolidification processes, nucleation dynamics is employed to delineate the proceeding of solid\u2013liquid interface. For the evaporation process, a thermodynamic model based on wave hypothesis is utilized. [27] The simulation focuses on the size effect of nanosized particles. Different particle size and laser fluence are used to study their effects on the phase change processes including melting, vaporization, and resolidification. Figure 1 shows the physical model under consideration. A laser pulse is deposited on a powder bed with initial temperature of Ti. Because of scattering, reflection, and penetration of the laser beams, it is assumed that at any vacant space in the powder bed, laser radiation distribute in all direction evenly. [25,26] For any 1Corresponding author. Contributed by the Heat Transfer Division of ASME for publication in the JOURNAL OF HEAT TRANSFER. Manuscript received February 12, 2011; final manuscript received July 26, 2011; published online November 21, 2011" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002225_rnc.3928-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002225_rnc.3928-Figure1-1.png", + "caption": "FIGURE 1 Three identical pendulums are connected by springs. Control torques u1 and u2 are applied to the first and third pendulums, and unknown disturbance torques d1, d2, and d3 are applied to all 3 pendulums. The objective is to control the angular displacements \u03b81 and \u03b83. In addition, Subsystem 1 control u1 cannot depend on Subsystem 2 states \u03b83 and \u03b8\u03073, and Subsystem 2 control u2 cannot depend on Subsystem 1 states \u03b81, \u03b82, \u03b8\u03071, and \u03b8\u03072", + "texts": [ + " Since Equations 10 to 12 satisfy assumptions (A1)-(A6), (A8), and (A9) in the work of Hoagg and Seigler14 and \u03b7k satisfies conditions (C1) to (C3) in the work of Hoagg and Seigler,14 it follows from theorem 2 part (ii) in the work of Hoagg and Seigler14 that there exist \u03bd\u03b5 > 0 and k\u03b5 > k0 such that for all [xT 0 \u03d5T d(0) \u03d5T \u03be (0)] T \u2208 B\u03bd\u03b5 and all k > k\u03b5, ||y \u2212 y\u2217||\u221e < \u03b52. Thus, for all [xT 0 \u03d5T d(0) \u03d5 T \u03be (0)] T \u2208 B\u03bd\u03b5 , all k > k\u03b5, and for all i \u2208 , ||yi \u2212 y\u2217,i||\u221e < ||Ei||\u03b52 < \u03b5, which confirms (ii). The following examples demonstrate D-FFL. The first example illustrates the local result Theorem 1, and the second example illustrates the global result Theorem 2. Example 1. The 3-pendulum system shown in Figure 1 consists of 2 nonlinear subsystems, which are denoted as Subsystem 1 and Subsystem 2. Each pendulum has mass m = 5 kg and length l = 75 cm, and we let \u03b81, \u03b82, and \u03b83 denote the angles of the 3 pendulums. The local outputs are y1 = \u03b81 and y2 = \u03b83. The local states are x1 = [\u03b81 \u03b82 \u03b8\u03071 \u03b8\u03072]T and x2 = [\u03b83 \u03b8\u03073]T. Furthermore, let d = [d1 d2 d3]T. The dynamics of Subsystem 1 are x\u03071 = \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 x1(3) x1(4) 3g l sin x1(1) \u2212 3ksp ml2 \u03b41 \ud835\udf15\u03b41 \ud835\udf15x1(1) 3g l sin x1(2) \u2212 3ksp ml2 ( \u03b41 \ud835\udf15\u03b41 \ud835\udf15x1(2) + \u03b42 \ud835\udf15\u03b42 \ud835\udf15x1(2) ) \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 + \u23a1\u23a2\u23a2\u23a2\u23a2\u23a3 0 0 3 ml2 0 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a6 u1 + \u23a1\u23a2\u23a2\u23a2\u23a2\u23a3 0 0 0 0 0 0 3 ml2 0 0 0 3 ml2 0 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a6 d, y1 = x1(1), where g = 9" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000247_aim.2010.5695728-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000247_aim.2010.5695728-Figure1-1.png", + "caption": "Fig. 1. Variables and parameters specific to the controlled plant.", + "texts": [ + " Common positive definite matrices are calculated for each T-S FC as solutions to appropriately defined linear matrix inequalities (LMIs) related to the FCS stability [19]. The main advantage of our approach is the relatively simple design method and controller structure that allow for the implementation of low-cost T-S FCs. The paper is organized as follows: Section II discusses the crisp and fuzzy mathematical modeling of the controlled plant, Section III is dedicated to the new T-S FCs, Section IV illustrates the real-time experimental results. The conclusions are presented in Section V. The INTECO ABS laboratory equipment consists of two wheels (Fig. 1). The lower wheel is accelerated to illustrate the car speed making the upper wheel will gain speed. Reaching the speed threshold value causes the upper wheel to initiate the braking sequence. Use is made of PWMcontrolled DC motors to drive the lower wheel and the brake on the upper wheel. Iron and plastic surfaces are forced to contact and cause the nonlinear slip-friction curve. The braking system contains a cable wound on the motor\u2019s shaft that acts upon a hand-brake mechanism with lever that determines the plates of the braking system to create T 978-1-4244-8030-2/10/$26" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002410_iris.2017.8250106-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002410_iris.2017.8250106-Figure1-1.png", + "caption": "Fig. 1. Robot kinematics.", + "texts": [ + " Assumption 5 is taken for the proposed CRE strategy based on frontier method using SRT presented in section III. Under these assumptions, the CRE of unknown environments for a team of robots, is explained in this paper. A more specific explanation will be presented in the following sections. In order to make a robot move in any direction, the turnand-go scheme is applied to the robot. A brief explanation of the robot kinematics is presented in this subsection. A k-th robot (k = 1, 2, \u00b7 \u00b7 \u00b7 , n) with radius of r has a position (xk, yk) \u2208 W and a heading angle, \u03b8k, as shown in Fig. 1. Each robot has two wheels and is equipped with a laser range scanners to measure its surrounding. The formulation of the robot kinematics is as follows: x\u0307k = Vc cos \u03b8k, (1) y\u0307k = Vc sin \u03b8k, (2) \u03b8\u0307k = \u03c9c, (3) where (xk, yk) is a position of k-th robot, and \u03b8k is a heading angle of k-th robot. In 1-3, the velocity and angular velocity of k-th robot, Vc and \u03c9c, can be written as Vc = Vl + Vr 2 , (4) \u03c9c = Vr + Vl D , (5) where D is robot diameter and Vl, Vr is left and right wheel speed, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002033_elma.2017.7955470-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002033_elma.2017.7955470-Figure3-1.png", + "caption": "Fig. 3. Regions of a coaxial magnetic gear", + "texts": [ + " The gear ratio of the second construction is G13=3, because p1=6 and p3=18. The dynamical conditions are these conditions in which the outer rotor rotates together with the inner rotor of the magnetic gear. In this state the dynamic torque of the magnetic gear is considered. In order to attain the highest torque density, the number of pole pairs on each of the rotors should satisfy the following relation [19]: .312 ppp The regions of interest in the magnetic gear\u2019s construction are represented in Fig.3. They are as follows: permanent magnet regions of the rotors, ferromagnetic parts, modulating steel segments, and both air gaps. Regions I, II, i, III, IV and V correspond to the permanent magnets mounted on the inner rotor, to the air gap between the inner rotor and the steel segments, to the steel segments, to the air gap between the steel segments and the outer rotor, and to the permanent magnets mounted on the outer rotor, respectively. Magnetic field distribution is described with the expression given below [20] ,2 M A A 0 dt d where 2 is Laplacian, A is the magnetic vector potential, \u03bc0 is the free space magnetic permeability (\u03bc0=4\u03c0\u00d710-7 H/m), \u03b3 is the electrical conductivity and M is the magnetization vector" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000209_tac.2009.2020643-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000209_tac.2009.2020643-Figure5-1.png", + "caption": "Fig. 5. .", + "texts": [ + " Since , we have for any control curve of (7). Therefore, by the similar method to Case 1, the criterion function changes its sign on any control curve . Thus, the system (7) is globally controllable by Proposition 3.2. We remark that in this subcase, the condition that is controllable (i.e. ) suffices for (7) being globally controllable. Subcase B: , . Without loss of generality, we assume that . By the similar method to Case 1, we can draw the trajectories of and the control curve in the phase plane as shown in Fig. 5. It is easy to know that the criterion function changes its sign on any control curve . Hence, the system (7) is globally controllable. Case 5: The eigenvalues of are two conjugate pure imaginary numbers. We assume , , , . For any control curve of (7), if does not pass through the origin, then there is a point in the outer side of (namely, the side which does not include the origin, see Definition 4.1 in Section IV). Since the trajectory of the system passing through is a circle and centered with the origin, must pass through twice at least as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001575_robot.2010.5509764-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001575_robot.2010.5509764-Figure6-1.png", + "caption": "Fig. 6 QuadHelix CAD prototype turning wheel movement", + "texts": [ + " The shaft needs only a small counter bearing (i), because the rope forces onto the shaft compensate themselves. All guidings and bearings are realized with high performance plastic bearings to reduce overall system weight. The elements for pre-tensioning the rope are not shown in this model. They are located between the pulleys and the turning wheel itself. These parts make up the first prototype of the QuadHelixDrive. The lightweight HMPE ropes for power transmission reduce the overall weight and thereby increase the massrelated torque-density. In Fig. 6 the movement of the actuator is shown while rotating the turning wheel. The motor unit turns the lower side DoHelix-coiling to rotate the turning wheel, while at the same time the upper side DoHelix-coiling is uncoiled through the rotation of the turning wheel. A first rough calculation of the available torque at the turning wheel axis starts with the shaft torque Mshaft. For that the nominal motor torque Mn is multiplied with the gearhead reduction Rgh and the maximum efficiency \u03b7gh of the gearhead" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001304_cjme.2012.05.947-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001304_cjme.2012.05.947-Figure3-1.png", + "caption": "Fig. 3. Machine with rotational head and table", + "texts": [ + " In the cutter coordinate system the cutter edges of the given cutter can be expressed as c c c c( , ) ( , ) ( , ) ( , ) ,k x k y k z k\u03b8 \u03b8 \u03b8 \u03b8 r i j k (4) and where \u03b8 is the angle between the positive Xc-axis and the line from the origin Oc to the projection of a given point on the cutter edge into the XcOcYc plane. Then it can be translated to the rotational coordinate system with the following expressions (5) CHINESE JOURNAL OF MECHANICAL ENGINEERING \u00b7949\u00b7 where 0 2( 1)\u03c0k N\u03b8 \u03b8 \u03bb \u03b1 and 0\u03b1 is initial angle of the offset vector. Finally, according to the kinetics of the machine tool, the cutting trace can be translated to the coordinate of the workpiece. Specifically, for a 5-axis machine tool with a tilt head and rotary table (Fig. 3), the cutting trace of a point on the cutting edge of cutter is written as w r r j w j( , , ) ( )[ ( )( ( ) ) ],k t t\u03b8 \u03b3 \u03b1 \u03c9 r B B B r O O O O (6) where r j j r0 0 ,O O O O and \u03b1 is the tilt angle of B-axis, \u03b3 is the rotational angle of C-axis, \u03c9 is the angular velocity of the cutter. For calculating differential cutting force, the cutter is divided into a series of chips, i.e., cutting elements along the cutter axis. At the same time, in order to compute the instantaneous uncut chip thickness, a line segment which is from cutter edge element and perpendicular to the cutter axis is prepared to intersect the cutter edge path surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003992_arso46408.2019.8948797-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003992_arso46408.2019.8948797-Figure4-1.png", + "caption": "FIGURE 4. Hysteretic creep excitation function.", + "texts": [ + "3, the chaotic characteristic of the neuron has a good robustness to the change of the hysteretic parameters, that is, neurons can be always in the chaotic state when the hysteretic parameters are changed in a large range. Furthermore, when there is the creep characteristic in the hysteretic parameter, the activation function can be described as: fcr (s,1r) = { (1+ exp[\u2212c(s\u2212 a+1r)])\u22121 , 1s(k) > 0 (1+ exp[\u2212c(s+ a+1r)])\u22121 , 1s(k) < 0 (6) 8618 VOLUME 4, 2016 Where 1r is the creep parameter, 1r \u2208 [\u2212r,+r], r > 0 is the creep amplitude. Some uncertainty can be caused in the hysteretic response of neuron by the creep characteristic. For instance, in Fig.4, the uncertain response fcr betweens fr\u2212 and fr+ will be output for the input s, that is fcr \u2208 [fr\u2212, fr+]. Thus, the uncertain response of the neuron D is caused by the creep characteristic. D = fcr (s,1r)\u2212 fc(s) (7) According to the Fig.4, Q is the upper boundary of the uncertainty D, that is |D| \u2264 Q, Q is the maximum amplitude of the output response of the neuron due to the creep. Various of neural network models can be obtained by the different coupling way of the neurons above. For instance, the mathematical model of the deterministic chaotic neural network can be described as follows: xi(k) = f [yi(k)] (8) yi(k + 1) = pyi(k)+ \u03b2 n\u2211 j=1,j6=i wijxj(k)+ Ii \u2212\u03b1 [xi(k)\u2212 I0] (9) Where the parameter \u03b2 is the coupling coefficient among the neurons, and wij is the weight between neuron i and neuron j" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003914_icems.2019.8922162-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003914_icems.2019.8922162-Figure1-1.png", + "caption": "Fig. 1. Schematic diagrams of the static eccentricity forms When considering static eccentricity, the air gap permeability can be computed.", + "texts": [ + " TABLE I FREQUENCY ORDER (k=0,1,2\u2026) \u03c9p2 kZp \u00b1\u03c52 16 \u00b1= k\u03bd \u03c9p2 kZp \u00b1+ )( 21 \u03c5\u03c5 1616,21 \u2212=+=\u2260 kk \u03bd\u03bd\u03bd\u03bd \u03c9p2 kZp \u00b1\u2212 )( 21 \u03c5\u03c5 16,16 21 kk =\u00b1= \u03bd\u03bd ( ) \u03c9\u03bc p\u00b11 kZp \u00b1\u00b1 )( \u03c5\u03bc 16,12 +=+= kk \u03bd\u03bc ( ) \u03c9\u03bc p\u00b11 kZp \u00b1)( \u03c5\u03bc 16,12 \u2212=+= kk \u03bd\u03bc ( ) \u03c9\u03bc\u03bc p21 \u00b1 kZp \u00b1\u00b1 )( 21 \u03bc\u03bc 12 += k\u03bc The length of air gap is function of the relative position between the stator and rotor as shown ( ) \u03b3\u03b8\u03b5\u03b4\u03b4\u03b8\u03b4 cos\u2212= (8) Where \u03b4 is the nominal length of air gap, namely the air gap length when the rotor is not eccentric, \u03b5 is the eccentricity ratio, \u03b8 is the relative position angle between the stator and rotor, 0=\u03b8 is set at the minimum air gap; \u03b3 is different eccentricity forms, and 1=\u03b3 represents the noncoaxial eccentricity, and 2=\u03b3 represents the elliptical eccentricity. The schematic diagrams of the static eccentricity forms discussed in this paper are demonstrated in Fig. 1, including the non-coaxial case and elliptical case. \u221e = \u221e = \u039b\u2212\u039b= \u2212\u039b=\u039b 0 0 0 )cos(cos),( )cos1()cos(),( k k k k kZt kZt \u03b8\u03b3\u03b8\u03b5\u03b8 \u03b3\u03b8\u03b5\u03b8\u03b8 (9) Where \u221e = \u039b 0 )cos(cos k k kZ\u03b8\u03b3\u03b8\u03b5 is defined as an eccentric magnetic guide portion. Since the air gap magnetic density can be expressed as the product of the magnetic potential and the air gap permeability, the radial magnetic potential expression of the motor can be derived from the stator, rotor magnetic potentials and air gap permeability. At the same time, it can be seen that the static eccentricity increases a series of force waves generated by the interaction between the stator, rotor magnetic potentials and the eccentric magnetic flux" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002870_joe.2018.8345-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002870_joe.2018.8345-Figure5-1.png", + "caption": "Fig. 5 Magnetic field distribution of the motor C", + "texts": [ + " To analyse the distribution of the electromagnetic force on the surface of the permanent magnet, the permanent magnet was divided into small pieces, so that the magnetic field parameters could be correctly calculated. The force of each small piece of permanent magnet was calculated separately, and then the force of each small permanent magnet was applied. Coupled into the structure field, the calculation accuracy of the vibration displacement could be significantly improved. Each permanent magnet was divided into 16 pieces here. The simulated magnetic field distribution in the motor is shown in Fig. 5. The radial magnetic density distribution on the centreline of the air gap was obtained by finite element calculation as shown in Fig. 6. When the rotor rotated, the magnetic density of the centre point of the permanent magnet surface periodically fluctuated, and the number of fluctuations was equal to the number of rotor slots, as shown in Fig. 7. Fig. 8 shows the distribution of the local electromagnetic force on the surface of the permanent magnet. Fig. 9 is the time FFT decomposition of magnetic density on the surface of a permanent magnet" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000825_auv.2012.6380732-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000825_auv.2012.6380732-Figure2-1.png", + "caption": "Fig. 2. Diagram illustrating the location of the front and rear vertical tunnel thrusters and the axis system.", + "texts": [ + " From equation (6) it can be seen that the state-variable vector contains the system outputs y and the change of the original state-variable vector (of the original model) \u2206xm. The statevariable vector used for feedback has the form: x(ki) = [ xm(k)\u2212 xm(k \u2212 1) y(k) ] (24) In this work the effect of measurement noise on a depth and pitch controller, built using the MPC algorithm, is evaluated. This system is an approximation of the low-speed hovering operational mode of the Delphin2 AUV; with the system inputs as the front and rear vertical tunnel thrusters and the system outputs of depth (z) and pitch (\u03b8), see Fig. 2. In order to effectively test this controller in simulation, a non-linear model is used as the system. This non-linear system can be modelled as two coupled subsystems; depth and pitch: 1) Depth Model: w\u0307v = 1 mz [Tvf cos \u03b8+Tvr cos \u03b8+(W\u2212B)\u22121 2 \u03c1V 2/3CDw|wv|wv] (25) wv = \u222b t 0 w\u0307v dt, z = \u222b t 0 wv dt (26) 2) Pitch Model: q\u0307v = \u2212 1 Iy [xTvfTvf+xTvrTvr\u2212zgW sin \u03b8\u22121 2 \u03c1V 2/3CDq|qv|qv] (27) qv = \u222b t 0 q\u0307v dt, \u03b8 = \u222b t 0 qv dt (28) Both the pitch and depth models, equations (27) and (25), have non-linear trigonometric and quadratic terms that need to be linearised for use with the MPC algorithm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002033_elma.2017.7955470-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002033_elma.2017.7955470-Figure2-1.png", + "caption": "Fig. 2. Dimensions of the geometry of the investigated coaxial magnetic gear", + "texts": [ + " MAGNETIC GEAR\u2019S CONSTRUCTIONS In this section two constructions of a coaxial magnetic gear have been considered. The dimensions of the gear for both constructions are the same. The general arrangement drawing of a coaxial magnetic gear is depicted in Fig. 1. The first construction has 4 inner permanent magnet pole pairs, 22 outer permanent magnet pole pairs and 26 steel segments. The second one has 6 inner permanent magnet pole pairs, 18 outer permanent magnet pole pairs and 24 steel segments. The sketch with the dimensions of the magnetic gear\u2019s construction is depicted in Fig. 2 and the dimensions are shown in Table I. The permanent magnets, which are used in the permanent magnet rotors, are made of NdFeB35 alloy [13]. The material used for the steel segments is a low carbon steel AISI 1008 [14]. The stack length of both constructions of the magnetic gear is 300 mm. More detailed information for the first construction of the coaxial magnetic gear can be found in [15]- [17]. III. PRINCIPLE OF OPERATION OF THE MAGNETIC GEAR The principle of operation of the magnetic gear is based on the modulation of the inner and outer rotors\u2019 magnetic fields of the gear through the steel segments" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001462_iros.2011.6094821-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001462_iros.2011.6094821-Figure1-1.png", + "caption": "Fig. 1. Four-rotor rotorcraft with crosswind", + "texts": [ + " If an aircraft is experiencing a crosswind, it will be pushed over or yawed away from the wind. In UAVs with propellers, the airflow generated by the lateral wind affects the airflow generated by the propellers introducing additional forces, fwk , on the aircraft 1. Thus, the magnitude of these forces depends on the induced wind speed in the propeller, Vp, and the incoming lateral airflow coming from the wind, Vw, i.e., fwk = (Vp,Vw), see [13], [12]. The quad-rotor rotorcraft is an underactuated mechanical system since it has four inputs and six degrees of freedom (see Figure 1). It has four motors which give special characteristics to it. The front and the rear motors rotate counterclockwise, while the other two motors rotate clockwise, whereby the gyroscopic effects and aerodynamic torques tend to cancel in hovering. The only forces acting in the corps (and applied in the z-direction in the body frame) are given by the translational forces fMi produced by motor i, the gravitational force g and the aerodynamic forces induced by the wind fwi in each rotor i. On the other hand, the gravitational force applied to the vehicle is fg = \u2212mgk\u0302, where k\u0302 denotes the unit vector codirectional with the z-axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003748_10402004.2019.1669755-Figure10-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003748_10402004.2019.1669755-Figure10-1.png", + "caption": "Figure 10. 2MNK9810 ultraprecision grinding machine.", + "texts": [ + "20E\u00fe 06), was 70mm in the positive direction of Z and 7mm in the negative direction of X, as shown in Fig. 8f. All of the simulated deformation data for node 24,183 at different pump pressures are shown in Fig. 9. The amount of deformation at this location will increase almost linearly with increasing pump pressure. 2MNK9810 is a Computerized Numerical Control (CNC) nanogrinding machine tool for grinding of high-performance parts, such as the flexible hinges of gyroscopes and inertia navigation components. Its configuration is shown in Fig. 10. This machine tool is composed of a column, a Z slide, XY slides, a machine bed, and a vibration control system. The Z slide adopts a closed hydrostatic guideway structure, including the column, guideways, hydrostatic block system, linear motor, slide table, motorized spindle and its fixture, linear encoder, and high-pressure pneumatic counterweight system, as shown in Fig. 11. The geometry and loads of the vertical slide are all symmetric about the Z axis and therefore only half of the vertical slide was simulated in order to reduce the computational intensity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003315_0954406219841076-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003315_0954406219841076-Figure9-1.png", + "caption": "Figure 9. 2 DoF helicopter workstation.", + "texts": [], + "surrounding_texts": [ + "Experimental results are provided to demonstrate the effectiveness of the proposed sliding mode controller based on GA. The Quanser system given in Figure 937 consists of a helicopter module mounted on a fixed base with two propellers driven by two DC motors with 24V and 15V to control the pitch and yaw angles respectively, two power amplifiers for driving the pitch and yaw motors. The pitch and yaw angles are measured using high-resolution encoders with a resolution of 4096 counts/revolution for pitch and 8192 counts/revolution for yaw angle. These measured signals are transferred through a slip ring. The control algorithm implemented in MATLAB 2015b communicates with the hardware using QUARC, which is similar to C like programming language. Without loss of generality, the system has two degrees of freedom: a motion around the yaw axis represented by the angle w and the rotation around the pitch axis denoted by . The input voltages to the DC motors present the control variables and the objective is to control the pitch and yaw angles so as to make the system to track the reference trajectory. Based on the Euler\u2013Lagrange formula, the equation governing the dynamics of the system is given by38 \u00f0Jeq,p \u00feml2\u00de \u20ac \u00bc KppVmp \u00fe KpyVmy Bp _ \u00fe \u00f0t\u00de \u00f0Jeq,y \u00feml2cos2\u00f0 \u00de\u00de \u20ac \u00bc KypVmp \u00fe KyyVmy By _ \u00fe@\u00f0t\u00de 8>< >: \u00f062\u00de where \u00f0t\u00de \u00bc ml2sin\u00f0 \u00decos\u00f0 \u00de _ 2 mglcos\u00f0 \u00de @\u00f0t\u00de \u00bc 2ml2 _ sin\u00f0 \u00decos\u00f0 \u00de _ ( \u00f063\u00de \u00f0t\u00de, _ \u00f0t\u00de, \u00f0t\u00de, and _ \u00f0t\u00de are respectively the pitch angle, the pitch velocity, the yaw angle, and the yaw velocity. Kpp,Kpy,Kyp, and Kyy are the thrust force constants. Vmp and Vmy are respectively the control input voltages to pitch and yaw motors. The plant parameters of the helicopter system are given in Table 5. To assess the performance of the proposed controller framework, the tracking control of the 2-DoF Quanser Aero is compared with that of the conventional SMC. Table 6 shows the parameters of both optimal SMC based on GA and classical SMC algorithms. Figures 10 and 11 illustrate the pitch and yaw responses, using each one of the controller. They highlight that the trajectory tracking of the Quanser Aero is ensured and each angle, for both cases settles quickly to the desired signal in spite of the presence of disturbances. However, as can be deduced from these figures, GA-based SMC provides better performances than the classical one in terms of convergence and precision. As demonstrated in the indicated zoomed figure in Figure 10, a highest value of error is obtained in case of the conventional SMC, in which the two matrices Q and R are chosen arbitrary. An improvement of the load disturbances rejection is ensured in case of the proposed design method. Thereafter, we conclude that the combination of SMC with both LQR method and GA seems to be a good choice for the control of the Quanser Aero. It should be noted that the maximum amplitude of the pitch and the yaw control inputs considered to obtain experimental results are respectively Vmp,max \u00bc 24V and Vmy,max \u00bc 15V. Figures 12 and 13 illustrate the evolution of the two components of the control input during the experiment. Firstly, it can be seen that the pitch and yaw motor voltages do not reach the saturation value while generating a required control signal to accelerate the pitch and yaw propellers. Secondly, as compared to the classical SMC, an improvement of the stability is ensured in case of combination of SMC with GA. The total variation (TV) minimization approach has been proved experimentally. Based on the different zoomed time history of Figures 12 and 13, it is clear that our proposed method can eliminate the chattering phenomenon. Detailed comparison of these controllers as regards to the transient performance is summarized in Table 7. The different values of time convergence, steady state errors and the total variation of control are presented. According to this table, all variables corresponding to the proposed algorithm were observed to converge faster and they have small amplitude oscillations than those of the classical SMC. Furthermore, the steady-state error for whole variables obtained when both methods are applied shows clearly that the suggested algorithm is more precise and offers faster convergence and better robustness than the classical one. Furthermore, it can be seen that the optimization of two matrices Q and R engender the minimization of the total variation of control and thereafter, the elimination of the chattering phenomena. Thus, this table definitely asserts not merely the efficiency as well as the quality of the designed controller but also its supremacy compared with the control algorithm developed without optimization." + ] + }, + { + "image_filename": "designv11_33_0000336_6.2010-2790-Figure17-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000336_6.2010-2790-Figure17-1.png", + "caption": "Figure 17. AFIT flapping mechanism. The 4-bar flapping mechanism is driven by a 17V DC hobby car motor (A) connected to an adjustable power supply for controlling wingbeat frequency. It is capable of operating at frequencies as high as 40 Hz, but was limited to 25 Hz for this research. With exception of the motor and metal crank (B) that connects the flapper arm (C) with the drive wheel (D) all components are made of rapid prototype plastic and printed on an Objet Eden350 3-D printer. Small coded markers are affixed to the mechanism and surroundings to provide photogrammetric references. The L-bracket (E) is utilized for the sole purpose of adding additional markers to the scene. An infrared emitter/sensor (not shown) triggers a set of strobe lights each time a small piece of reflective tape (F), attached to the drive disk, passes by the emitter. Note the presence of the green foam within the flapper arm illustrated in the inlay (F). The final variant of the wing clamp employing a \u201csandwiching\u201d method was \u201cbuilt-in\u201d to the mechanism\u2019s flapper arm.", + "texts": [ + " Because we did not actively control vacuum pressure during test, and since all chambers have a finite leak rate, the pressure at the end of each test varied between 2-3 Torr. hawkmoth. 16 American Institute of Aeronautics and Astronautics Our ability to test in air and vacuum using a common testing apparatus made the opportunity of accomplishing a full scale aeroelastic test too tantalizing to resist. Our goal in testing was to corroborate the findings of Combes & Daniel6 (C&D) by using an improved methodology. The experimental apparatus consisted of a vacuum chamber (Figure 15), flapping mechanism (Figure 17), and a simple stroboscopic camera system (Figure 16). The vacuum chamber is simple enough and described in detail in the caption below Figure 15, but the design of the mechanism and the camera system merit further discussion. In order to flap the wing we obviously need a mechanism since half of the testing takes place in vacuum and a hawkmoth, or any other animal for that matter, could not perform nor survive in such an environment. Our mechanism is effectively a 4-bar linkage designed in SolidWorks and fabricated with an Objet 3-D printer" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000020_1.3610022-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000020_1.3610022-Figure5-1.png", + "caption": "Fig. 5 Equilibrium forces in two-degree-of-freedom chain", + "texts": [ + " The auxiliary loops can then be added, as required or as desired, to give direct readings of displacements at points of interest. Equilibrium Analysis Equilibrium Configuration Under App l ied External Forces. T h e p r o b - lem to be considered now is the determination of an appropriate equilibrium configuration for a single or multiple-loop, single or multiple-degree-of-freedom kinematic chain, which is in equilibrium under the actions of internal spring forces and external conservative forces. In the 2-degree-of-freedom problem of Fig. 5 the expressions shown for the equilibrium forces Pk and P\u201e as functions of chosen generalized coordinates x and y, are easily derived. Inverting the expressions for Ph and P, to give explicit relations for x and y as functions of Ph and P\u201e is more difficult, and may not be feasible even for this simple problem. Given particular data, an implicit solution for x and y could be carried out by assuming a range of x and y-values and evaluating corresponding Pk and P\u201e-forces, tabulating these forces, and then scanning these results to establish approximate equilibrium x and " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001673_sav-2010-0501-Figure10-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001673_sav-2010-0501-Figure10-1.png", + "caption": "Fig. 10. Third optimization step.", + "texts": [], + "surrounding_texts": [ + "In this study, an analytical and numerical model is worked out to simulate the possible dynamic behaviour of a suspension system as well as the state of the stress and the strain energy density in a lower suspension arm. The multiaxial criterion of the uniaxial case of strain energy density equivalent to the multiaxial one is used to calculate the part fatigue life. A model of the negative feedback and forward path of the vehicle\u2019s lower suspension arm, which is subjected to very great dynamic stress, is developed. Using this model, the maximum force of impact between the road and the tire of a vehicle measured in a 0.5 second fraction is estimated at 10 KN. However, the maximum value of the force transmitted by the tire is about 4 KN. Numerical calculations show that the maximum negative feedback force is 1200N in the damping case and 3750N in the rigid case. To filter the critical element and to extract the fatigue life, a Matlab interface is generated to locate automatically the critical elements without applying the Newton-Raphson algorithm in every element of the mesh. This filter takes into account the multipoint excitation shifted in time. The shifted excitation gives a tangle of the stress signals of the material mesh elements. The elasto-plastic nonlinear case is described by applying the Ramberg-Osgood uniaxial relation binding stress to deformation. Following the first iteration, all elements with a fatigue life less than 10% of the critical element are removed; between 5 and 11% of the weight is lost. Unlike previous methods, the rejection ratio developed in this work is directly linked to the notion of fatigue. This method allows the calculation of the optimum weight for the part in relation to its fatigue life and its eigen mode in a reasonable time period using either a coarse or a refined mesh." + ] + }, + { + "image_filename": "designv11_33_0003238_s40684-019-00095-4-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003238_s40684-019-00095-4-Figure9-1.png", + "caption": "Fig. 9 a Experiment setup of a single-link manipulator. b Diagram of information interchange. The dotted black line describes the reading data from MP2300\u00a0s after the experiment through ETHERNET, the blue solid line transfer angle, angular velocity and torque through MECHATROLINK, the red solid line denotes calculating angle from the motor encoder, and the green line shows actuating the motor by calculated torque", + "texts": [ + " Thirdly, energy consumption of the NNC is significantly higher than the ESC. In our perspective, neural network weights keep going to different local minimums because of randomly choosing the initial weights in the NNC, which leads to fluctuation in control torque and further results in loss of accuracy and overuse of energy. However, in the proposed controller, the initial weights are given the estimation values that may have fewer local minimums surrounding with. Experimental testbed is given in Fig.\u00a09a. The single-link manipulator is equipped with a Yaskawa@ \u03a3\u00a0\u2212\u00a07 motor. The product number of the motor and the servopack is SGM7J01A and SGD7SR90A. We can utilize ladder diagram programing system to assign desired torque by controller MP2300\u00a0s. Information exchange is shown in Fig.\u00a09b. Leading signal is chosen as q0 = \u03c0/4\u00a0\u2212\u00a0\u03c0/4\u00a0cos\u00a0(2\u03c0t/5)rad that keeps same with Example 1. Control parameters are chosen as K1 = 10, K2 = 0.0318, K3 = 9.54 \u00d7 10\u22124, K4 = 5, k5 = 100, k6 varies from 200 to 1000, k7 = k8 = k9 = 0, \u03b8M = VM = WM = 109, \ud835\udf19 = ||e|| + 0.1||e\u0307|| , \u03d5m = 0.02, \u03d5M = 0.1 and \u03bcM = 30. The control frequency is 1000\u00a0Hz. Sub neural networks are assigned values trained in Example 1. Similar with literature [9], the control torque is additionally compensated by pre-identified Coulomb friction model and viscous friction model in order to reduce the influence of friction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003494_012194-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003494_012194-Figure2-1.png", + "caption": "Figure 2. Kinematic diagram of the strip-block brake: 1 \u2013 brake pulley, 2 \u2013 brake band, 3 \u2013 improved linings.", + "texts": [ + " Mass transfer determines the energy levels of friction surfaces in the interaction of their microprotrusions (dependences (11) and (12)). The fourth case differs from the third one in the values of the gradients (grad\u03c6\u044d)\u0410f1 and (grad\u03c6\u044d)\u0410f2 since the energy levels of the semiconductors in the friction pair are different. their pulsed frictional interaction destroys the charges of the transverse electric field. Semiconductor elements operating in the transistor mode are installed in the body of the polymer linings of the friction brake nodes. According to the kinematic schemes of the strip-block brake (Figure 2), the improved linings 3 are mounted on brake bands 2, which are attached to the balancer at one end (on the side of the leaving branch of the band), and to the crank necks of the crankshaft at the other end (on the side of the entering branch). The figure 2, the following conventions are used: Se, Sl \u2013 tension of the entering and leaving branches of the band; Fw \u2013 worker's effort; Rp, r \u2013 radii: the working surface of the rim of the pulley; crankshaft; \u03c9 is the angular velocity of the brake pulley. MEACS2018 IOP Conf. Series: Materials Science and Engineering 560 (2019) 012194 IOP Publishing doi:10.1088/1757-899X/560/1/012194 Serial strip-block brakes drawworks work as follows. By moving the crank, the crankshaft is rotated, as a result of which the driller tightens the brake bands 2 with advanced friction linings, 3 and they sit on the brake pulleys 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002729_s11837-018-2982-1-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002729_s11837-018-2982-1-Figure1-1.png", + "caption": "Fig. 1. EIKOS specimen-electrode alignment. (a) The specimen carrier with specimen in place aligned to the 1-mm electrode and (b) the three-axis optical microscope alignment station.", + "texts": [ + " The specimen/electrode stage is fixed and directly attached to a cryogenic cooler. The use of a 532-nm laser leads to performance comparable to the LEAP 30009 HRTM for many materials systems (relative to a UV laser system) and simplifies design and cost of ownership. The use of a curved reflectron19,20 in the fundamental flight path design enables high spectral quality in both voltagepulsed and laser-pulsed modes. This system (called the EIKOSTM21) uses a wiregeometry sample in a holder as shown in Fig. 1a. An integrated counter electrode with a 1-mm aperture is affixed to a carrier that is aligned over the top of the specimen (prior to insertion into the vacuum chamber) using a three-axis, computeraided alignment center (Fig. 1b). The specimen is first crimped into a carrier called a stub. This can be accomplished with simple tweezers and pliers, but a station to assist with uniform crimping and total length control is included. The specimen in the stub is then loaded into the holder, called a puck, that is similar in dimension to a LEAP puck, but it is not physically compatible. An XY stage sets the specimen apex in the center of the counter electrode aperture, and a pin below the stub is used to set the height such that the specimen protrudes through the aperture", + "3 to 4 results in a small decrease ( 18%) in maximum analyzable specimen radius. For a convenient electrode diameter of 1000 lm (dashed line in Fig. 2a), the optimum protrusion (minimum k) is within the minimal variation in the range of 100\u2013300 lm (Fig. 2b). Positioning of the specimen to take advantage of Larson, Ulfig, Lenz, Bunton, Shepard, Payne, Rice, Chen, Prosa, Rauls, Kelly, Sridharan, and Babu this optimal geometry with placement errors resulting in less than 2.5% relative error in the predicted k value is easy to achieve using the apparatus shown in Fig. 1b. Unlike in the LEAP where the specimen must be carefully positioned in the center of the local electrode aperture to avoid clipping data from the detector, specimen positioning away from the aperture center was found have no measurable impact on data quality. Experimental field enhancement of the EIKOS geometry may be compared to that of traditional remote electrode, non-LEAP, atom probes through a simple experiment. An aluminum specimen was run with and without an integrated electrode. When the integrated electrode is removed, the detector serves as the counter electrode and should provide field enhancement similar to that of a very large aperture or very distant counter electrode" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003437_ceit.2018.8751810-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003437_ceit.2018.8751810-Figure1-1.png", + "caption": "Fig. 1. Overall structure of AR.Drone", + "texts": [ + " A carbon-fiber frame has been used to keep the parts together. Adjacent propellers rotate in the opposite directions to prevent spinning around the central-axis. Depending on the type of motion desired, different levels of energy should be applied to the motors driving the propellers. Drones have relatively simple structure, yet they are capable of performing a wide variety of movement patterns and exhibiting high maneuverability. They have six degrees of freedom, which consist of three translational and three rotational components. As seen in Figure 1, pitch is corresponding to a rotational movement along y-axis and generates a translational 978-1-5386-7641-7/18/$31.00 c\u00a92018 IEEE movement along the x-axis. Similarly, roll is corresponding to a rotational movement along x-axis, which results in a translational movement along the y-axis. Initially, AR.Drone was designed as a hobbyist\u2019s toy which can be controlled by Android and iOS via Wi-Fi connection. However, embedded two circuit boards enable the use of AR.Drone for research related applications [10], [11]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003397_s00170-019-03894-w-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003397_s00170-019-03894-w-Figure5-1.png", + "caption": "Fig. 5 Types of tool with respect to tooth profile error", + "texts": [ + " Manufacturing of gears on universal machine tools is located in the area of NC form milling [1], but the tooth surface of gears is different from the free-form surface. The influence of tooth profile error related to the processing parameters such as the type of tool, the feeding strategy of tool, and the basic parameters of gear will be discussed in the following sections. In the process of free-form milling of gears, many types of cutting tool, such as end mill, ball-end mill, and taper ball-end mill, are available. Different types of cutting tools result in different processing efficiencies as well as different surface structures of tooth flank. Figure 5 gives an overview of the influence of tool types with respect to tooth profile error. Note that the side edge of end mill is positioned tangential to the tooth surface while the top arc edge of ball-end mill is tangent to the tooth surface. Depending on the selection of tool, different types of milling tool result in different machining times. As shown in Fig. 6, the relation between cutting pass number and residual profile errors is presented. In this example, the residual profile errors are varied from\u0394t = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003359_pedes.2018.8707719-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003359_pedes.2018.8707719-Figure4-1.png", + "caption": "Fig. 4. Cross section of the analyzed PMSM with buried magnets.", + "texts": [ + " With the resulting fault network, we can simulate for example inter-turn faults, phaseto-phase faults or phase-to-ground faults. In contrast to the FEA simulation, this implementation allows a simulation with variable angular velocity and time-variant fault components. Furthermore, the computation time of the analytical model is about 200 times faster. We use a 2D FEA in this paper to compare the simulation results of the analytical model. The FEA simulation is performed with the software Flux-2D from Altair and considers the nonlinear magnetic circuit. Fig. 4 shows the cross section of the analyzed PMSM with buried magnets. The electric circuit of the FEA model equals the electric circuit of the analytical model. It is essential to take the complete cross section of the PMSM for the FEA model into account, without utilizing any periodicity. Otherwise, the stator winding faults would also be repeated periodically. The FEA model is also used to determine the flux linkage \u03a8PM and the inductance Ld and Lq, by linearization at the nominal operating point" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003891_s00498-019-00251-w-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003891_s00498-019-00251-w-Figure4-1.png", + "caption": "Fig. 4 Drawing of when both legs i and i+1 are on the ground", + "texts": [ + " , n} and E = {(i, i + 1) : i = 1, . . . , n} where n is the number of spokes and it is understood that all indices are taken modulo n. The remaining data, (X , S,\u0394, f ), depend on the geometry of the wheel. Consider the rimless wheel with different spoke lengths and different inter-spoke angles. For i = 1, . . . , n, let \u03b4i and i be half the angle between spokes i and i + 1 and the spoke length of spoke i , respectively. We will assume that the grade of the slope being walked down has a constant angle of \u03b1, see, Fig. 4. In what follows, x1 will be the angle of the wheel from vertical (with respect to the spoke in contact with the ground) and x2 = x\u03071 is the angular velocity. Each vector field, fi , is that of an inverted pendulum of length i , x\u0307 = fi (x) = [ x2 \u03b6i sin(x1) ] , (13) where \u03b6i = g/ i and g is the acceleration due to gravity. The impact surface, S(i,i+1), is given by configurations where both legs i and i + 1 are in contact with the ground. That is, S(i,i+1) = {(x1, x2) : x1 = \u03c0/2 \u2212 \u03bei \u2212 2\u03b4i \u2212 \u03b1} " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002131_metroaerospace.2017.7999605-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002131_metroaerospace.2017.7999605-Figure2-1.png", + "caption": "Fig. 2. Example of measurement system based on electronic balance. [8].", + "texts": [ + " On the other shaft extreme, an electronic balance is attached. By means of the shaft, the thrust force exerted by propeller, due to the motor rotation, is transmitted on the electronic balance. By using this test bench, it is possible to measure the thrust force exerted by propeller for each motor rotation speed and to determine the relationship between them. This relationship is used for optimizing the mechanical and the electrical RPAS design and the control method. An example of test bench, implementing this system, is reported in Fig. 2. Another measurement system for analyzing the thrust force is depicted in Fig. 3 [6]. In this case, the drive shaft of the BLDC motor is connected to the drive shaft of a load motor. Motor-propeller configurations with different load conditions can be emulated and the corresponding load torque values are measured. For taking into account as the atmosphere conditions, such as wind direction and speed, can affect the thrust force, another test bench is based on the electronic weight balance method placed in a wind tunnel [6], [7]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001852_2016024-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001852_2016024-Figure1-1.png", + "caption": "Fig. 1. Piston ring pack geometry of a four stroke motorbike engine.", + "texts": [ + " The asperity 18-page 3 interactions at the dead centres were taken into account by means of the Greenwood-Tripp model. To study the lubricating ability of the compression ring close to the actual conditions, variation of the lubricant oil density and viscosity is considered in this study. In this paper a 2D axisymmetric geometry is assumed and the basic dimensions of the top compression ring and the cylinder inner liner were measured by a 3D coordinate measuring machine having an accuracy of 1 \u03bcm per axis. In Figure 1, a schematic view of the piston ring pack and the top compression ring shape are illustrated. The top ring axial profile was obtained by an ideal parabola in the present analysis. However, compression rings profiles that are not symmetrical caused significant changes in the contact, hence it is necessary to analyse the real profiles for the future investigation in this study. The first compression ring is a chromium plated steel ring characterized by radial width W = 3.35 mm and a thickness of B = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002186_s0965542517080073-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002186_s0965542517080073-Figure3-1.png", + "caption": "Fig. 3.", + "texts": [ + " If the distance obtained for a given is greater than the prescribed one, is decreased, and the computation is repeated. The decrease of is repeated until becomes acceptable. 3. APPLICATION TO THE PROBLEM OF DETERMINING THE WORKSPACE OF A ROBOT We use the method described above for determining the workspace of a parallel robot (see [10]). The workspace of a robot is the set of all positions of the tool controlled by the robot. We consider this problem using a planar robot with two links of variable length [11] (Fig. 3). We want to approximate the set that can include the point to a certain accuracy. The two bars of the robot are fixed at the points 1 and 2 using mechanisms that make it possible to vary the lengths and ensure the unobstructed rotation of the bars about the points where they a fixed. The workspace of this robot is determined by a system of inequalities of form (1) with the following functions : (9) \u03b4 \u03b4 X , \u2264 ,( ) ( )I I Eh Q X h Q Q , \u2264 ,( ) ( )E I Eh Q X h Q Q IQ EQ n ,( )I Eh Q Q \u03b4 \u03b4 \u03b4 ,( )I Eh Q Q \u2286 2X R z ,1 2l l ( )jg x , = + \u2212 , , = \u2212 \u2212 ,2 2 max 2 min 2 2 2 1 1 2 1 2 1 2 1 2 1 1 2( ) ( ) ( ) ( )g x x x x l g x x l x x , = \u2212 + \u2212 , , = \u2212 \u2212 \u2212 ,2 2 max 2 min 2 2 2 3 1 2 1 0 2 2 4 1 2 2 1 0 2( ) ( ) ( ) ( ) ( ) ( )g x x x l x l g x x l x l x \u03b4 = " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000945_ssd.2010.5585533-Figure10-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000945_ssd.2010.5585533-Figure10-1.png", + "caption": "Figure 10. Mesh of the stator lamination.", + "texts": [ + " Moreover it is to be noted that the two claw plates are magnetically decoupled. In order to reduce the computation time, the FEA study domain is limited to a one pair of poles of the CPAES. Figures 7 and 8 show the stator and the rotor study domains, respectively. The meshing of both stator and rotor study domains has been automatically generated in the Ansys-Workbench environment. Figure 9 shows a mesh of the stator yoke and of the two associated magnetic collectors. A mesh of the stator lamination is illustrated in figure 10. Figure 11 shows a mesh of the rotor claws and the as sociated magnetic rings. 3.2.1. Main Flux Paths The flux moves through the magnetic circuit of the CPAES within three dimensional (3D) paths. These are described as follows: \u2022 axially in the stator yoke (see figure 12), \u2022 radially down through the collector and crossing the air gap down to the magnetic ring (see figure 12), \u2022 radially then axially in the magnetic ring (see figure 12), \u2022 axially-radially in the claws (see figure 13), \u2022 radially up in air gap facing the stator laminations (see figure 14), \u2022 radially up in the stator teeth and circumferentially in the stator core back (see figure 15), \u2022 radially down in stator teeth facing the two adjacent claws (see figure 16), \u2022 radially-axially in the adjacent claws (see figure 16), \u2022 axially then radially in the magnetic ring (see figure 17), \u2022 radially up through the air gap and the mag netic collector on the other side of the machine (see figure 17), \u2022 axially in the other side of the stator yoke (see figure 18)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000949_detc2011-47346-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000949_detc2011-47346-Figure1-1.png", + "caption": "FIGURE 1. A schematic of two-dimensional lumped parameter model of a planetary gear system", + "texts": [ + " This discrepancy between static and dynamic response emphasizes the importance of dynamic analysis when seeking an optimum TPM to reduce planetary gear vibration. Contrary to the expectation of further vibration reduction, increased dynamic response is observed when the both optimum sun-planet and ring-planet TPMs giving minimum response when applied individually are combined. A two-dimensional lumped-parameter model of a planetary gear developed by Lin and Parker [13] and Ambarisha and Parker [14] is adopted as a basic model without tooth profile modification (TPM). TPM will be included in the basic model and used to study dynamic response with TPM. Figure 1, from [14], shows a schematic of the planetary gear model. The matrix equation of motion with N planets is Mx\u0308+K(x, t)x = F x = [xc,yc,uc,xr,yr,ur,xs,ys,us,\u03b61,\u03b71,u1, \u00b7 \u00b7 \u00b7 ,\u03b6N ,\u03b7N ,uN ] T , (1) where M is the inertia matrix and K = Km(x, t)+Kb is the stiffness matrix including the nonlinear, time-varying mesh stiffness Km(x, t) and the bearing stiffness Kb (refer to [13] for detailed system matrices). ksn(t) and krn(t) in [13] are replaced with ksn(x, t) krn(x, t), which are ksn(x, t) = ksn(t)\u0398(\u03b4sn) = { ksn(t) \u03b4sn \u2265 0, 0 \u03b4sn < 0, n = 1,2, \u00b7 \u00b7 \u00b7 ,N krn(x, t) = krn(t)\u0398(\u03b4rn) = { krn(t) \u03b4rn \u2265 0, 0 \u03b4rn < 0, \u03b4sn = ys cos\u03c8sn \u2212 xs sin\u03c8sn \u2212 \u03b6n sin\u03b1s \u2212\u03b7n cos\u03b1s +us+un \u03b4rn = yr cos\u03c8sn \u2212 xr sin\u03c8rn + \u03b6n sin\u03b1r \u2212\u03b7n cos\u03b1r +ur \u2212un, (2) where \u03c8sn = \u03c8n \u2212\u03b1s and \u03c8rn = \u03c8n +\u03b1r, \u03c8n is the circumferential angle of nth planet, and \u03b1s and \u03b1r are the pressure angles of the sun-planet and ring-planet meshes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003907_pee.2019.8923240-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003907_pee.2019.8923240-Figure5-1.png", + "caption": "FIGURE 5. (a) Position trajectories of the multi-vehicle system; (b) Velocity trajectories of the multi-vehicle system.", + "texts": [ + "0367 0.3767 0.2374 0.7099 \u22120.2053 0.2833 0.7099 2.4161 . Solid blue lines in Figs. 5(a) and (b) respectively show the velocity and position trajectories of the multi-vehicle system from t = 0s to t = 40s. Using the star, asterisk, square and diamond to represent the states of each vehicle, the TVF of the multi-vehicle system under switching interaction topologies is illustrated at t = 10s using bold dash-dotted lines, at t = 25s using bold dashed lines and at t = 40s using bold dotted lines. From Fig. 5, we can observe that both the velocities and positions of the multi-vehicle system reach parallel rectangle formations and the parallel rectangles keep rotation. In Fig. 6, the switching signal, formation error, and coupling weights are displayed respectively. Fig. 6(a) shows that the interaction topology G\u03b4(t) of the multi-vehicle system switch every 1s among G1, G2, G3 and G4 randomly. From Fig. 6(b), the formation error of the multi-vehicle system converges to zero which means that the TVF is achieved" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003836_tie.2019.2950863-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003836_tie.2019.2950863-Figure4-1.png", + "caption": "Fig. 4: IM geometric model (ro = 0.2 mm, \u03d5s = 0 &\u03b8r = 0).", + "texts": [ + " To solve the magneto-static problem by FEA, the magnetic vector potential can be obtained by \u2207\u00d7 ( 1 \u00b5 \u2207\u00d7A ) = J Also, using Stoke\u2019s theorem and in two dimensional cylindrical coordinates (A = Az (\u03b8r, ro)), the mutual-inductance between the kth and jth stator phase winding (Lskj (\u03b8r, ro)) can be given by Lsk,j (ro, \u03b8r) = \u222b Az(\u03b8r, ro)dz Similarly, the mutual-inductance between the kth stator winding and the jth rotor bar (Msr k,j (\u03d5s, \u03b8r, ro)) can be obtained by the magnetic potential vector corresponding to proper points. The IM parameters are presented in Table. I. In addition, the geometric model of the IM utilized in FEA is shown in Fig. 4. Because of the non-uniform air-gap length due to static eccentricity, the magnetic flux in shorter air-gap length has a larger magnitude. The 3-D self inductance of the first rotor bar (Lr1,1 (\u03b8r, ro)) and the mutual inductance of the stator windings (Lsa,b (\u03d5s, ro)) are represented in Figs. 5a and b, respectively. By increasing the rotor eccentricity (ro), the winding inductances of the stator increase as a result of effective air-gap length reduction. Also, due to the non-uniform air-gap in the eccentricity condition, the stator winding inductances change into the quasi-sinusoidal shapes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003995_nigeriacomputconf45974.2019.8949624-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003995_nigeriacomputconf45974.2019.8949624-Figure1-1.png", + "caption": "Figure 1: Three Phase Induction Motor", + "texts": [ + " In [7] and [8], a PI controller was designed to ensure the global synchronization of some directed networks and to increase the robustness of the whole system respectively This paper provides a method of obtaining an explicit formula for calculating the proportional and integral gains of the PI controller and tests the design on a threephase induction motor model. This paper will show that the closed loop speed control of the induction motor is stabilized by the calculated controller gains. The rest of the paper is organized into sections. Section 1 gives the T introduction while section 2 describes the model of the three-phase induction motor and formulation of the controller gains. Section 3 provides the results and discussion. Section 4 concludes the paper. The schematic diagram of a line-operated three-phase induction motor is illustrated in Figure 1. Its equivalent in d \u2013 q coordinates is illustrated in Figure 2. The model of the threephase induction motor with the d-axis aligned with the rotor flux linkage is provided in equations (1) \u2013 (10) [9], [10]. [ \ud835\udc49\ud835\udc5e\ud835\udc60 \ud835\udc49\ud835\udc51\ud835\udc60 \ud835\udc490\ud835\udc60 ] = \ud835\udc3e\ud835\udc60 [ \ud835\udc49\ud835\udc4e\ud835\udc60 \ud835\udc49\ud835\udc4f\ud835\udc60 \ud835\udc49\ud835\udc50\ud835\udc60 ] (1) [ \ud835\udc56\ud835\udc5e\ud835\udc60 \ud835\udc56\ud835\udc51\ud835\udc60 \ud835\udc560\ud835\udc60 ] = \ud835\udc3e\ud835\udc60 [ \ud835\udc56\ud835\udc4e\ud835\udc60 \ud835\udc56\ud835\udc4f\ud835\udc60 \ud835\udc56\ud835\udc50\ud835\udc60 ] (2) \ud835\udc3e\ud835\udc60 = [ \ud835\udc50\ud835\udc5c\ud835\udc60 \ud835\udf03 \ud835\udc50\ud835\udc5c\ud835\udc60(\ud835\udf03 \u2212 2\ud835\udf0b 3\u2044 ) \ud835\udc50\ud835\udc5c\ud835\udc60(\ud835\udf03 + 2\ud835\udf0b 3\u2044 ) \ud835\udc60\ud835\udc56\ud835\udc5b \ud835\udf03 \ud835\udc60\ud835\udc56\ud835\udc5b(\ud835\udf03 \u2212 2\ud835\udf0b 3\u2044 ) \ud835\udc60\ud835\udc56\ud835\udc5b(\ud835\udf03 + 2\ud835\udf0b 3\u2044 ) 1 2 1 2 1 2 ] (3) [ \ud835\udc49\ud835\udc5e\ud835\udc60 \ud835\udc49\ud835\udc51\ud835\udc60 ] = \ud835\udc5f\ud835\udc60 [ \ud835\udc56\ud835\udc5e\ud835\udc60 \ud835\udc56\ud835\udc51\ud835\udc60 ] + \ud835\udc51 \ud835\udc51\ud835\udc61 [ \ud835\udf06\ud835\udc5e\ud835\udc60 \ud835\udf06\ud835\udc51\ud835\udc60 ] + \u03c9\ud835\udc52 [ 0 \u22121 1 0 ] [ \ud835\udf06\ud835\udc5e\ud835\udc60 \ud835\udf06\ud835\udc51\ud835\udc60 ] (4) [ \ud835\udc49\ud835\udc5e\ud835\udc5f \u2032 \ud835\udc49\ud835\udc51\ud835\udc5f \u2032 ] = \ud835\udc5f\ud835\udc5f \u2032 [ \ud835\udc56\ud835\udc5e\ud835\udc5f \u2032 \ud835\udc56\ud835\udc51\ud835\udc5f \u2032 ] + \ud835\udc51 \ud835\udc51\ud835\udc61 [ 0 \ud835\udf06\ud835\udc5f \u2032 ] + (\ud835\udf14\ud835\udc52 \u2212 \ud835\udf14\ud835\udc5f) [ 0 \u22121 1 0 ] [ 0 \ud835\udf06\ud835\udc5f \u2032 ] (5) [ \ud835\udf06\ud835\udc5e\ud835\udc60 \ud835\udf06\ud835\udc51\ud835\udc60 0 \ud835\udf06\ud835\udc5f \u2032 ] = [ \ud835\udc3f\ud835\udc60 0 0 \ud835\udc3f\ud835\udc60 \ud835\udc3f\ud835\udc5a\ud835\udc60 0 0 \ud835\udc3f\ud835\udc5a\ud835\udc60 \ud835\udc3f\ud835\udc5a\ud835\udc60 0 0 \ud835\udc3f\ud835\udc5a\ud835\udc60 \ud835\udc3f\ud835\udc5f \u2032 0 0 \ud835\udc3f\ud835\udc5f \u2032 ] [ \ud835\udc56\ud835\udc5e\ud835\udc60 \ud835\udc56\ud835\udc51\ud835\udc60 \ud835\udc56\ud835\udc5e\ud835\udc5f \u2032 \ud835\udc56\ud835\udc5e\ud835\udc5f \u2032 ] (6) \ud835\udc51 \ud835\udc51\ud835\udc61 \ud835\udf06\ud835\udc5f \u2032 + \ud835\udf06\ud835\udc5f \u2032 \ud835\udf0f = \ud835\udc3f\ud835\udc5a\ud835\udc60 \ud835\udf0f \ud835\udc56\ud835\udc51\ud835\udc60 (7) \ud835\udf14\ud835\udc60\ud835\udc59 = \ud835\udf14\ud835\udc52 \u2212 \ud835\udf14\ud835\udc5f = \ud835\udc3f\ud835\udc5a\ud835\udc60 \ud835\udf0f\ud835\udf06\ud835\udc5f \ud835\udc56\ud835\udc5e\ud835\udc60 (8) \ud835\udf06\ud835\udc5f \u2032 = \ud835\udc3f\ud835\udc5a\ud835\udc60\ud835\udc56\ud835\udc51\ud835\udc60 (9) \ud835\udc47\ud835\udc52\ud835\udc5a = \ud835\udc43 2 \ud835\udf06\ud835\udc5f \u2032 \ud835\udc3f\ud835\udc5a\ud835\udc60\ud835\udc56\ud835\udc5e\ud835\udc60 \ud835\udc3f\ud835\udc5f \u2032 (10) The equation that describes the mechanical dynamics of the machine is given in (11)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000289_vppc.2009.5289719-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000289_vppc.2009.5289719-Figure8-1.png", + "caption": "Figure 8. The size of the vehicle model used for simulations", + "texts": [ + " It is expected that the influence on the steering system of the driver becomes large because the failure at the time of a rectilinear-propagation run may induce a big imbalance torque between the right and left wheels. Then, first, the failure transient analyses at the time of the rectilinear-propagation acceleration which may induce the most dangerous traffic accident are carried out. In the transient analyses, \u03b2 and \u03b3 which are related to the drift and spin of vehicles (Fig. 7), and the vehicle trajectory changes with these parameters are investigated using the Prius (Fig. 8 and Table III) as an analysis vehicle. B. Failure Caused during Rectilinear-Propagation while Accelerating on Dry Roads Fig.9 shows transient chracteritics when the front-wheel drive system of the FRID EV (henceforth, the conditions of failure are the same), and the rear-left wheel drive system of Fig.1(a) and the front-left wheel drive system of Fig. 1(b) failed at time t1 during rectilinear-propagation while accelerating on a dry road. In two in-wheel motor drive type EVs, with this failure, the longitudinal (driving) force which propels a left-hand side wheel is lost instantaneously" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002459_robio.2017.8324729-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002459_robio.2017.8324729-Figure3-1.png", + "caption": "Fig. 3: Catenary curve applied to the mathematical model for the cable", + "texts": [ + " In general, Catenary number is a key parameter to determine the shape of the cable and it can be derived by using the tension of the both endpoints of the cable. However, in practice, it is very difficult to obtain proper tension in the case of a thin and light cable because of the sensor noise and resolution. Therefore, we propose to use only positions of the UAVs that are located on the endpoints of the cable. In general, Catenary curve is expressed as in eq. (1) that is determined in the Cartesian coordinates \u03a3 whose origin is located on the lowest point of the loose cable as shown in Fig. 3 [8]\uff0e C : z = a ( cosh x a \u2212 1 ) (1) where a [m] denotes Catenary number which is an important variable to determine the shape of Catenary curve. The Catenary number is generally expressed by the tension T of the endpoint and angle \u03b8 of the cable as: a = Tcos\u03b8 wg (2) where w[kg/m] denotes linear density of the cable and g[m/s2] denotes gravitational acceleration. In this paper, we propose a method to derive the above Catenary number without using tension information because of the problem of sensor noise and resolution. Its detail derivation is described in section III-B. As shown in Fig. 3, positions of endpoints P and Q are defined as (xp, zp) and (xq, zq) with respect to the Cartesian coordinates \u03a3, respectively. Then the valid shape of the cable can be determined from eq. (1) with the condition of xq < x < xp. However, in practice, the position of the lowest point of the loose cable changes while UAVs are flying because both or one side of the endpoint position change. Besides, it is technically difficult to measure the lowest point during the flight operation. Therefore, this paper derives a mathematical model of Catenary curve based on the coordinates \u03a3Base whose origin is located on the endpoint of one side instead of using the information of the lowest point position as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001586_lars.2010.36-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001586_lars.2010.36-Figure3-1.png", + "caption": "Fig. 3. Diagram for |\u2206\u0303l| computation", + "texts": [ + " However, a estimated value of |\u2206\u0303l| can be calculated using a bezier curve of third degree [7] between consecutive samples, where the initial orientation of the curve is equal to the robot orientation in the previous sample and the final orientation is equal to the actual robot orientation. Each control point of the curve is located in a straight line formed by the position and angle of the robot in each sample, with a distance of the robot position equal of a third of the euclidean distance between the samples, as shown in Fig. 3. The estimated value |\u2206\u0303l| can be achieved by numerical integration of the bezier curve. Note that the control points should agree with the movement direction of the robot (forward or backward). The system non-linearity is modeled by two symmetric dead zones, one for each system input. The dead zone is described by Fig. 4 and can be written as: m(t) = u(t) \u2212 D if u(t) > D u(t) + D if u(t) < \u2212D 0 otherwise (10) where D > 0 is the dead zone constant and dead zone slopes have unitary gain (its always possible to fix one parameter in the Hammerstein model)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001673_sav-2010-0501-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001673_sav-2010-0501-Figure9-1.png", + "caption": "Fig. 9. Second optimization step.", + "texts": [], + "surrounding_texts": [ + "Currently, the lower arm suspension of a vehicle is generally manufactured from steel. Figure 5 illustrates a vehicle suspension containing two suspension arms. The applicability of a weight optimization algorithm to this part already in service is verified. We propose a direct method of weight optimization to obtain a lighter automobile part with a safer fatigue life and a natural frequency away from the PSD frequency range. The optimization is achieved by removing elements of the automobile part that have the minimum energy density sum variation. We developed a strategy that determined the critical elements and their coordinates, and allowed the isolation of the element with the maximum sum of positive variation of the strain density energy. Contrary to the static case, the rough maximum value of the strain energy density does not necessarily occur in the critical element. Our method allows the application of the Newton-Raphson algorithm to only one element instead of applying it to all the mesh points in the structure. In order to extract the number of cycles to rupture, the nonlinear Mansson-Coffin equation is used. This filter generalizes the case where the excitation is multipoint and shifted in time, giving a tangle of the mesh material element signals. This filter is based on the following algorithm: xc and zc are the critical element coordinates having : M(ic, jc) = Mnc ( Max ( tf\u2211 m=0 \u2206U(m,n) )) where: m = time parameter k = simulation duration tf =\u22822 k= k! 2!(k\u22122)! n = spatial parameter to make ic and jc in the same column nc \u2206U(m,n) = U(l, n) \u2212 U(p, n), when U(l, n) > U(p, n) Unlike the static case where the absolute value of strain energy causes rupture through fatigue, it is the variation of the strain energy density which causes rupture by fatigue in the dynamic stress case. To apply the Miner law, the rainflow cycle must be extracted. The problem of the direct passage of the PSD of \u03c3(t) to the counting of the rainflow cycles is then solved by the rigorous application of a theorem based on the definition of a rainflow cycle [25] and on the Markov chain theory. A rainflow cycle, as illustrated in Fig. 6 can be mathematically characterized in the following way: let us consider \u03c3(t) where t \u2208 [0, T ] and the stress maximum M i of level K occur at time ti. The limits (m\u2212 i ,Mi) and (Mi,m + i ) can be defined where: \u2013 m\u2212 i is the minimum of \u03c3(t) and Mi is between the last passage to the negative slope of \u03c3(t) and the maximum Mi. This minimum is on the left of Mi and occurs at time t\u2212i \u2013 m+ i is the minimum \u03c3(t) which is between Mi and the first passage to positive slope of \u03c3(t) by the level K . This minimum is on the right of Mi and occurs at time t+ i . If there is no passage of \u03c3(t) by the level K before or after time t i, then t\u2212i = 0 and t+i = T . The rainflow extracted at time ti is then defined either as the extent (mrfc i ,Mi), or (Mi,m rfc i ). This minimum mrfc i is given by applying the condition: mrfc i = { max(m\u2212 i ,m+ i ) if t\u2212i > 0 m+ i else (15) Anomalies which may be present in the strain energy density signal of the critical element can be avoided if the rainflow cycles are counted using the Markov method. Successive equalities of two values of strain energy density at consecutive moments could be a problem. The anomaly is corrected by the WAFO Matlab interface dat2tp. When the loadings are composed of cycles of various amplitudes and different average values, it is necessary to measure the total damage produced by these cycles. Fatemi and Yang [25] present a complete review of the laws of damage calculation which were developed from the linear damage rule suggested by Palmgren. The mathematical formulation under which it is currently known was proposed by Miner and it is expressed as Eq. (16): D = n\u2211 i=1 ni Ni (16) Random stress history is described as a sequence of blocks of constant amplitude. Each block i is composed of Ni cycles of amplitude. Partial fatigue life Ni corresponding to this stress amplitude is determined from the Wo\u0308hler curve or using the strain energy density approach. Failure is predicted when damage D is equal to 1. It is thus necessary to find the number of times the random loading is repeated before D is equal to 1(?). It is necessary to find the Bf number which one must multiply by D to reach rupture. B f is calculated using Eq. (17): Bf = 1\u2211 i ni Ni because BfD = Bf (\u2211 i ni Ni ) = 1 (17) In this study the number of cycles to rupture is calculated from the density of uniaxial strain energy (SENER from S = strain, ENER = energy) equivalent of the multiaxial case. SENER reaches a maximum value of 150 J/m 3. The shortest fatigue life number NR of the part is calculated using the Miner law to be 5.2 \u00d7 10 7 cycles caused by continuous excitement. The part weighs about 3.5 kg. Low strain density energy areas are shown in Fig. 7. An initial analysis already enables us to identify areas with low SENER variation. These areas can be removed without affecting the part\u2019s performance or reducing significantly its life expectancy or its natural frequencies. Abaqus 6.4 software, calculates natural frequencies based on the theory of elasticity and uses a deformable body such as a non negligible mass spring. Frequency analysis of the most complex structures is laborious and depends on complex system conditions such as: boundary conditions, static equilibrium conditions of the excited system and system mass distribution. In this paper, the static equilibrium state as an embedded system where the specimen is forced to no motion in the six degrees of freedom and elasticity is due to specimen weight. The first two modes have frequencies of 416 Hz and 1754 Hz, respectively (Table 3). The reference geometry weight is around 35 kg. The reference geometry of the first two modes is far from the PSD frequency range describing the road profile. Figure 7 shows critical areas as well as areas of low SENER variation. It is known that natural frequencies are proportional to flexural rigidity and inversely proportional to part mass as described above. As material in low intensity SENER variation areas is removed. The part\u2019s geometry changes and its mass is reduced (Fig. 8). The element removal is symmetrical and must respect the objective function: k\u2211 i=1 (SENER(t)\u2212SENER(t\u2032)each element Max( k\u2211 i=1 (SENER(t)\u2212SENER(t\u2032)each element) \u3008RJ with SENER(t) > SENER(t) t and t\u2032are time parameters n = number of mesh elements k = n! 2!(n\u22122)! (18) The rejection ratio RJ is directly linked to the notion of fatigue. In order to avoid the extraction of rainflow cycles, RJ takes into account the SENER variation, simplifies calculations and adjusts to loading randomness. It also takes into account fatigue life calculations in all mesh elements and optimizes the part\u2019s weight in a reasonable time period. SENER fluctuations are due to mesh ratio variation, directly linked to the integration implicit scheme used in this study. The mass decreases by 4.6% from 3.5 kg to 3.34 kg. The first natural frequency mode varies from 416 Hz to 439 Hz and the second mode from 1754 Hz to 1883 Hz, moving away from the frequency range of the PSD (Fig. 8). Predicted fatigue life of the part is about 5 \u00d7 107 to 108. Two other stages of material removal were carried out to reduce weight. Figures 9 and 10 show second and third stage optimization results. The part\u2019s mass decreases without affecting either its fatigue life or its first mode significantly. The mass decreases from 3.5 kg to 3.1 kg, which corresponds to a 11.42% weight loss. All critical elements are located in the embedded area. Figure 11 shows the geometrical evolution of the part during the optimization process. Numerical results are summarized in Table 4. The results of the optimization study are comparable to Haiba et al.\u2019s [1] work on weight optimization of a simple mechanical part. However, their fatigue study is based primarily on simple experimental method of the Wo\u0308hler curve extrapolation in the calculation of fatigue life. They took into account only one parameter: stress variation. Our approach takes into account two parameters: stress and strain, using the multiaxial criterion of the strain energy density equivalent to the uniaxial case. Haiba et al. [1] defined the fatigue rejection ratio as: (Fatigue life)min (Fatigue life)max < RR where (Fatigue life)min and (fatigue life)max are the minimum and maximum fatigue life of the structure, respectively. In their study the critical areas can be strengthened by a detailed optimization study using genetic algorithms and by the use of the rejection ratio developed in the present study. Our proposed optimization approach for making parts lighter is analogous to the method of Haiba et al. [1]. Unlike in Haiba et al.\u2019s [1] work, the rejection ratio developed in the present study is linked directly to the notion of fatigue. It calculates the fatigue life of all mesh elements, and uses a more rational criterion of the strain density energy. The rejection ratio developed here simplifies the calculations and adapts to load random aspect. It avoids rainflow cycle extraction and fatigue life calculation in all elements. It also allows the optimization of a part\u2019s weight compared to its fatigue life and its eigen mode in a reasonable time." + ] + }, + { + "image_filename": "designv11_33_0000777_s0025654411040017-Figure10-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000777_s0025654411040017-Figure10-1.png", + "caption": "Fig. 10. Fig. 11.", + "texts": [], + "surrounding_texts": [ + "By setting \u03b4 = 0 in Eqs. (3.6) and (3.7) and by introducing new variables z1 and z2 by formulas z1 = z + a1\u03d5 and z2 = z \u2212 a2\u03d5, we obtain the equations z\u2022\u20221 + a11(c1z1 + \u03b31z \u2022 1) + a12(c2z2 + \u03b32z \u2022 2) = 0, z\u2022\u20222 + a21(c1z1 + \u03b31z \u2022 1) + a22(c2z2 + \u03b32z \u2022 2) = 0, (4.1) MECHANICS OF SOLIDS Vol. 46 No. 4 2011 a11 = 1 + a2 1 \u2212 a1f1, a12 = 1 \u2212 a1a2 \u2212 a1f2, a21 = 1 \u2212 a1a2 + a2f1, a22 = 1 + a2 2 + a2f2. (4.2) The dimensionless variables ak and fk (k = 1, 2) in formulas (4.2) are the ratios of the corresponding variables introduced in the preceding sections to the radius of inertia \u03c1; i.e., we assume that the radius of inertia \u03c1 is equal to unity. The variables z1 and z2 have the following physical meaning. These are small deviations of the points O1 and O2 from their equilibrium positions described in Sec. 3. Equations (4.1) and (4.2) and relations (1.1) with the use of the Lyapunov theorem permit solving the above stability and instability problem for any fixed v. Let us write out the characteristic polynomial of the linear system (4.1) as p4 + p3(a11\u03b31 + a22\u03b32) + p2(a11c1 + a22c2 + \u03b31\u03b32\u0394) + p(\u03b31c2 + \u03b32c1)\u0394 + c1c2\u0394 = 0, (4.3) where \u0394 = a11a22 \u2212 a12a21. Applying the Routh\u2013Hurwitz theorem to the polynomial (4.3), we arrive at the following result. Proposition 1. For the asymptotic stability of the equilibrium in the case of rectilinear and uniform motion of the vehicle, it is necessary and sufficient that the rolling friction coefficients f1 and f2 (and hence the velocity v of the vehicle translational motion determined by (1.1)) satisfy the inequalities a11\u03b31 + a22\u03b32 > 0, a11c1 + a22c2 + \u03b31\u03b32\u0394 > 0, \u0394 > 0, (4.4) (a11\u03b31 + a22\u03b32)[a11\u03b32c 2 1 + a22\u03b31c 2 2 + \u03b31\u03b32\u0394(\u03b31c2 + \u03b32c1)] \u2212 (\u03b31c2 + \u03b32c1)2\u0394 > 0, \u0394 = a11a22 \u2212 a12a21 = a(a + f2 \u2212 f1), a = a1 + a2. (4.5) The dimensionless parameters aij (i, j = 1, 2) are determined by formulas (4.2). Consider several particular cases where Proposition 1 applies. Proposition 2. If the system is completely symmetric, i.e., if \u03b3k = \u03b3 = 0, ck = c, fk = f, k = 1, 2, a1 = a2 = a0, (4.6) then inequalities (4.4) hold for each f ; i.e., the asymptotic stability of the vehicle equilibrium considered above is preserved for any wheel rolling friction coefficients and hence for any velocities of its translational motion. The proof is by a straightforward verification of inequalities (4.4) under conditions (4.6). In this case, it follows from (4.5) that \u0394 = a2 > 0 and the first two inequalities in (4.4) are satisfied, because they are reduced to the following ones: 2\u03b3(1 + a2 0) > 0, 2c(1 + a2 0) + \u03b32a2 > 0. The last inequality in (4.4) has the form c2(1 \u2212 a2 0) 2 + 4\u03b32c2a2 0(1 + a2 0) > 0. Thus, the stability conditions (4.4) are satisfied for any f \u2265 0, and the proof of Proposition 2 is complete. Proposition 3. Assume that all conditions (4.6) except for the last are satisfied; i.e., a1 =a2 (the case of incomplete symmetry; i.e., the car body center of mass is displaced). Then, for a1 < a2, the stability conditions (4.4) are satisfied for every f \u2265 0, and for a1 > a2, the stability is violated for f > f0 = 1 a1 \u2212 a2 [ 2 + a2 1 + a2 2 + \u03b5(a1 + a2)2 \u2212 \u221a \u03b52(a1 + a2)4 + 4(a1 + a2)2 ] > 0, \u03b5 = \u03b32 c . (4.7) Remark. The quantity f0 in inequality (4.7) can be arbitrarily small for some values of the system parameters. For example, as \u03b5 \u2192 0 (small damping or large rigidity of the wheel) and for a1a2 = 1 and a1 > 0, we have f0 = (a1 \u2212 1)(1 + a2 1)/(a1 + a2 1), which can be arbitrarily small as a1 \u2192 1. MECHANICS OF SOLIDS Vol. 46 No. 4 2011 The proof is by a straightforward verification. Under the conditions given in Proposition 3, the instability inequalities (4.4) become \u03b3[2 + a2 1 + a2 2 + (a2 \u2212 a1)f ] > 0, c[2 + a2 1 + a2 2 + (a2 \u2212 a1)f ] + \u03b32(a1 + a2)2 > 0, \u0394 = (a1 + a2)2 > 0, x2 + 2x\u03b5(a1 + a2)2 \u2212 4(a1 + a2)2 > 0, x = 2 + a2 1 + a2 2 + (a2 \u2212 a1)f. (4.8) For a2 > a1, these inequalities are necessarily satisfied for any f > 0. (For the first three inequalities in (4.8), this is obvious, and for the last, this follows from the inequality x > 2 + a2 1 + a2 2 > 2(a1 + a2).) But if a2 < a1, then the first inequality in (4.8) is violated for f > f1 = (2 + a2 1 + a2 2)/(a1 \u2212 a2), the second is violated for f > f2 > f1, the third is always satisfied, and the last is violated for f \u2208 [f0, f00], where f0 is given by the formula in (4.7) and f00 is given by the same formula with the plus sign at the radical. It is easily seen that 0 < f0 < f1 < f00, which implies that all inequalities in (4.8) are violated for f > f0. The proof of Proposition 3 is complete. Proposition 4. Let f2 = 0, \u03b30 = 0, and f1 = f > 0; i.e., assume that the fore (driving) wheel does not experience any rolling resistance and there is no dissipation in it. In this case, if a1a2 < 1, then the instability arises for f > a1, and if a1a2 > 1, then the instability arises for f > (a1a2 \u2212 1)/a2. Proposition 5. Let f1 = 0, \u03b31 = 0, and f2 = f > 0; i.e., assume that the rear (driven) wheel does not experience any rolling resistance and there is no dissipation in it. In this case, if a1a2 < 1, then the instability arises for f > (1 \u2212 a1a2)/a1, and if a1a2 > 1, then the system is stable for any f > 0. The proof of Propositions 4 and 5 is by a straightforward verification of inequalities (4.4) in Proposition 1 under the assumptions stated above. Let \u03b31 > 0, \u03b32 = 0, and \u03b32 > 0. Then the following assertion holds. Proposition 6. Under the above conditions, the asymptotic stability domain D1 in the plane of the parameters (f1, f2) is given by the following inequalities. If 0 < a1a2 < 1, then 0 < f2 < 1 \u2212 a1a2 a1 , 0 < f1, c1a1f1 \u2212 c2a2f1 < c0, f1 \u2212 f2 < a1 + a2. (4.9) If 1 < a1a2 < \u221e, then 0 < f1 < a1a2 \u2212 1 a2 , 0 < f2, c1a1f1 \u2212 c2a2f2 < c0, (4.10) c0 = c1(1 + a2 1) + c2(1 + a2 2). (4.11) The proof of Proposition 6 is by a straightforward verification of inequalities (4.4) under the restrictions given in the statement. Note that if one of the inequalities in conditions (4.10), (4.11) is not satisfied, then the vibrations under study are unstable. Now let \u03b31 = \u03b32 = 0; i.e., assume that there is no dissipation in the wheels. In this case, Eq. (4.3) is biquadratic, the stability of system (4.1) can be only nonasymptotic, and the conditions for such stability are equivalent to the conditions that this biquadratic equation has only pure imaginary roots. These considerations lead to the following result. Proposition 7. For \u03b31 = \u03b32 = 0, the stability domain D0 in the plane of parameters (f1, f2) is given by the following inequalities: f1 > 0, f2 > 0, c0 > c1a1f1 \u2212 c2a2f2, a1 + a2 > f1 \u2212 f2, [c0 \u2212 (c1a1f1 \u2212 c2a2f2)]2 > 4c1c2(a1 + a2)(a1 + a2 \u2212 f1 + f2). (4.12) The proof of Proposition 7 follows from well-known necessary and sufficient conditions for the existence of pure imaginary roots of a biquadratic equation. Proposition 8. The inclusion D1 \u2282 D0 holds. MECHANICS OF SOLIDS Vol. 46 No. 4 2011 The proof of Proposition 8 is by a straightforward substitution of the extreme values in the linear inequalities (4.9) or (4.10) into inequalities (4.12) and verification that the latter are satisfied. Remark. Proposition 8 means that this nonsymmetric addition of dissipation narrows down the stability domain in the plane of the parameters (f1, f2). Graphically (schematically), the domains D1 and D0 in various possible cases are shown in Figs. 2\u201315, where digit 1 corresponds to the line f1 \u2212 f2 = a1 + a2, digit 2 corresponds to the line c1a1f1 \u2212 c2a2f2 = c0 = c1(1 + a2 1) + c2(1 + a2 2), digit 3 corresponds to the line f2 = |1 \u2212 a1a2|/a1, digit 4 corresponds to the line f1 = |1 \u2212 a1a2|/a2, digit 5 corresponds to the line f1 \u2212 f2 = \u2212|1 \u2212 a1a2|(a1 + a2)/(a1a2), and digit 6 corresponds to the parabola (c0 \u2212 c1a1f1 + c2a2f2)2 \u2212 4c1c2(a1 + a2)(a1 + a2 \u2212 f1 + f2) = 0. Figures 2, 4, 6, 8, 10, 12, and 14 correspond to the case of \u03b31 = \u03b32 = 0, while Figs. 3, 5, 7, 9, 11, 13, and 15 correspond to the case of \u03b31 = 0, \u03b32 = 0 or \u03b31 = 0, \u03b32 = 0 for all possible distinct values of the parameters a1, a2, c1, and c2. Now consider the case of weighted wheels, m1 = 0 and m2 = 0. Since we assume that the car body mass is m = 1, the quantities m1 and m2 are the respective dimensionless ratios of the wheel masses to the car body mass. Simple calculations show that the form of the vibration equations (4.1) remains the same and the coefficients aij (i, j = 1, 2) are determined by the functions a11 = \u03c12 \u2212 (a1 + am2)(\u2212a1 + f1) \u03c12 0 , a12 = \u03c12 \u2212 (a1 + am2)(a2 + f2) \u03c12 0 , a21 = \u03c12 + (a2 + am1)(\u2212a1 + f1) \u03c12 0 , a22 = \u03c12 + (a2 + am1)(a2 + f2) \u03c12 0 , where \u03c12 0 = m0\u03c1 2 and m0 = 1 + m1 + m2. Then one can readily compute that \u0394 = a11a22 \u2212 a12a21 = a(a + f2 \u2212 f1)/\u03c12 0. Proposition 1 remains valid under the above variations in aij , i, j = 1, 2, and \u0394. In a similar way, one can restate Propositions 2\u20138. Equations (4.1) and (4.2) describing the vehicle vibrations under the action of the rolling friction forces admit the following analogies (interpretations) to well-known systems in mechanics. Each of the coefficient matrices multiplying zk and z\u2022k (k = 1, 2) in (4.1) with (4.2) taken into account consists of MECHANICS OF SOLIDS Vol. 46 No. 4 2011 two parts. The first part (it is said to be unperturbed) does not contain the friction coefficients f1 and f2. The second (perturbed) part depends on these coefficients linearly. As follows from the above derivation of the equations of motion, the origin of the perturbed part significantly depends on the following two external causes. One of them is the interaction between the deformable wheels and the rough road, which causes the rolling resistance torques Mk1 and Mk2; the second cause is the torque M1 from the engine (external energy source) to the driving wheels ensuring their (and the driven wheels) permanent rolling if there are resistance torques mentioned above. If these causes are absent, then the system motion occurs for f1 = f2 = 0, and Eqs. (4.1) contain only the first parts of the above-mentioned matrices for which the solutions have the property of (asymptotic) stability (a conservative system MECHANICS OF SOLIDS Vol. 46 No. 4 2011 under the action of some forces with complete or incomplete dissipation). But if f2 1 + f2 2 = 0, then these causes result in the appearance of additional terms in Eqs. (4.1), which, according to [6], can be interpreted as positional (circular) forces (for the corresponding terms with z1 and z2); gyroscopic and nonconservative (depending on the velocities) forces (for the corresponding terms with z\u20221 and z\u20222); conservative positional forces (for the corresponding terms with z1 and z2); and dissipative (Rayleigh) MECHANICS OF SOLIDS Vol. 46 No. 4 2011 MECHANICS OF SOLIDS Vol. 46 No. 4 2011 forces (for the corresponding terms with z\u20221 and z\u20222). In the general case, such a decomposition of forces was performed in [7]. Thus, the existence of the above-mentioned external causes leads to the appearance of additional forces of the described structure, whose destructive character is well known in mechanics of gyroscopic systems, structural mechanics (in particular, in mechanics of bridges), and other adjacent fields; it is described in the literature [8\u201310]. Even if there forces do not actually lead to the vehicle instability, they still significantly change the frequencies of its vibrations in the vertical plane, and this fact cannot be ignored when solving the corresponding problems. These facts were already considered earlier by the authors in the paper [11]." + ] + }, + { + "image_filename": "designv11_33_0001223_0954406211409805-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001223_0954406211409805-Figure4-1.png", + "caption": "Fig. 4 The turning machine", + "texts": [ + " Resorting to the presented MCS, P\u00f0Md 4Mr \u00de j can be conveniently computed. As for KPR, it is a bilateral probability constraint problem (6). The direct use of the above MCS is impossible, unless some changes are made, shown as follows RKDPR \u00bcmin P\u00f0 5 \"5 \u00de j \u00bcmin\u00bdP\u00f0 \"5 \u00de P\u00f0 \"5 \u00de j \u00bcmin\u00bdP\u00f0 \"5 \u00de \u00fe P\u00f0 \" \u00de 1 j \u00f011\u00de From (9), the performance function of KDPR is G1 \u00bc \" and G2 \u00bc \u00fe \", respectively. Using the above MCS, P\u00f0 \"5 \u00de and P\u00f0 \" \u00de can be worked out. Then, RKDPR in (9) can be computed. As shown in Fig. 4, the turning machine is assembled on the machine frame at the assembling Point1 and Point2. The driver (composed of Cylinder1 and Cylinder2) generates driving force to overcome the RF derived from resistance loads. The driving speed is about 0.015 m/s and last 8 s. During the running process, the DER and KPR of the angle displacement errors are concerned. The mean values and standard deviations of the main influential factors are given in Table 1. Proc. IMechE Vol. 225 Part C: J. Mechanical Engineering Science at PURDUE UNIV LIBRARY TSS on May 19, 2015pic" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002850_icra.2018.8461156-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002850_icra.2018.8461156-Figure2-1.png", + "caption": "Fig. 2. Planar coordinate definition for a generic unstructured environment.", + "texts": [ + " If the spatial distribution of objects in the environment is modeled as a continuous function of the body-referenced viewing angle \u03b3, the optic flow field can be written as, Q\u0307(\u03b3,x) = \u2212\u03c8\u0307 + [d(\u03b3,q)]\u22121(u sin \u03b3 \u2212 v cos \u03b3), (1) where d(\u03b3,q) is the radial distance to the nearest point in the visual field at angle \u03b3, u and v are body-referenced forward and lateral velocity components respectively, x = {q, q\u0307} is the vehicle state with q = {x, y, \u03c8} as the vehicle pose with respect to the environment (Fig. 2). \u00b0 For motion restricted to a plane, it is well-known that wide-field optic flow patterns are spatially periodic that reside in L2[\u2212\u03c0, \u03c0], the space of square-integrable and piecewise-continuous functions, and can be modeled using the first N harmonics of the Fourier series as [18], Q\u0307WF (\u03b3,x)\u2248a0 2 + N\u2211 n=1 (an cosn\u03b3+bn sinn\u03b3) , an= 1 \u03c0 \u222b \u03c0 \u2212\u03c0 Q\u0307(\u03b3,x) cosn\u03b3 d\u03b3, bn= 1 \u03c0 \u222b \u03c0 \u2212\u03c0 Q\u0307(\u03b3,x) sinn\u03b3 d\u03b3. (2) For vehicle flight close to the centerline of a straight corridor, the tangential optic flow profile resembles a sinewave like pattern, with the peak amplitude proportional to the forward speed and inversely proportional to the perpendicular distance of the vehicle to the wall" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002941_icelmach.2018.8506824-Figure10-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002941_icelmach.2018.8506824-Figure10-1.png", + "caption": "Fig. 10. Temperature difference distributions between the initial and the model that considers radiation. The case without radiation, even air-flow speed and homogenous loss distribution is used as reference and compared with the respective cases with radiation and (a) homogeneous loss distribution and even air-flow speed, (b) uneven air-flow speed and homogeneous loss distribution, (c) even air-flow speed and inhomogeneous loss distribution, (d) uneven air-flow speed and inhomogeneous loss distribution.", + "texts": [ + " The investigated models yield different results, with discrepancies that can reach 10% in temperature values. With more realistic model representations, the initial image of temperature distribution can vary significantly. On a second step, the simplified radiation model is introduced in the boundary definitions. The same cases are investigated considering the loss and air-flow distribution. It can be observed that the addition of radiative heat transfer affects the overall temperature distribution by shifting it to lower levels, as depicted in Fig. 10(a). The areas where the shift is most pronounced are at the top and bottom parts which are furthest away from the cooling channels, as indicated in Fig. 10(a) with region (I). On the side regions, the shift is less pronounced, as indicated in Fig. 10(a) with (H). Comparing Fig. 10(b) and Fig. 10(a) it can be seen that the lower-shifting is still present, and the effect on the temperature distribution incurred by the uneven flow speed in the axial channels is still present. It is though observed that the regions where the lower-shifting occurs cover a slightly larger part of the stator, indicated in Fig. 10(b) with regions (J) and (K). The effect of partial channel blockage is also here highlighted. As the inhomogeneous loss distribution is accounted for, as illustrated in Fig. 10(c), elevated temperatures are still encountered at the tooth tips, close to the vertical mid-line. However, the radiative heat dissipation affects the outer surface temperature of the stator by lowering it, when compared to Fig. 9(c). This can be observed by comparing the ranges presented in Fig. 9(c) and Fig. 10(c) in combination with comparing Fig. 10(c), region (L) and Fig. 9(c), region (E). Finally, the temperature difference compared to the reference case, as inhomogeneous loss distribution, uneven flow distribution and radiative heat transfer is incorporated in the model, is presented in Fig. 10(d). Here it can be seen that radiation further affects the temperature distribution when compared to Fig. 9(d). Effects from the inhomogeneous loss distribution, such as the elevated temperatures in region (M) in Fig. 10(d) are present, as well as the effects from the air flow distribution, as indicated by region (N) in Fig. 10(d). Overall, the skewed thermal profile is still present, while the peak temperature is shifted lower by 0.02p.u. and the lowest temperature is shifted lower by 0.009p.u. What can be noted is that the lowshifting of 0.02p.u. occurs at the tooth tips and that there is a general shift lower for the whole teeth region. Thus, the inclusion of radiative heat transfer in the model, affects the temperature distribution in the entire stator body. In situations when temperature sensors are placed in the slots, this will naturally have an influence on the conclusions that can be drawn when comparing simulation results with measurements" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003648_j.matpr.2019.06.208-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003648_j.matpr.2019.06.208-Figure2-1.png", + "caption": "Fig. 2. (a) The geometry and moving path of laser heat source; (b) mesh elements of the work piece model", + "texts": [ + " ( ) (5) where E is heat energy, AC is absorption coefficient. The value of parameters in the equations (4) and (5) are shown in next section. 3.1 Numerical simulation setup 1764 P. Ninpetch and P. Kowitwarangkul / Materials Today: Proceedings 17 (2019) 1761\u20131767 Numerical simulation was performed by commercial simulation software ANSYS 18.1. The material used in this study is AISI 304 stainless steel plate with a dimension of 50x70x2 mm3. The geometry and moving path of laser heat source are shown in Fig.2(a). A computational mesh of around 70,000 cells is used for the simulation. The mesh is shown in Fig.2(b). The parameters of moving laser heat source are listed in Table 1. The temperature history during the process is measured by the probes at six points as shown in Fig.1(a). The distance between each point is 1 mm. The AISI 304 stainless steel plate with a dimension of 50x70x2 mm3 were used as specimens in this experiment. Fig.3(a) shows the laser source device, JenLas\u00ae fiber ns 50, at Center Innovation of Design and Engineering for Manufacturing (Col-DEM), KMUTNB which has maximum power of 50 W", + " At 70 sec and 101 sec, the comet-shape area of temperature distribution are expand due to the heat accumulation during the back and forth moving of laser heat source on the specimen. Fig.4. Temperature distribution in the numerical specimen model from the scenario 1 at time (a) 30 sec; (b) 70 sec; (c) 101 sec. The numerical simulation results from test scenario 1 and 2 were validated with the experimental results as shown in Fig.5. The graphs show the temperature history at the probe point no.1 (see Fig.2). The temperature peak took place after the moving of heat source passed that area. The temperature profile of scenario 1 (hatch spacing 0.5 mm) is slightly higher than that of scenario 2 (hatch spacing 1 mm), e.g., the second peak of scenario 1 and 2 are 184 oC and 176 oC respectively. The comparison of the results from numerical simulations and experiments shows good agreement with minor difference. The peak of the results from experiment are slightly lower than that of the simulation. Fig. 6 shows the effect of moving laser heat source parameters, e", + "68 mm/s, the first temperature peak of the graph is increased from 128 oC to 194 oC. 1766 P. Ninpetch and P. Kowitwarangkul / Materials Today: Proceedings 17 (2019) 1761\u20131767 Different hatch spacing slightly affects the temperature history at probe point no.1 as can be seen from Fig.6. At the second temperature peaks in the graph, with the increasing of hatch spacing from 0.5 mm to 1.0 mm, the temperature peak is decreased for around 5-10 oC. Fig.7 shows the temperature histories at six probe points (see Fig. 2) from the numerical simulations of scenario 3 and 5 which have power intensity of 200 W/mm2 and 480 W/mm2 respectively. At the probe point no.6 which is the position at the first forth path of laser beam, the first peak temperature was increased from around 1200 oC to 2500 oC after the power intensity was increased from 200 W/mm2 to 480 W/mm2. At the probe point no.1 to no.5, the temperature of both peaks is not so much different, while at the probe point no.6, the second temperature peak is much lower than the first peak since the laser source moved pass through that point only the first forth path (see Fig. 2). 5.Conclusion In this research, the effects of the moving laser heat source parameters such as speed, direction, hatch spacing and power intensity on the temperature distribution of the AISI 304 stainless steel plate were investigated through P. Ninpetch and P. Kowitwarangkul / Materials Today: Proceedings 17 (2019) 1761\u20131767 1767 the numerical simulations and experiments. The effect of changeable direction moving laser heat source with low power intensity in the range of 200-480 W/mm2 were studied through the numerical simulation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002260_j.ifacol.2017.08.764-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002260_j.ifacol.2017.08.764-Figure1-1.png", + "caption": "Fig. 1. The quadrotor-manipulator system with corresponding frames", + "texts": [ + " Weighted adjacency matrix A is defined such that aij = aji is a positive weight if (i, j) \u2208 E , while aij = 0 if (i, j) /\u2208 E . Let Laplacian matrix L = [lij ] \u2208 R n\u00d7n associated with A be defined as lii = \u2211n j=1,j \ufffd=i aij and lij = \u2212aij , where i \ufffd= j. Accordingly, if G is connected and undirected, then (L \u2297 Ip)x = 0 or xT (L \u2297 Ip) = 0 if and only if xi = xj , where xi \u2208 R p, x = [x1; . . . ;xn]. Note that xT (L \u2297 Ip)x= 1 2 n \u2211 i=1 n \u2211 j=1 aij\ufffdxi \u2212 xj\ufffd 2 . Consider a quadrotor-manipulator system consisting of a quadrotor with a p-DOF manipulator as shown in Fig. 1. Let \u2211 w denotes the world-fixed inertial reference frame, \u2211 b denotes the quadrotor body-fixed reference frame. pq = [x; y; z] \u2208 R 3 is the position of the mass center of the quadrotor in \u2211 w. With the roll/pitch/yaw angles of the quadrotor represented by \u03c6 = [\u03d5; \u03b8;\u03c8] \u2208 R 3, the rotation matrix Rb from \u2211 b to \u2211 w is given as Rb = [ c\u03b8c\u03c8 \u2212 s\u03d5s\u03b8s\u03c8 \u2212c\u03d5s\u03c8 s\u03b8c\u03c8 + s\u03d5c\u03b8s\u03c8 c\u03b8s\u03c8 + s\u03d5s\u03b8c\u03c8 c\u03d5c\u03c8 s\u03b8s\u03c8 \u2212 s\u03d5c\u03b8c\u03c8 \u2212c\u03d5s\u03b8 s\u03d5 c\u03d5c\u03b8 ] , (1) where c\u03b3 and s\u03b3 denote, respectively, cos \u03b3 and sin \u03b3. Moreover, let p\u0307q denotes linear velocity of the quadrotor and \u03c9 = [\u03c9x;\u03c9y;\u03c9z] \u2208 R 3 denotes angular velocity of the quadrotor which can be represented by \u03c9 = Q\u03c6\u0307" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002880_s00170-018-2687-1-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002880_s00170-018-2687-1-Figure2-1.png", + "caption": "Fig. 2 The geometric model of the worktable assembly of the dry gear hobbing machine", + "texts": [ + " The aim of this study is to provide refined numerical model of the worktable assembly in the dry hobbing machine to discover the heat transfer process of the worktable assembly in the cutting process, as well as both the static and transient thermal-induced deformation characteristics of the worktable assembly. This study is based on the worktable assembly of a YDE3120CNC dry hobbing machine shown in Fig. 1. This worktable assembly mainly consists of the shell, worktop, shield, spindle, fixture, bearings, and servo motor. The typical process parameters of this gear hobbing machine and the geometric parameters of the coupled hob are listed in Table 1. A geometric model of the worktable assembly with a simplified structure is shown in Fig. 2. For a dry gear hobbing machine in the cutting process, there are three heat sources, i. e., the friction heat of bearing, the heat generated from the servo motor, and the cutting heat. To cool the motor, the dry hobbing machine is equipped with a water chiller whose cooling capacity is significantly higher than the heat generated from the servo motor. Hence, the motor heating was ignored in the modeling, instead, considering that the water outlet temperature of the chiller is 25 \u00b0C, the outer surface temperature of the spindle part of the motor was defined as 27 \u00b0C", + " 6b, the largest y-axis-direction deformation appeared at the workpiece, which was 43.09 \u03bcm. The reason is that the workpiece was fixed by the fixture and the center above it, which limited the displacement in the axial direction. Therefore, the thermal-induced deformation has to happen at other directions. To further comprehend how the temperature and deformation fields vary with time, the transient-state thermal-structure characteristics of the worktable assembly were analyzed as well. Eight points from the surface of the worktable assembly were selected and monitored. Figure 2 and Table 6 show the locations of these eight points. As shown in Fig. 7, temperature of these eight points varies with time and location: the closer to the workpiece, the faster the temperature rises. At the initial stage of large-volume gear hobbing, the temperature of the worktable assembly kept increasing and subsequently leveled off. It is because that the compressed air cools the worktable assembly by the convection heat transfer. The heat transfer rate mainly depends on the difference between the air temperature and surface temperature of each component" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003852_j.mechmachtheory.2019.103628-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003852_j.mechmachtheory.2019.103628-Figure7-1.png", + "caption": "Fig. 7. The motion schematic diagram of a typical CNC hypoid grinding machine.", + "texts": [ + " The position vector of grinding wheel top center in gear coordinate system g can be expressed as follows: V g w = V g r + r w \u2217 M ( g 2 , \u03bc) n g r (38) The axis vector of grinding wheel k w in gear coordinate system g can be expressed as follows: k g w = M ( g 2 , \u03bc) g 1 (39) Another coordinate system \u2032 g = { O g , i \u2032 g , j \u2032 g , k \u2032 g } is established to rigidly connect with gear at the cross point of gear pair, in which i \u2032 g is in the same direction as the axis vector of the gear V g , and j \u2032 g is in the opposite direction with j g . The motion schematic diagram of a typical CNC hypoid grinding machine is shown in Fig. 7 , which is a five-axis machine tool. The grinding wheel is located on the left side of the machine tool and is driven up and down by Y-axis and back and forth by Z-axis. The gear is located on the right side of the machine tool. A-axis is located on the B-axis, and the gear rotation is driven by the A axis, which can swing around the Y-axis. The B axis is located on the X-axis, which can be driven by it to move left and right longitudinally. Let machine tool coordinate system m = { O m , i m , j m , k m } have the same axis direction as \u2032 g , and the coordinate origin is located at O m which is located on the A-axis (workpiece spindle), and its distance to the end of shaft A is R 0 ", + " The axis vector of grinding wheel is transformed into machine tool coordinate system m : k m w = M m k g w (40) M m = \u23a1 \u23a3 \u2212cos \u03b3 0 sin \u03b3 0 \u22121 0 sin \u03b3 0 cos \u03b3 \u23a4 \u23a6 (41) The position vector of grinding wheel top center is also transformed into machine tool coordinate system m : V m w = M m V g w + ( R 0 + L B + L M ) i m (42) where L B and L M are arbor length and mounting distance of gear. Since the grinding wheel axis is in the direction of k m and cannot be adjusted, the angle of the grinding wheel axis relative to the gear axis needs to be obtained by adjusting the gear axis. As shown in Fig. 7 , the grinding wheel axis can be transformed to the direction of k m by two steps. First, the original grinding wheel axis is rotated angle \u03c8 1 around the coordinate axis i m of machine tool, so that it can be located in the i m O m k m plane. Then, it is rotated by angle \u03c8 2 around the coordinate axis j m of machine tool. k m = M ( j m , \u03c8 2 ) M ( i m , \u03c8 1 ) k m w (43) \u03c8 1 = ta n \u22121 ( k m w \u00b7 j m / k m w \u00b7 k m ) (44) \u03c8 2 = \u2212ta n \u22121 ( k m w \u00b7 k m / k m w \u00b7 i m ) (45) In order to keep the relative position of the gear and grinding wheel, both the gear axis vector and the position vector of the grinding wheel top center should be transformed at the same time", + " The final parameters of the grinding wheel are used to check the interference of the non-working side of the grinding wheel again as shown in Fig. 15 . It can be seen that there is no interference, and the minimum gap for concave side grinding is 0.449 mm, the minimum gap for convex side grinding is 0.217 mm. Both locate at the top of the grinding wheel. The distance increases gradually from the top of grinding wheel to the top of tooth. According to the motion principle of CNC hypoid grinding machine shown in Fig. 7 , a five-axis tooth grinding simulation platform for the FFHHG is constructed, as shown in Fig. 16 . The distance R 0 is 350 mm, L B and L M are 102.538 mm and 65 mm, respectively. Then the grinding wheel location of tooth grinding can be calculated according to the method described in Section 5 , which are shown in Figs. 17 and 18 . At the same time, the grinding wheel model (the green part) is built according to Table 4 , and the imported gear model (the orange part) is shown in Fig. 10. There is no feeding during the first grinding, and the grinding wheel should be tangent to the tooth surface theoretically" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001135_1077546309353917-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001135_1077546309353917-Figure2-1.png", + "caption": "Figure 2. A two-node beam element with a rigid longitudinal motion.", + "texts": [ + " With the longitudinal elastic motions neglected, the longitudinal loads may be derived from the rigid body inertia forces, and can be expressed as P \u00bc PR Aaox l x\u00f0 \u00de \u00fe 1 2 A _ 2 l 2 x2 \u00f06\u00de where PR is an external longitudinal load acting at the right hand end of an element, and aox is the absolute acceleration of the point O in the x direction shown in Figure 1. In view of the high axial stiffness of a beam, it is reasonable to consider the beam as being rigid in its longitudinal direction. A two-node beam element having a rigid longitudinal motion is shown in Figure 2, where u1 is the longitudinal displacement; vi (i\u00bc 1, 2) are the transverse displacements; i (i\u00bc 1, 2) are the rotations; mi (i\u00bc 1, 2) are the curvatures. Because the DFE method is to approximate the displacements as polynomials, the longitudinal and transverse deflections are given as Longitudinal deflection: u \u00bc u1 \u00f07\u00de Transverse deflection: v \u00bc \u00bdN f eg \u00f08\u00de at The University of Iowa Libraries on June 21, 2015jvc.sagepub.comDownloaded from where u1 is a nodal variable, which is constant with respect to the x direction; [N] is the matrix of shape functions; { e} is the vector of nodal variables, which depends on an assumed polynomial to approximate the transverse deflection", + " To simply demonstrate the derivation of shape functions, the mechanism is considered as each link discretized as two equal two-node elements, and the curvature distribution of each element is approximated as a linear function, i.e., \u00f0i\u00deN\u00f0C\u00de \u00bc 1 x l , x l h i \u00f022\u00de ef g \u00bc m1 m2 T \u00f023\u00de where m1 and m2 represent the curvature at the end nodes for a beam element. The input link can be treated as a cantilever, i.e., there is no transverse deflection and the rotation at point O (Figure 3), which can be utilized to determine the integration constants. The coupler can be treated as a pinned-pinned beam, but there are transverse deflections X2 and X3 (Figure 3) appearing at both pins. Based on Figure 2, the occurrence of the transverse deflection X2 and X3 is due to the transverse deflection of point A for the input link (see equations (14)\u2013(15)). Similar to the coupler, the output link can be treated as a pinned-pinned beam, but there is a transverse deflection X4 appearing at point B. Also, there is no bending stress at points A, B and D, i.e., the curvatures are zero at these locations, which leads the number of nodal variables for the mechanism as four. Therefore, the shape functions for each link can be summarized as \u00f01\u00deN \u00bc x2 2 x3 6l2 , x3 6l2 , 0, 0 \u00f024\u00de \u00f02\u00deN \u00bc l2x 2 \u00fe l22 3 , l22 6 \u00fe l2x 2 \u00fe x2 2 x3 6l2 , 0, 0 \u00f025\u00de \u00f03\u00deN\u00f0D\u00de \u00bc 5l22 6 S2\u00fe S3 S2\u00f0 \u00dex R3 , l22 2 S2\u00fe \u00f0S3 S2\u00dex R3 , l3x 2x \u00fe x3 6l3 , 0 2 664 3 775 \u00f026\u00de \u00f04\u00deN\u00f0D\u00de \u00bc 5l22 6 S2 \u00fe S3 S2\u00f0 \u00de l3 \u00fe x\u00f0 \u00de R3 , l22 2 S2 \u00fe S3 S2\u00f0 \u00de l3 \u00fe x\u00f0 \u00de R3 , l23 3 \u00fe x2 2 x3 6l3 , 0 2 6666664 3 7777775 \u00f027\u00de at The University of Iowa Libraries on June 21, 2015jvc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001264_vppc.2012.6422500-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001264_vppc.2012.6422500-Figure3-1.png", + "caption": "Fig. 3: Exploded view of 3-D structural model of the motor", + "texts": [ + " The list of natural frequencies and modes of a motor were used as an input to an acoustic boundary element solver. It has been known that modal shapes of low circumferential mode order numbers have significant effect on the airborne noise [1]. Therefore, it is important to avoid running a motor near such frequencies that the stator-frame vibrates with the low order numbers. Since electric vehicles move over a range of speeds, it is impossible not to run a traction motor at its resonant frequencies. However, the resonance does not occur by quickly passing over dangerous rotational speeds. Fig. 3 shows an example of a three dimensional model of a motor used in the modal analysis. Resonant frequencies and modal shapes were calculated by ANSYS and some examples of 2-D and 3-D modal shapes with 2nd circumferential mode are presented in Fig. 4. Since natural frequencies and vibration behavior of a motor are significantly dependent on the way the motor is fixed and supported, boundary conditions were appropriately applied in order to capture the actual fixture circumstance. Radial magnetic forces with both fundamental and high order harmonics and natural frequencies with modal shapes were incorporated into an acoustic boundary element solver, LMS Virtual" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002871_s10118-019-2180-9-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002871_s10118-019-2180-9-Figure2-1.png", + "caption": "Fig. 2\u00a0\u00a0\u00a0\u00a0The\u00a0experimental\u00a0setup:\u00a0(a)\u00a0A\u00a0scheme\u00a0of\u00a0DIC\u00a0experimental\u00a0arrangement;\u00a0(b) A\u00a0specimen\u00a0with\u00a0the\u00a0random\u00a0and\u00a0high\u00a0contrast\u00a0pattern\u00a0on\u00a0its\u00a0surface", + "texts": [ + " \u00a0In \u00a0order \u00a0to \u00a0measure \u00a0the \u00a0exact elongation\u00a0of\u00a0polyamide\u00a0specimens\u00a0(PA66;\u00a0trade\u00a0name:\u00a0Zytel 101), \u00a0the \u00a0long-range \u00a0extensometer \u00a0with \u00a0the \u00a0total \u00a0measurement\u00a0range\u00a0of\u00a0about\u00a01.2\u00a0m\u00a0was\u00a0employed. For\u00a0the\u00a0purpose\u00a0of\u00a0estimating\u00a0the\u00a0material\u00a0compressibility, a\u00a0non-contact,\u00a0optical\u00a0method\u00a0based\u00a0on\u00a0digital\u00a0images\u00a0correlation\u00a0(DIC)\u00a0was\u00a0utilized.\u00a0The\u00a0DIC\u00a0method\u00a0allows\u00a0one\u00a0to\u00a0obtain\u00a0the \u00a0maps \u00a0of \u00a0displacements \u00a0and \u00a0macroscopic \u00a0 deformations\u00a0measured\u00a0on\u00a0a\u00a0surface\u00a0of\u00a0the\u00a0tested\u00a0specimen.\u00a0The\u00a0DIC experimental\u00a0setup\u00a0consists\u00a0of\u00a0a\u00a0black\u00a0and\u00a0white\u00a0digital\u00a0camera \u00a0with \u00a0a \u00a0suitable \u00a0lens, \u00a0and \u00a0a \u00a0PC \u00a0with \u00a0installed \u00a0software (Fig.\u00a02a). \u00a0 https://doi.org/10.1007/s10118-019-2180-9 \u00a0 The \u00a0presented \u00a0research \u00a0used \u00a0the \u00a0commercial \u00a0program Vic2D.\u00a0Before\u00a0the\u00a0test,\u00a0random\u00a0spots\u00a0with\u00a0high\u00a0contrast\u00a0must be \u00a0applied \u00a0to \u00a0the \u00a0surface \u00a0(Fig. \u00a02b). \u00a0After \u00a0selecting \u00a0the \u00a0field for \u00a0analysis, \u00a0the \u00a0specimen\u2019s \u00a0surface \u00a0is \u00a0divided \u00a0into \u00a0smaller subareas \u00a0characterized \u00a0by \u00a0a \u00a0unique \u00a0distribution \u00a0of \u00a0gray shades.\u00a0A \u00a0computer \u00a0algorithm \u00a0compares \u00a0the \u00a0digital \u00a0 photographs\u00a0of\u00a0a\u00a0non-deformed\u00a0sample\u00a0with\u00a0the\u00a0photos\u00a0taken\u00a0after the\u00a0deformation\u00a0and\u00a0finds\u00a0the\u00a0new\u00a0positions\u00a0of\u00a0analyzed\u00a0subareas\u00a0by \u00a0utilizing \u00a0the \u00a0method \u00a0of \u00a0best \u00a0fit. \u00a0As \u00a0a \u00a0result, \u00a0a \u00a0displacement\u00a0map\u00a0or\u00a0macroscopic\u00a0deformations\u00a0are\u00a0obtained\u00a0for the\u00a0entire\u00a0analyzed\u00a0area" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001811_roman.2017.8172348-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001811_roman.2017.8172348-Figure1-1.png", + "caption": "Fig. 1. Illustrative representation of a joint mechanism with backlash between motor and link.", + "texts": [ + " V, we present experimental results using the prototype of the left arm of Romeo robot from SoftBank Robotics, showing: i) the benefits with respect to the classical residual; ii) the effectiveness of contact detection both with the robot at a given static posture or in motion; and iii) the use of our residual extension to allow physical robot manipulation and collaboration. Joint backlash is an undesired effect of joint mechanisms; it is usually due to a play in the transmission between the motor shaft and the link shaft, as illustrated in Fig. 1. The effect of this play is that during motion the torque generated by the motor \u03c4M is not always transferred to the link. Hence, the toque acting on the link side \u03c4L is not equal to the motor torque. It is straightforward to figure out that the backlash produces an error in the classical residual. This error is negligible when the backlash is very small, as commonly found in industrial robots; but it becomes significant with large joint backlash. To illustrate this effect, we simulated a single joint with a non negligible backlash controlled so as to generate a sinusoidal joint positioning" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003948_iecon.2019.8927305-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003948_iecon.2019.8927305-Figure7-1.png", + "caption": "Fig. 7. Axes of the selected teeth for harmonic response investigation.", + "texts": [ + " Utilizing the full model allows investigation of harmonic response at 2L where is the strongest excitation. Since resulting natural frequencies are usually far from 2L, harmonic response solved on a reduced model would not be investigated at 2L in that case. Harmonic response is detected in the nodes of finite element mesh located on the outer circumference of the stator lamination (Case A) or the motor frame respectively (Cases B and C). Node position is always selected as close as possible to the respective tooth axis in accordance with Fig. 7. Horizontal and vertical axes correspond to x and y axes of Cartesian coordinate system. Response in two diagonal axes is investigated considering resulting structural modes and the machine pole number. Quantity used for harmonic response representation is acceleration therefore higher frequencies are highlighted in the presented vibration spectra. Acceleration is chosen due to better comparison with measurement carried out using accelerometers. Frequency spectra of radial and tangential forces acting on a single tooth and harmonic response detected in the corresponding mesh node are depicted in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002795_gt2018-77151-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002795_gt2018-77151-Figure2-1.png", + "caption": "Fig. 2 Pad FEM structure model and constraint condition [21].", + "texts": [ + " Incidentally, there is a notable thermal convection between the pad surfaces and the oil oil churning in a feed groove [28]. Hence, the heat transfer coefficients for both the pad leading edge and trailing edge surface facing an oil feed groove are relatively high, hgr=1,750 W/m 2 K as cited in Ref. [28]. A constant temperature equal to 40 \u00b0C is set as the oil supply temperature in all pad leading and trailing edge surfaces facing an oil feed groove. An ambient temperature at 50 \u00b0C is assumed for any other surface i.e., back and side surfaces of a pad. Fig. 2 displays a mesh for a pad finite element (FE) model and the applied constrain conditions. Table 1 sums the pad FE model using eight-node brick elements. Radial displacements are constrained along the pivot line and two parallel side lines, as defined in Ref. [13]. In addition, there is a fully fixed point at the center of the pivot line. Only elastic displacements of the pad surface in contact with the film are needed. Hence, a Guyan reduction or static condensation technique is appropriate for an efficient computation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000451_978-3-642-29329-0_6-Figure6.14-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000451_978-3-642-29329-0_6-Figure6.14-1.png", + "caption": "Fig. 6.14 Bifurcation diagram of the normal form (6.6) in the case s \u00bc 1", + "texts": [ + "8), we can obtain the asymptotic form of the homoclinic bifurcation curve in the original parameter space, that is HOMe2 \u00bc \u00f0p1; p2\u00de : p2 \u00bc 1 7 p1; p1 0 : In order to find the asymptotic expansion of the other homoclinic orbit, we first translate our system in order to move the equilibrium e2 (non-saddle) to the origin of the state space, by the translation 1 \u00bc z1 p1; 2 \u00bc z2: Then, through the following change of coordinates and time rescaling 1 \u00bc z1p1; 2 \u00bc z2p 3=2 1 ; t \u00bc t ffiffiffiffiffi p1 p ; and the parameter transformation g1 \u00bc p1 p2ffiffiffiffiffi p1 p ; g2 \u00bc ffiffiffiffi p1 p ; (6.9) we reduce again our system to (6.7). At this point, by repeating the above analysis, we obtain the asymptotic formula of the second homoclinic curve which, in the original parameters, reads HOMe1 \u00bc \u00f0p1; p2\u00de : p2 \u00bc 6 7 p1; p1 0 : We are now in the condition to sketch the bifurcation diagram in the neighborhood of the codimension-two point, as reported in Fig. 6.14. The bifurcation found in Sect. 6.3.1.3 corresponds to passing from the region between the curves(HOMe2, TC) to the region between the curves(He2 ,HOMe1 ). Thanks to the fact that all the bifurcation curves are asymptotically linear, with different slopes, the above described passage is generically possible, provided the original parameters are a suitable combination of \u00f0p1; p2\u00de. 1. Andrzejewski, R., Awrejcewicz, J.: Nonlinear Dynamics of a Wheeled Vehicle. Springer (2005) 2. Arctic Falls Proving Grounds: http://www" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003896_j.triboint.2019.106096-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003896_j.triboint.2019.106096-Figure2-1.png", + "caption": "Fig. 2. Two-disc test rig photograph and technical drawing of its test chamber used for the recreation of thermally induced WEL in oil-lubricated contacts using bearing inner rings made of 100CrMn6.", + "texts": [ + " Since this is undesired, the experiments with the round focus are carried out in pulse mode. In this mode the laser beam also acts on the specimen for a defined time. However, not continuously while the machine table is moving but intermittently (Fig. 1b). In the used configuration the shutter of the system achieved reproducibly a minimum pulse duration of 3 ms. The comparability of the experiments in pulse mode with the experiments in feed mode is confirmed in random tests. In this study, an in-house developed two-disc test rig (Fig. 2) is used to perform WEL accelerations tests under defined tribological conditions [29]. The functional principle corresponds to that of a disc-disc-tribometer. Both discs are driven by separate servo motors, the lower disc operates in speed control and the upper disc operates in torque control. The upper shaft is pre-loaded with a braking torque which is opposite to the torque resulting from the shear stress in the rolling/sliding contact. To avoid an uncontrolled spin of the upper shaft, an overrunning clutch blocks the shaft in the direction of rotation against the braking torque" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001540_j.1747-1567.2011.00755.x-Figure11-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001540_j.1747-1567.2011.00755.x-Figure11-1.png", + "caption": "Figure 11 Finite element model of the drop tester.", + "texts": [ + " Numerous factors that influence impact pulses exist, such as different drop heights and strike surfaces. If the rigidity relationship between the strike surface and the impact pulse can be clarified, this could be considered as a guideline for adjusting the strike surface in obtaining the standard JEDEC conditions or other in-house design conditions. The entire finite element model, which included a rigid base, strike surface, and drop table, was created according to actual dimensions as shown in Fig. 11. Figure 12 shows the boundary conditions and constraints. To reduce computational time, the distance of the drop table to the strike surface was only 1 mm, and nearimpact velocity was applied as an initial velocity according to the curve fitting equation in Fig. 6. Three factors, drop height, strike surface thickness, and strike surface material, were considered during the drop simulation. The benchmark test condition, strike surface thickness, and material properties are summarized in Table 2. A one-way analysis of variance (ANOVA) was adopted in performing the factors effect, assuming that one factor varied and the others remained as the benchmark" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002311_j.ifacol.2017.08.973-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002311_j.ifacol.2017.08.973-Figure2-1.png", + "caption": "Fig. 2. Twin Rotor MIMO System", + "texts": [ + " (13) There exist \u03b1i, \u03b2i, \u03c3i and Di(p) such that control system (6) with observer (10) provides convergence of tracking error to the bounded area with accuracy \u03b4 = \u221a 1 \u03bbmin(P1) [( V (0)\u2212 \u03b8 \u03c2 ) e\u2212\u03c2T + \u03b8 \u03c2 ] , (14) where V = \u03b5TP1\u03b5+\u03b7TP2\u03b7 is a Lyapunov function, P1 and P2 are solutions of equations A TP1+P1A = \u2212Q1, H TP2+ P2H = \u2212Q2 respectively, Q1 and Q2 are symmetric positive defined matrices, \u03b8 = sup\u03d52(t) \u03c5 + \u22062 \u03c5 , \u03c5 is a some positive number, \u03c2 is a positive number depending on plant parameters. A series of experiments were conducted for verification of obtained results. Mechatronic laboratory bench \u201dTwin Rotor MIMO System\u201d (TRMS) (Fig. 2) was used for experimental research. TRMS is a laboratory helicopterlike system with two degrees of freedom and opportunity of independent two-channel control. The system includes two DC motors: one of them provides movement in vertical plane (pitch angle) and the other one is for motion in horizontal plane (yaw angle). TRMS is controlled by independent voltage levels on the armature of the motors (in the range [-2.5V ... 2.5V]). The pitch and yaw angles are measured by optical encoders. Consider mathematical model of laboratory bench" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001547_raad.2010.5524611-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001547_raad.2010.5524611-Figure3-1.png", + "caption": "Fig. 3. Inertia matrix is approximated as a diagonal one based on the symmetric shape of the helicopter. The inertia around one axis can be determined by the following formula:", + "texts": [], + "surrounding_texts": [ + "Based on the Newtonian law the linear dynamic model for the orientation can be described as x bl \u0398 \u03a9\u2212\u03a9= )( 2 2 2 4\u03d5 (5) y bl \u0398 \u03a9\u2212\u03a9= )( 2 1 2 3\u03d1 (6) z d \u0398 \u03a9\u2212\u03a9+\u03a9\u2212\u03a9 = )( 2 3 2 4 2 1 2 2\u03c8 (7) where \u0398 is the inertia around the axes of the coordinate system. The equations are valid in the frame of the helicopter. These equations remain acceptable in the world frame, if the orientation of the helicopter remains horizontal. Therefore these can be considered as a linearized model around 0=\u03b7 , where \u03b7 is the 3D Euler(RPY) orientation of the helicopter relative to world frame. In this dynamic model, the effect of the aerodynamic friction and the gyroscopic effect considered to be negligible. The aerodynamic friction is in linear connection with \u03b7 which is approximated with zero. The gyroscopic effect is a function of the rotor inertia r\u0398 , this can be ignored, because the mass of each rotor is 0.4% of the full mass of the helicopter and this value is also in connection with \u03b7 ." + ] + }, + { + "image_filename": "designv11_33_0000009_cae.20327-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000009_cae.20327-Figure3-1.png", + "caption": "Figure 3 Movement of the oscillating group.", + "texts": [ + " There is also the option of representing the variables as a function of any variable, so the displacement of the center of gravity of the drum in the front plane z y can be obtained, which is known as the signature of the movement. Finally, the third module is the post-processor that enables the student to see a 2D or 3D motion movie of the oscillating group and analyze if its movement is right, or the oscillating group strikes the cabinet. It is an application programmed with AutoLISP, which through the main window is fed by the drawing and position files generated by the geometric and dynamic modules, respectively. This application creates the frames, Figure 3, that are afterwards linked by means of Autodesk Animator Pro to produce the movie that simulates the 2D or 3D movement of the oscillating group. This module is a perfect complement of the dynamic module because it makes the comprehension of the results easier, and the user can check whether the design is acceptable. The design task conducted by the students is a very challenging, real-life problem: the students are asked to design an oscillating group for a washing machine. It is delivered as a laboratory activity of the first-year subject Dynamics offered within the Industrial Engineering degree of the University of Zaragoza (Spain)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000980_icra.2012.6225322-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000980_icra.2012.6225322-Figure3-1.png", + "caption": "Fig. 3. Left: Given view points fixed to an object. Positions of these view points are centers of faces of a geodesic dome represented by icosahedron. Right: Grasp position example. Grasp positions are expressed in object coordinates. When re-grasping, we solve inverse kinematics of which the input is the transformation of two grasp positions.", + "texts": [ + " executed simulation of the re-grasping task of known shape objects [9]. In this section, we show the outline of our re-grasping observation system. Fig.2 shows the system of construction of 3D shape model with observation based on dual-arm regrasping, and Algorithm.1 describes the detail of this system flow. Robots observe an object from the view points fixed to the objects coordinates. These view points are centers of faces of a geodesic dome represented by regular icosahedron (the left figure of Fig.3 ). We solve inverse kinematics that robots can Algorithm 1 Algorithm of observation based on dual-arm re-grasping 1: i\u2190 0 : the number of re-grasping 2: V : given view points 3: P : all of observed 3D point cloud 4: Pi : all of observed 3D point cloud in i-th grasp 5: Pi,j : observed 3D point cloud from a view point vj in i-th grasp 6: IsF inishObservation = false 7: while !IsF inishObservation do 8: for all vj \u2208 V do 9: if exists IK solution to observe from vj then 10: move robot 11: get 3D point cloud on object Pi,j 12: append Pi,j to Pi 13: end if 14: end for 15: Voxel Filtering to P , P i 16: if checkfinishfn(P ,Pi) == true then 17: IsF inishObservation\u2190 true 18: else 19: re-grasping 20: append Pi to P 21: i\u2190 i+ 1 22: end if 23: end while observe from each view point", + " Using the observed 3D point cloud, robots check that the observation of objects is finished. We mention the detail of the functions for check finish observation later. When the observation has not finished yet, compute grasp position and select the next grasp position. We use OpenRAVE for grasp position computation[10]. The target model for grasp computation is convex hull of 3D point cloud observed by then. If no next grasps is avaiable, robots give up observation. We use the information of grasp positions, which is shown in the right figure of Fig.3 . Grasp positions are represented in object coordinates. We show this selection method in section IV. Now that robots have selected the next grasp position, execute re-grasping. In the real world, robots may fail to grasp. So robots detect grasp failure and if failure is detected, return to III-C. It is possible to express this motion planning of a sequence of re-grasping position through the graph search problem. We call this graph dual-arm observation graph. Fig.4 shows an overview of the dual-arm observation graph" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000663_j.aca.2012.05.014-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000663_j.aca.2012.05.014-Figure1-1.png", + "caption": "Fig. 1. Schematic diagram of self constructed EC. CEM: cation exchange membrane. CER: cation exchange resin. FD: fluorescence detector.", + "texts": [ + " Further dilution of the stock solutions was performed with the eluent to obtain a series of standard solutions with concentration in the range of 0.01\u20135 mg L\u22121. The entire preparation process should be completed quickly because of the poor stability of folates in an aqueous medium and open air [14]. The stock standard solutions were stored in polypropylene bottles at \u221220 \u25e6C, and the calibration standard solutions were freshly prepared daily. Ultra pure Milli-Q water (Millipore, Molsheim, France) was used throughout the experiment. 2.2. Home-made EC Fig. 1 shows the structure of the home-made EC, which consisted of two quadrate PTFE blocks placed on top of the other (70 mm \u00d7 34 mm \u00d7 16 mm) in a symmetrical structure separated by two pieces of cation-exchange membranes (CEM, 70 mm \u00d7 34 mm \u00d7 0.5 mm for each piece, Shanghai Chem. Co., Shanghai, China). A Ru/Ti electrode (49.6 mm \u00d7 10 mm \u00d7 0.5 mm, kindly provided by Qingdao Shenghan Chromatography Tech Equipment Co., LTD., Qingdao, China) was embedded into the chamber in each block. Each electrode had two holes (0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001093_icra.2012.6225142-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001093_icra.2012.6225142-Figure5-1.png", + "caption": "Fig. 5. Overview of the Prototype Model of Convex Version", + "texts": [ + " INITIAL PROTOTYPES The initial prototypes based on the principle shown above were built. We made three types of gear mechanisms, each of which is described from the next section. In this section, we describe the configuration of the planar version of the omnidirectional driving gear mechanism. The actuators that drive the mechanism are placed in world coordinates so the plate that moves omnidirectionally on the flat plane can be configured in a lightweight manner. The overall view of the manufactured prototype of the convex version is shown in Fig. 5. The driving unit of the prototype of the concave version is shown in Fig. 6. To activate this prototype of concave version of the omnidirectional gear driving unit, two pairs of spur gears are deployed inside of the tube structure of the concave omnidirectional gear. One pair is deployed to its rotational direction, and the other is deployed to its translational direction. Only one spur gear is enough to activate this omnidirectional gear, however, to cancel any undesired outer torque, a pair of spur gears is deployed to move its one axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001264_vppc.2012.6422500-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001264_vppc.2012.6422500-Figure5-1.png", + "caption": "Fig. 5: Actual motor set to the dynamometer (left) and motor geometry for simulations (right)", + "texts": [ + " (3) where [n] is the modal matrix and {MRSP( O))} is the modal response at a given frequency, 0) [7]. As a result, the sound pressure level can be obtained by combining (2) and (3). p(OJ) = iOJ{ATV(OJ)}i[nHMRSP(OJ)} (4) III. MODEL VALIDATION To determine the effectiveness of using the model to predict the acoustic noise of electric motors, experiments and numerical simulations of a traction motor at various operation conditions were conducted. A spoke-type interior permanent magnet synchronous motor with 10 poles and 12 slots was used for the test. Fig. 5 shows the actual picture and 3-D modeled geometry for numerical simulations, respectively. Sound pressure level was measured at the point 1m from the center of the motor using B&K type 4191 free-field microphone and Audio Analyzer system. A-weight acoustic filter was applied to both measured sound pressure level and predicted noise level. Each measurement was replicated three times for given operating conditions. The motor was operated within its maximum power region and sound pressure level was measure at six levels of speeds (2500, 3000, 4000, 5000, 6000, and 7000rpm) as shown in Table I" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003466_j.mechmachtheory.2019.06.017-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003466_j.mechmachtheory.2019.06.017-Figure3-1.png", + "caption": "Fig. 3. Definition of the rack-cutter: (a) normal section;(b) pitch plane.", + "texts": [ + " The profile crowned tapered involute pinion and shaper are obtained. The limiting conditions (pointing and undercutting) of the tooth surfaces are determined to avoid singularities when generating the tapered involute pinion. The rack-cutter is used to generate both the tapered involute shaper and the tapered involute pinion, then the former is utilized to generate the face-gear, while the latter is applied to meshing with the face-gear. A rack-cutter with helix angle and parabolic profiles is shown in Fig. 3 . The tooth surfaces ri of the rack-cutter in coordinate system S r are expressed as follows: r ri ( u i , l i ) = M rb ( l i ) M ba r ai ( u i ) r ai ( u i ) = \u23a1 \u23a2 \u23a3 a i u 2 i \u2212u i 0 1 \u23a4 \u23a5 \u23a6 (1) M rb ( l i ) = \u23a1 \u23a2 \u23a3 cos \u03b2 0 sin \u03b2 l i sin \u03b2 0 1 0 0 \u2212 sin \u03b2 0 cos \u03b2 l i cos \u03b2 0 0 0 1 \u23a4 \u23a5 \u23a6 (2) M ba = \u23a1 \u23a2 \u23a3 \u00b1 cos \u03b1n \u00b1 sin \u03b1n 0 \u00b1( L d cos \u03b1n + u 0 i sin \u03b1n ) \u2212 sin \u03b1n cos \u03b1n 0 u 0 i cos \u03b1n \u2212 L d sin \u03b1n 0 0 1 0 0 0 0 1 \u23a4 \u23a5 \u23a6 (3) where u i , l i ( Fig. 3 (a) and (b)) are the surface parameters of the rack-cutter; subscript i = p, s represents the tapered involute pinion and shaper, respectively. a i means the parabola coefficient of the parabolic profile of the rack-cutter. Parameter u 0 i determines the shifting distance of the tangent point of parabolic profile to the conventional linear profile. Matrices M rb ( l i ) and M ba describe the coordinate transformation from coordinate system S a to coordinate system S r . The plus sign in matrix M ba corresponds to the coordinate transformation for left tooth surface of the rack, while the minus sign corresponds to the right tooth surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003983_j.ifacol.2019.12.310-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003983_j.ifacol.2019.12.310-Figure1-1.png", + "caption": "Fig. 1 Body-fixed reference frame (LEFT) and NED coordinate system (RIFHT)", + "texts": [ + " The control objectives and the control system are presented in Section 3 and 4 respectively. Simulation results are presented in Section 5 followed by the conclusion in Section 6. 2. VEHICLE MODEL This section describes the 6-DOF manoeuvring model for an underactuated AUV moving in 3D space and formulates the problem of 3D path following of space curves. Let [ , , ]TX x y z represent the position vector from the North- East-Down (NED) coordinate origin { , , }i i ix y z to the origin of a body-fixed frame which is centre of buoyancy, { , , }b b bx y z as illustrated in Fig. 1. The translational and rotational velocities of the vehicle are presented as [ , , ]Tu v wv and [ , , ]Tp q r\u03c9 expressed in the body frame with respect to the inertial frame. The kinematic equations are (Woolsey, 2011): \u02c6 X Rv R R (1) where \u0302 denotes the 3x3 skew-symmetric matrix satisfying a\u0302b = a\u00d7b for vector a and b. The vehicle model is based on following assumptions: Assumption 1: The motion of the AUV is described by 6 degree of freedom (DOF), which are surge, sway, heave, roll, pitch and yaw. Assumption 2: The AUV is port-starboard symmetric. Assumption 3: The vehicle\u2019s centre of mass (CM) is located at cmr , and the origin on the body reference frame is located at the centre of buoyancy (CB); see Fig. 1. components : a steady, circulating flow component Vs(X) and an unsteady and uniform flow component Vu(t) (Thomasson, 2000). These two components can be expressed in body-fixed reference frame as follows: ( , ) ( ) ( , ) ( ) T s s T u u v R X R V X v R t R V t (2) The complete flow field is ( , , ) ( , ) ( , )f s uv R X t v R X v R t . Let v the generalized vehicle\u2019s velocity while vs and vu represent the steady and circulating flow and unsteady and uniform flow respectively as follows: , 0 and 0 T T T s s u uv v v The vehicle\u2019s velocity relative to the flow can be expressed as follows: r r T s uv v v v where r s u is the vehicles velocity relative to the flow expressed in body-fixed frame" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002954_ccta.2018.8511496-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002954_ccta.2018.8511496-Figure1-1.png", + "caption": "Fig. 1. Coordinate frames", + "texts": [ + " The AUV is axially symmetric and its rudder \u03b4\u03c8 and elevator \u03b4\u03b8 commands are executed by moving a double gimbal mechanism to actuate the ring-wing and thruster. The vehicle is stable in roll axis and hence no roll input is needed. Structural flexibility of the AUV is not considered in the mathematical model, and the acceleration due to gravity and water density are assumed to be constants. The two reference frames used in deriving the AUV\u2019s six DOFs equations of motion are the North-East-Depth earthfixed frame FE and the body-fixed frame FB , see Fig. 1. The displacement vector from OE to OB is expressed in FE as ro = [ xe ye ze ]T E (1) and the displacement vectors from OB to the vehicle\u2019s center of gravity cg and the center of buoyancy cb are expressed in FB as rcg = [ xg yg zg ]T B , rcb = [ xb yb zb ]T B (2) The attitude roll, pitch, and yaw angles are denoted by \u03c6, \u03b8, and \u03c8, respectively. The translational surge, sway, and heave velocities of the cg are denoted in FB by u, v, and w, and the roll, pitch, and yaw angular rates of FB relative to FE are denoted in FB by p, q, and r, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002779_1464419318789185-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002779_1464419318789185-Figure2-1.png", + "caption": "Figure 2. The relationship between the target tooth and the rack cutter.", + "texts": [ + ",23 it is important to deal with micro-geometrical form and their effects upon gear vibration in practice. Actually, all gear tips are modified or relieved in order to reduce sudden impacts at tip\u2013root or tip\u2013tip interactions. However, it should be noted that this paper ideally deals with the involute geometry to simplify the model, which is an acceptable assumption. It is also assumed that the target tooth is produced by the generating method using a rack cutter. The relationship between the tooth and the cutter is shown in Figure 2, in which the global coordinate system is fixed on the center of the pitch circle and the local coordinate system is fixed on the rack. The coordinates of the cutting points on the rack are given as yc \u00bc 0:25 mn= cos L\u00fe cm mn tan t\u00f0 \u00de = tan t \u00fe 1= tan t\u00f0 \u00de xc \u00bc yc cm mn\u00f0 \u00de tan t 0:25 mn= cos 8>< >: \u00f01\u00de and the coordinates of the involute profile nodes on the target tooth (Figure 3(a)) are determined as xt \u00bc L\u00fe xc\u00f0 \u00de cos L=r\u00f0 \u00de r\u00fe yc\u00f0 \u00de sin L=r\u00f0 \u00de yt \u00bc L\u00fe xc\u00f0 \u00de sin L=r\u00f0 \u00de \u00fe r\u00fe yc\u00f0 \u00de cos L=r\u00f0 \u00de \u00f02\u00de Moreover, the coordinates of the face nodes on the target tooth (Figure 3(a)) are obtained as x \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi x2t \u00fe y2t p sin \u2019 i\u00f0 \u00de y \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi x2t \u00fe y2t p cos \u2019 i\u00f0 \u00de ( \u00f03\u00de where mn is the normal module, is the helical angle, cm is the modification coefficient, t is the transverse pressure angle, r is the radius of the reference circle, and L is the lateral displacement of the local coordinate system on the rack calculated as L \u00bc Lh i 1\u00f0 \u00de \u00fe L2 \u00f04\u00de Here, Lh \u00bc L1 L2\u00f0 \u00de=ne is the searching step size to determine the face nodes, in which L1 is the lateral displacement of the rack at the start point of the cutting process, L2 is the rack displacement at the end point of the cutting process, ne is the number of the nodes along the involute tooth profile, i is the identifier ranging from 1 to ne, and \u2019 is the angle between the center line (as shown in Figure 2) and the line that goes through the involute tooth profile node and the center of the global coordinate system. Combining equations (1) to (4), the involute profile nodes and the face nodes of the target tooth are accurately obtained as shown in Figure 3(a). Because the functions of the transition curve, the addendum circle, and the body curve could be conveniently obtained, the rest nodes on the (front) end face are sequentially determined based on the basic geometric parameters. Finally, all the nodes of the tooth mesh model are generated through axially extruding and simultaneously rotating all the nodes on the (front) end face" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002933_icelmach.2018.8507125-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002933_icelmach.2018.8507125-Figure7-1.png", + "caption": "Fig. 7. Stator system multi-physics model", + "texts": [ + " Some low-order subharmonics have also been suppressed. Radial force density and its harmonic contents, to some extent, can reflect the electromagnetic vibration for the machine, but this is an indirect analysis method. A direct approach to calculate the vibration is to apply the radial force on tooth tips and calculate the deformation and acceleration of the stator mechanical system. A multi-physics model in structural field is built for the stator system, including stator core, concentrated windings and frame, as is pictured in Fig. 7. The contribution of end caps and flanges is considered through constraints of bolts in the frame. With the help of 3-D FEA, we can first obtain the inherent properties of the system through modal analysis, including the mode shapes and natural frequencies. We need to focus on low-order mode shapes, including the zeroth (also called breathing mode), second, third and fourth circumferential mode, because their natural frequencies are low and generally concerned with significant vibration. The natural frequencies are 1160" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001090_20100901-3-it-2016.00149-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001090_20100901-3-it-2016.00149-Figure2-1.png", + "caption": "Fig. 2. Bifurcation set in plane (k, a), for b = 0, w = 5, yr = 2, showing the straight line (red line) k = ayr and the subcritical Hsub curve (blue line).", + "texts": [ + " Note that at k = ayr there appears a transcritical bifurcation where the two points (ayr 2, yr) and (xd, yd) collide. As a result of this bifurcation, the pseudo-equilibrium point p becomes a singular sliding point located at the corner of \u03a3as to enter this region. For the planar differential system (17) however, this bifurcation is standard. For ayr < k the determinant D(ay2 r , yr) is positive. Assuming ayr < k, the stability also requires k > kc = (awyr 2 \u2212 1)/(2ayr), which corresponds with the condition T (x, yr) < 0. In short, the stability conditions are ayr < k, and ayr(wyr \u2212 2k) < 1. In Fig. 2, the corresponding bifurcation set in the parameter plane (k, a) is depicted, showing the straight line k = ayr (labelled T, and drawn in red) that corresponds to D(ay2 r , yr) = 0, and the Hsub curve corresponding to T (ay2 r , yr) = 0. At the points of the Hsub curve, it can be rigorously shown (like in Ponce and Pagano [2009b]) that a subcritical Hopf bifurcation takes place. Thus, at the right of this curve, we have the (local) stability operating region, i.e. a stable pseudo-equilibrium point surrounded by an unstable limit cycle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002035_s12206-017-0515-4-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002035_s12206-017-0515-4-Figure9-1.png", + "caption": "Fig. 9. 3D-FEM model of the analytical models for acoustic analysis.", + "texts": [ + " 7, with a slightly small excitation force of 12480 Hz, which is close to the resonance frequency of the 5th mode in Fig. 6(e). Consequently, the deformations of the SPM_CMG and the FC_CMG are generated at 12480 Hz. At this frequency, the magnitude of the generated deformation of the FC_CMG is 1.37 times higher than that of the SPM_CMG. On the basis of the result of the vibration analysis on the outside surface of the LSR, the acoustic noise is calculated by the 3D-FEM. The vibration velocities at each frequency for the harmonics of the force are imported to each node on the outside surface of the LSR. Fig. 9 shows the acoustic FEM models for the SPM_CMG and FC_CMG. Fig. 10 shows the Sound pressure level (SPL), which is a logarithmic measure of the sound pressure relative to a reference value of 20 \u03bcPa on point A, as shown in Fig. 9. In the FC_CMG, the highest noise of 63.78 dBA occurs at 12480 Hz, which is close to the resonance frequency. This value is 7.62 dBA higher than the highest noise 56.16 dBA in the SPM_CMG. Moreover, the total SPL values of the SPM_CMG and FC_CMG are 56.9 and 64.39 dBA, respectively. Thus, the total SPL value of the FC_CMG is 7.49 dBA higher than that of the SPM_CMG. This paper compares the SPM_CMG with the FC_CMG in terms of their vibration and noise characteristics. In the FC_CMG, the radial flux density\u2019s 8th, 18th, 26th and 44th harmonics, which are absent in the SPM_CMG are additionally generated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003469_j.engfailanal.2019.06.024-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003469_j.engfailanal.2019.06.024-Figure3-1.png", + "caption": "Fig. 3. Tool performance test.", + "texts": [ + " The eccentric mechanism consists of two eccentric rings, and the main shaft power is transmitted to the electromagnetic clutch through the couplings. When the electromagnetic clutch is electrified, it meshes with the deceleration device, and the power is transmitted to the eccentric mechanism after the deceleration, thus realizing the rotation of the inner and outer eccentric rings [5]. In order to test the performance of the wellbore trajectory control tool, our team developed several experiments on the dynamic performance test bench of downhole tools, as shown in Fig. 3. The cantilever bearing of borehole trajectory control devices failure twice during the experiments, as shown in Fig. 4. After disassembling the faulty bearing, we found that both the rolling body and the cage of the bearing were damaged. It can be seen that the cages are badly worn and deformed, and it has a certain degree of fracture, at the same time, there are obvious deformation caused by extrusion damage at one end of the cage. The other end of the cage has no obvious wear marks, and the wear and extrusion condition are obviously better than the first end", + " 13, and the calculated errors under different grids on the basis of minimum grid stress are shown in Table 2. Considering the calculation time and error value, the element size of cage is selected to 0.5mm, the element size of needle is selected to 0.3mm in subsequent analysis, 6. Experiments and result analysis 6.1. Experimental verification In order to verify the theoretical and FE results, the laboratory simulation experiments are carried out on the dynamic performance test bench of downhole tools (as shown in Fig. 3), the specific parameters areTable 1. Since the experiments are based on existing sensors of the tool, the verification should depend on the deflection angle of the drill bit, which is considered to the deflection angle of the lower end of the main shaft (x=0) [30]. Taking maximum offset distance of the eccentric mechanism as the value of 6mm and other parameters into the above equation, it can be obtained that the offset force P= 12,409 N, the deflection angle of the lower end of the main shaft is 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001479_j.phpro.2011.03.153-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001479_j.phpro.2011.03.153-Figure1-1.png", + "caption": "Figure 1: Schematic cross section of zinc coated steel sheets in overlap configuration", + "texts": [ + "de High speed imaging can further be used to gain a better understanding of the fluid dynamics of the melt pool and the spatter formation. For the simulation OpenFOAM is used. This software is written in C++ and is based on the Finite Volume approach. To simulate the laser deep penetration welding process [1, 2 and 3] many coupled physical effects have to be considered. In a first step the governing physical effects have to be determined. The keyhole and melt pool dynamics strongly depend on the geometry and the heat flux of the melt pool (Fig. 1). The flow inside the melt pool and the energy dissipation inside the fluid and the solid phase are modelled as a system of coupled nonlinear partial differential equations. The flow characteristic of the molten steel is described as an incompressible fluid by the Navier-Stokes equation. The free surface of the molten steel is treated by the Volume of Fluid (VOF) method. The VOF method is a numerical approach for tracking and locating the free surface of the fluid-air interface [4]. Within the model the energy flux and the physics of the phase transformation are considered" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003124_2019-01-0028-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003124_2019-01-0028-Figure2-1.png", + "caption": "FIGURE 2 Schematic of a single-cone synchronizer [6]", + "texts": [ + " FIGURE 1 Components of power loss in vehicle transmission \u00a9 S A E In te rn at io na l The function of synchronizers is to eliminate the relative rotational speeds between shafts and gear wheels in order to permit a smooth gear change. It is widely employed for gear shifting in Manual Transmissions (MT), Dual-Clutch Transmissions (DCT) and Automated Manual Transmissions (AMT). Like wet clutches [6], during engagement and disengagement processes of synchronizers, frictional heat is generated which causes deterioration and wear of the friction surfaces. Therefore, for a better service performance and longer life span of components, lubricants are appended between the synchronizer ring and the cone (Figure 2) to absorb heat and to eliminate dry friction. When the shift sleeve is in neutral position (disengaged), both the synchronizer ring and the cone sets are continuously rotating and the gap between them is fully or partly filled with lubricant. The relative motion between the ring and the cone induces viscous shearing of the lubricant film in the friction surface gap, which brings a drag torque on both friction surfaces. This drag torque transmitted by synchronizers is considered as a load-independent power loss because of the independence from the transmitted load" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002851_ssrr.2018.8468638-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002851_ssrr.2018.8468638-Figure2-1.png", + "caption": "Fig. 2. Assumed model", + "texts": [ + " In addition, it is assumed that the robot climbs directly from the lower side to the higher side of the slope and the twodimensional surface of the robot\u2019s side view is considered. Rajabi et al. proposed a geometric climbing condition when a tracked vehicle climbs over a step on flat ground[8]. In previous research, we applied this method to a circular cross-section obstacle on a slope, and derived the tip-over condition[11]. In this research, we extend this study and derive the geometric tip-over condition when a tracked vehicle with sub-tracks climbs over a circular cross-section obstacle on a slope. Fig. 2(a) shows the state when a tracked vehicle with subtracks climbs a circular cross-section obstacle on a slope. As described in previous research, the tip-over condition can be derived by the following equations, using the distance dG, the distance dO, and the angle \u03b8 between the main track and the slope. dG = dO (1) ddG d\u03b8 = ddO d\u03b8 (2) This is based on the observation that the robot can climb over the obstacle if its center of gravity reaches just above the contact point between itself and the obstacle. Equation (1) holds when the robot climbs over an obstacle of any diameter. Equation (2) holds when the robot climbs over an obstacle with the maximum diameter that it can climb over. The detailed reason why Equation (2) holds is described in [11]. This condition is the same regardless of whether the robot has sub-tracks. On the other hand, the distances dG and dO are different when the robot has sub-tracks. These are shown by the following equations, using the character in Fig. 2(b). dG = lr cos (\u03b8 \u03b8r + \u03d5) rr sin\u03d5+ lm 2 cos (\u03b8 + \u03d5) + \u221a X2 + Y 2 cos ( \u03b8 + \u03d5+ tan\u22121 Y X ) (3) dO = { lr cos (\u03b8 \u03b8r) + rm sin \u03b8 d 2 (1 + cos \u03b8) tan\u03d5 + d 2 (1 + cos \u03b8) rr lr sin (\u03b8 \u03b8r) + rm cos \u03b8 tan \u03b8 } cos\u03d5 (4) The distances dG and dO vary according to the sub-track angles \u03b8f and \u03b8r. X and Y indicate the position of the robot\u2019s center of gravity and are represented by the following equations. These values will vary according to sub-track angles \u03b8f and \u03b8r because the sub-tracks also have mass. X = [ Mmxm +Mf { lm 2 + ( lf 2 + xf ) cos \u03b8f yf sin \u03b8f } +Mr { lm 2 ( lr 2 xr ) cos \u03b8r + yr sin \u03b8r }] /M (5) Y = [ Mmym +Mf {( lf 2 + xf ) sin \u03b8f + yf cos \u03b8f } +Mr {( lr 2 xr ) sin \u03b8r + yr cos \u03b8r }] /M (6) The maximum diameter of an obstacle that a robot with a certain sub-track angle can climb, and the angle \u03b8 at the moment when the robot climbs over the obstacle are derived by combining the above equations and solving them simultaneously" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001361_transcom.e93.b.2901-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001361_transcom.e93.b.2901-Figure3-1.png", + "caption": "Fig. 3 Neighbor selection with dodecahedron (a) and proof of RNG \u2286 PYG for k \u2265 8 (b).", + "texts": [ + " In this section, we propose an YG-based topology control with Platonic solid (PYG) for 3D wireless ad-hoc sensor networks. In PYG, each node broadcasts a HELLO message including its location information with the maximum transmit power. After collecting HELLO messages, each node u divides the 3D space, a sphere centered at node u, into k equal cones by using a Platonic solid (i.e., a regular k-hedron) and selects the neighbor with the lowest link cost (e.g., the shortest distance or the lowest minimum transmit power) in each cone. For example, shown in Fig. 3(a), node u divides its 3D space into 12 equal cones by using a dodecahedron and selects the closest neighbor in each cone. (It seems that node u selects more than one neighbor in some cones because 3D cones overlap on figure.) After the neighbor selection, each node u broadcasts its selected neighbor set, uN. When a node u is received vN from node v, it compares two selected neighbor sets (i.e., uN and vN) and inserts certain edges into two edge sets: default edge set, EPYG, and symmetric edge set, ES PYG. Both sets are defined as follows: edge(u, v) \u2208 EPYG iff. v \u2208 uN or u \u2208 vN edge(u, v) \u2208 ESPYG iff. v \u2208 uN and u \u2208 vN (1) That is, an edge(u, v) is remained in GPYG = (V, EPYG) if more than one of two nodes u and v select the other node, whereas it is remained in GS PYG iff both node u and v select each other. Here, SPYG denotes the symmetric YG with Platonic solid and is similar to the symmetric Yao graph in 2D [2]. GPYG is connected for all Plastic solids (i.e. k \u2265 4). For example, shown in Fig. 3(a), although node w is not selected by node u (and by other nodes), it can select node u (and/or other nodes) and connect itself to the network. However, GS PYG is connected when using an octahedron, dodecahedron, and icosahedrons (i.e. for k \u2265 8). This result can be proved by comparison with RNG as shown in Fig. 3(b). It is well known that RNG guarantees the connectivity of 3D wireless ad-hoc networks [5]. In RNG, an edge(u, v) is remained if a region, looks like a Rugby ball, does not contain any other node, where the region is formed as the intersection of two spheres centered at two nodes u and v, and with radius \u2016u, v\u2016 (Refer Fig. 1(b)). SPYG includes an edge(u, v) if nodes u and v have the lowest link cost in C(v, u) and C(u, v) respectively (e.g., two nodes u and v are closest to each other in each 3D cones). That is, in SPYG, an edge(u, v) is remained if a region, formed as the union of two 3D cones with slant height \u2016u, v\u2016, does not contain any other nodes. Note that the angle \u2220avc = \u03c0 steradian in Fig. 3(b). The Rugby-ball-like region of RNG encloses two 3D cones of SPYG if the apex angle of the cones is equal or smaller than \u03c0/2. The apex angle of the 3D cone made by a regular k-hedron is 4\u03c0/k steradian. Therefore, the region (where edge(u, v) is removed if other node w is located in) of RNG can enclose that of SPYG if k \u2265 8. This means that the edge set of RNG includes that of SPYG for k \u2265 8. Since RNG guarantees the network connectivity, it is proven that GS PYG is connected for k \u2265 8 Previous topology control algorithms use Euclidean distance as the link cost for the topology construction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003864_s11012-019-01081-5-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003864_s11012-019-01081-5-Figure4-1.png", + "caption": "Fig. 4 Linear (a), planar (b), and spatial (c) translational motion of the 3-CPCR manipulator", + "texts": [ + " Displacements of dimension 0 (trivial case): \u2022 Identity {I}, or the null displacement; it is obtained by blocking the six inputs, ai and ai. 2. Displacements of dimension 1: \u2022 Rectilinear translational motion {Tu}; it is obtained either by blocking all except one of the translational inputs (ai), or by blocking the rotational inputs and relating the three translational inputs in such a way that only one is independent. If a translation oriented according to the components of a certain vector (s, t, v) is required, it is necessary to link the three translational inputs among themselves (see Fig. 4a). Taking into account Eq. (30), the following relation among the inputs is obtained: _a1=v \u00bc _a2=t \u00bc _a3=s \u00f040\u00de 3. Displacements of dimension 2: \u2022 Planar translational motion {Tu,v}; it is obtained either by blocking all except two of the translational inputs (ai), or by blocking the rotational inputs and relating the three translational inputs by a linear equation. Similarly to the rectilinear translation, if a planar translational motion included in a plane perpendicular to a certain vector of components (s, t, v) is required, the three translational inputs must be linked to each other (Fig. 4b). Taking into account Eq. (34), the following relation among the inputs is obtained: _a1v\u00fe _a2t _a3s \u00bc 0 \u00f041\u00de 4. Displacements of dimension 3: \u2022 Spatial translational motion {T3}; this is achieved by blocking all rotational inputs and actuating the three translational inputs, as shown in Fig. 4c. Derived from the original 3-CPCR, the resulting manipulator is the 3-PPCR. The directions of translation are given by Eq. (34). \u2022 Planar motion {Fu,v}. As explained in the previous section, only in configurations where one of the rotational inputs is actuated and the remaining two rotational inputs are kept constant and equal to p=2 does the axis of rotation of the moving platform translate in a direction perpendicular to that axis. This axis is always parallel to one of the coordinate axes. By conveniently selecting the two translational inputs in such a way that they define two translational motions in accordance with the coordinate axes and perpendicular to the axis of rotation, planar motion is achieved" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000908_0020-7403(65)90021-4-Figure12-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000908_0020-7403(65)90021-4-Figure12-1.png", + "caption": "FIG. 12. Effect of geometr ical imperfec t ions on the post-cr i t ical equi l ibr ium pa ths of cont inuous f rameworks .", + "texts": [ + " cons tan t of the rest ra ining spring and c the spring constant at D, and i f0 and ~ are posi t ive as shown, the equi l ibr ium pa th is g iven exact ly by t ~ o ~ .m r) -rj] e o ~ r c s i n t~o-~-~ W ~ cos ~ sin 0 (37) where W L w - (38a) 2c and C o ] \u00a2 = - - C For small values of 0 and ~ equat ion (37) can be wr i t ten approx imate ly w = l + k - ( l + k ) - k 0 ( l + \u00bd s i n y ) t a n y + . . . The pa th has ex t remes for posi t ive value of 0 \u00b0 when 0 = _+ k(l + \u00bd sin~.,) t a n ~ (38b) (39) (40) The posi t ive sign of 0 corresponds to the m a x i m u m or peak load pa ramete r and the nega t ive sign to the min imum, see Fig. 12. The peak value o f w is given by Wr~x = 1 + k - (2 + \u00bd sin 7) ~/[k(l + k) tan 7] O\u00b0t + ' \" (41) and the min imum value by Wml n = 1 + k + (2 + \u00bd sin y) ~/[k( 1 + k) t an ~,] ~i + . . . (42) The equi l ibr ium pa ths for a nega t ive va lue of 0 \u00b0 are shown dashed in Fig. 12. One branch favours the stable mode from the onset of loading. The o ther approaches the perfect pa th f rom above as indicated in the figure. Both forms are in agreement wi th the exper imenta l results in Ref. 13. 680 S . J . BaITVEC In the case of pin-jointed f rames n, 12 the model in Fig. 11 is modified by omi t t ing the spr ing a t A. Then without the spr ing the equi l ibr ium pa th is ob ta ined f rom (37) by se t t ing k = 0. On expand ing , approx imate ly , w = 1 - - 0 - \u00bd(tan= 7+- ~t) 02+ \"\" (43) E q u a t i n g the first der iva t ive wi th respect to 0 to zero, gives to the order of 0 \u00b0i, 00t " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003110_j.cad.2019.01.001-Figure12-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003110_j.cad.2019.01.001-Figure12-1.png", + "caption": "Fig. 12. Cut view of the mesh around a propeller: In rows, (a, b) horizontal and (c, d) longitudinal cut views. In columns, (a, c) initial and (b, d) final meshes.", + "texts": [], + "surrounding_texts": [ + "In this example, we compare the results obtained by applying our method using successive small rotations and only one large rotation. To this end, we first rotate the mesh around a propeller, see Figs. 12(a) and 12(c), seventy-two degrees in increments of one degree. Note that this angle corresponds to the angle between two consecutive blades of the propeller. At each increment, we solve the non-linear problem using the proposed augmented Lagrangian method. Then, we compare the results by imposing one large rotation to thewholemesh. Themesh is composedof 180167 linear tetrahedra and 32868 nodes. At each step, the augmented Lagrangian method is initiated using the optimized mesh of the previous step. Nevertheless, note that the mesh quality is computed taking into account the initial mesh. Similarly, the Lagrange multipliers and the penalty parameter are initialized using the values of the previous step. Figs. 12(b) and 12(d) show the mesh at the end of the displacement process. In Fig. 13, we show the whole rotation process in increments of six degrees. Note that the mesh accommodates the movement of the propeller in order to obtain a valid mesh composed of highquality elements. As the propeller rotates, it drags the elements at the outer part of the mesh in order to obtain a valid mesh. Fig. 14 shows the evolution of the minimum quality element along the rotation process. At the first iteration, the minimum quality is one, and it starts to decrease as the propeller rotates. Although in the first iterations the mesh quality rapidly decreases, in the rest of the morphing process the minimum quality roughly remains constant. At the end of the morphing process, the worst element has a quality of 0.40. Note that in this example we have applied the rotation of the propeller in small steps. Nevertheless, we have also morphed the mesh from the starting position to the final one in one step. Thus, we apply the proposed formulation to rotate the mesh around a propeller 72 degrees. The whole process takes four augmented Lagrangian iterations, and nine non-linear solves. The final configuration is computed in 636 s. Fig. 15 shows the final configuration of the propeller by evolving the rotation in increments of one degree 15(a), and by directly imposing the final configuration 15(b). Note that both configurations present similar results of the final mesh. The maximum difference between corresponding nodes is 3.7 units, and the whole geometry measures 400 units. Thus, the relative error between the two meshes is less than 1%. Furthermore, the minimum element quality of both meshes is similar. In the case of sub-stepping, the minimum quality is 0.4, while in the case of directly imposing the final configuration, the minimum quality is 0.39." + ] + }, + { + "image_filename": "designv11_33_0000945_ssd.2010.5585533-Figure18-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000945_ssd.2010.5585533-Figure18-1.png", + "caption": "Figure 18. 3-D flux path trough the magnetic ring, the collector and the stator yoke.", + "texts": [ + " These are described as follows: \u2022 axially in the stator yoke (see figure 12), \u2022 radially down through the collector and crossing the air gap down to the magnetic ring (see figure 12), \u2022 radially then axially in the magnetic ring (see figure 12), \u2022 axially-radially in the claws (see figure 13), \u2022 radially up in air gap facing the stator laminations (see figure 14), \u2022 radially up in the stator teeth and circumferentially in the stator core back (see figure 15), \u2022 radially down in stator teeth facing the two adjacent claws (see figure 16), \u2022 radially-axially in the adjacent claws (see figure 16), \u2022 axially then radially in the magnetic ring (see figure 17), \u2022 radially up through the air gap and the mag netic collector on the other side of the machine (see figure 17), \u2022 axially in the other side of the stator yoke (see figure 18). 3.2.2. Leakage Flux Paths Not all the flux produced by the excitation winding and the armature contributes to the EMF generation. Different leakage fluxes have been distinguished in the CPAES. The two main ones are \u2022 the leakage flux linking adjacent claws, \u2022 tow dimensional flux paths which flow through the magnetic circuit as follows: - axially in the stator yoke, - radially down through the magnetic collector and air gap el, - radially then axially in the magnetic rings holding the claws, - axially in the claw, - radially in air gap e2 and in the stator teeth" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001432_iros.2011.6094612-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001432_iros.2011.6094612-Figure1-1.png", + "caption": "Fig. 1. Generic layout of the IMU Design", + "texts": [], + "surrounding_texts": [ + "The IMU used in this method consists of two tri-axial accelerometers and one uni-axial rate gyro. The basic geometry of the setup is shown in Fig. (1). The IMU is intended to be mounted to the rotating limb some distance L from the axis of rotation, which is not known. The distance between the accelerometers is dL and is precisely controlled during manufacturing. The direction along the limb is defined as x; the direction tangent to the arc of rotation is defined as y; and, the direction perpendicular to x and y is defined as z. Without loss of generality, the gravitational vector is constrained to the x, y plane for this discussion. The accelerometers are oriented so they record the accelerations in the x, y, and z directions. The measured acceleration will be defined by the accelerometer number first and the axis of measurement second; for example, the acceleration measured by accelerometer number one in the x direction will be denoted as a1,x. If the accelerometers are mounted askew from the correct orientation, it is assumed the measured readings will be rotated into the proper frame before implementing the algorithm discussed below. The rate gyro, RG, is oriented so it measures angular velocity about the z axis; exact placement of the rate gyro along the IMU is not required as its measurement is independent of placement. Animal joint kinematics allow the assumption that all joint rotation axes are approximately parallel limiting the accelerations to centripetal, angular, gravitational, and translational (no coriolis terms). It is assumed that the measurements will be taken in an inertial frame; so, any pure translational accelerations have to arise during a gait cycle. Since animal joints are rotary in nature, any purely translational accelerations occur when two or more joints move at exactly the correct rate to cause the linking segment to experience a purely linear motion. It is assumed that this exact rotational rate match-up will occur only momentarily during a gait cycle and can be neglected when calculating the inclination angle. At the moments in the gait cycle where this assumption does not hold, the calculated inclination will be biased as it will report the direction of the vector created by combining the linear acceleration and gravity and not gravity alone. Because a Kalman filter is being used to combine this estimate of inclination with a integration estimate, brief changes in the calculated inclination due to translation acceleration or erroneous sensor measurements will be filtered out. Given these assumptions, the accelerations recorded by the accelerometers are described by (1). a1,x = (L+ dL)\u03b8\u03072 + g cos (\u03b8) a1,y = \u2212(L+ dL)\u03b8\u0308 \u2212 g sin (\u03b8) a2,x = L\u03b8\u03072 + g cos (\u03b8) a2,y = \u2212L\u03b8\u0308 \u2212 g sin (\u03b8) (1) Where L is defined as the distance from the measurement center of a2 to the instantaneous center of rotation as projected onto the x axis; dL is the distance between the measurement centers of a1 and a2; and, \u03b8, \u03b8\u0307, and \u03b8\u0308 are, respectively, the angle relative to gravity, the first time derivative of that angle, and the second time derivative of that angle. The difference in the two accelerometer measurements gives the angular acceleration as defined in (2). \u03b8\u0308 = a2,y \u2212 a1,y dL (2) The relationships defined in (1) can be solved for \u03b8 independent of distance and dynamics in four different solution regimes: 1) Static: \u03b8\u0307 = \u03b8\u0308 = 0 2) Constant Velocity: \u03b8\u0307 6= 0, \u03b8\u0308 = 0 3) Initial Acceleration: \u03b8\u0307 = 0, \u03b8\u0308 6= 0 4) Dynamic: \u03b8\u0307 6= 0, \u03b8\u0308 6= 0 The solution to the static regime, denoted \u03b8Static, is the trivial solution given in (3) and is the classic method used to measure inclination in the works discussed in the introduction. \u03b8Static = Atan2 (\u2212ay, ax) (3) It can be shown that the solution to both the constant velocity and the initial acceleration regimes is equivalent to the solution to the dynamic regime. The solution to the dynamic regime, denoted \u03b8Dynamic, is derived as follows: Rearranging the equations in (1) and solving for length terms gives: L+ dL = a1,x \u2212 g cos (\u03b8) \u03b8\u03072 L+ dL = \u2212a1,y + g sin (\u03b8) \u03b8\u0308 L = a2,x \u2212 g cos (\u03b8) \u03b8\u03072 L = \u2212a2,y + g sin (\u03b8) \u03b8\u0308 (4) which can be combined to remove the dependence on distance: \u2212a1,y + g sin (\u03b8) \u03b8\u0308 = a1,x \u2212 g cos (\u03b8) \u03b8\u03072 \u2212a2,y + g sin (\u03b8) \u03b8\u0308 = a2,x \u2212 g cos (\u03b8) \u03b8\u03072 (5) the ratio of these equations removes the dependence on angular acceleration and velocity: a1,y + g sin (\u03b8) a2,y + g sin (\u03b8) = a1,x \u2212 g cos (\u03b8) a2,x \u2212 g cos (\u03b8) (6) cross multiplying gives: a1,ya2,x + a2,xg sin \u03b8 \u2212 a1,yg cos \u03b8 = a2,ya1,x + a1,xg sin \u03b8 \u2212 a2,yg cos \u03b8 (7) which can be rewritten as: A cos (\u03b8) +B sin (\u03b8) = D (8) where: A = a2,y \u2212 a1,y, B = a2,x \u2212 a1,x, D = a1,xa2,y \u2212 a1,ya2,x g the cosine and sine terms can be combined and can be rewritten as: A cos (\u03b8) +B sin (\u03b8) = C sin(\u03b8 + \u03b4) = D (9) where: \u03b4 = Atan2 (A,B) , C = sign ( A sin \u03b4 )\u221a A2 +B2, sign (x) \u2261 x |x| Finally, (9) can be solved for \u03b8Dynamic according to (10). \u03b8Dynamic = Atan2 ( D/C, \u03b6 \u221a 1\u2212 D2 C2 ) \u2212 \u03b4 (10) where: \u03b6 = sign [ cos ( \u03b8\u0302 + \u03b4 )] , \u03b8\u0302 = estimate of current inclination angle. This calculation of \u03b8Dynamic provides an unbiased dynamic measurement of angle relative to the gravitational vector because it is independent of angular and centripetal acceleration and their respective moment arms. Determining when to use \u03b8Dynamic or \u03b8Static and calculating a good \u03b8\u0302 will be discussed in further detail in section III." + ] + }, + { + "image_filename": "designv11_33_0002963_roman.2018.8525601-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002963_roman.2018.8525601-Figure1-1.png", + "caption": "Fig. 1: Structure of a one-section cable-driven continuum robot.", + "texts": [ + " Finally, Section VI concludes the paper with remarks on the future implications. Continuum robots are flexible structures that have the ability to bend their shapes smoothly to suit both the task and the environment. Cable-driven and pneumatic-driven actuations are commonly applied in continuum robots to achieve the desired tip bending. Although cable-driven continuum robots are the focus of this research, application of the proposed DGMP framework to pneumatic-driven robots is straightforward. The structure of a one-section cable-driven continuum robot is shown in Fig. 1. Three cables are attached to the robot\u2019s tip through guidance disks to provide the required moment by varying their respective lengths. The constantcurvature approach [18] is assumed to find the forward kinematics that relate the tip pose bTi of section i to its cables lengths li with respect to its base. Based on this assumption, a one-section continuum robot is represented as a segment of a circle located in 3D space. This arc is characterized by its length si, curvature \u03bai and direction of curvature \u03c6i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002459_robio.2017.8324729-Figure10-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002459_robio.2017.8324729-Figure10-1.png", + "caption": "Fig. 10: Overview of the developed quadrotor UAV", + "texts": [ + " This section describes flight experiments for the proposed shape estimation method explained in Section III and obstacle avoidance flight proposed in Section IV. In the experiment, one end of the cable is connected on a prop and the other end is attached on the UAV as shown in Fig. 8. This experimental setup imitates the relative motion between UAVs for fundamental verification. The motion capture system is used to observe the position of UAV and the measured position is fed back to the UAV for flight control as shown in Fig. 9. The UAV used in this research is originally developed to satisfy the specification of cable handling. Fig. 10 shows the overview of the developed quadrotor UAV. The UAV is autonomously controlled by simple PID feedback controller on the originally developed on-board flight controller system. Firstly, we carry out the verification of the shape estimation method of the cable proposed in Section III. To verify the shape estimation method of the cable, multiple markers are attached on the cable itself for comparison between the real shape and the estimated shape. The UAV flies to hover and move arbitrary directions and shape estimation algorithm tries to estimate the shape of the cable in real time during the flight" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001290_j.mseb.2010.12.001-Figure10-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001290_j.mseb.2010.12.001-Figure10-1.png", + "caption": "Fig. 10. Temperature distribution in model of conducting element (Cu/Ag foil strip, width 620 m, thickness 50 m) with hole cut by use of Nd:YAG laser ( = 355 nm; 30 ns) due to current flow of 4 A.", + "texts": [], + "surrounding_texts": [ + "F kanth \u2013 mad\nb\nQ\nw t\nb fi c a m i fi w w\nF t f\nig. 6. High resistivity microareas made by laser material removing on two-layered power density 1.4 \u00d7 108 W/m2; (c) \u2013 metallographic cross-sectional view of HRMA\ny the equation:\n= 1\n(1 + \u02db(T \u2212 T0))(V2 X + V2 Y ) (7)\nhere 0 \u2013 is resistivity in 293 K, \u02db \u2013 thermal coefficient of resisivity.\nThe simulation of heating of conductive elements with holes y means of finite elements method in COMSOL environment conrmed an extreme rise of temperature near the hole due the electric urrent of 4 A (Fig. 8). Computations were performed for the switchble current of 4 A.The temperature of the narrowing has exceeded elting temperature of base metal by about 150 K. Investigations nto switching properties of fuse links with such elements con-\nrmed their ability to break of the short-circuit currents. However, ith regards to the regarding value of prearcing integral the results ere not satisfactory.\nig. 7. Microholes made by Nd:YAG laser, = 1064 nm, imp = 4 ms, focus length of he lens system 50 mm: (a) \u2013 in Cu/Ag foil, thickness 70 m, Eimp = 3.2 J; (b) \u2013 in Ag oil, thickness 70 m, Eimp = 3.9 J.\nal/Ag wire by Nd:YAG laser pulse, i = 4 ms: (a) \u2013 power density 1 \u00d7 108 W/m2; (b) e on two-layered kanthal/Ag wire, power density 1.4 \u00d7 108 W/m2.\n3.2. Microholes in thin foil made by thermal and ablative laser cutting\nThe rise of mass to be evaporated in a switch arc due to a toroidshaped edge of the hole (Fig. 7) was the main reason of less than excellent switching properties of such fuse links.\nYet another approach to creation of narrowings and holes sequence in stripes of thin metallic foil called for application of laser microcutting. Holes and narrowings were produced in Ag and two-layered Cu/Ag foils with thickness 25\u201350 m and width 300\u2013650 m. Two laser systems were used: the Nd:YAG laser (Coherent AVIA 355-14, = 355 nm, imp = 30 ns, average power 6 W, frequency 50 kHz) and the fiber laser (SPI Pulsed Fiber Laser 20 W, = 1070 nm, average power 6\u201316 W, imp = 60 ns, frequency 25\u2013250 kHz). The beam of the Nd:YAG laser, focused to a spot of 30 m diameter, was moved with a speed 400 mm/s, in a shielding gas (nitrogen) atmosphere. The beam of the fiber laser was focused by lens with the focus length 163 mm to a spot approximately 35 m. It was scanned with a speed 100\u2013800 mm/s without a protective gas. The samples were fixed to the ceramic holder, which allowed good thermal conditions and proper adjustment. There were free spaces in the holder below processing holes, which significantly improved the quality of final structure.The size of the smallest possible structures fabricated via the demonstrated techniques could be compared to the double spot size, which for the above mentioned laser system was about 60 m. To attain an even smaller size, the focal length of about 100 mm or 53 mm is needed, along with a higher expanding rate. The process had an ablative character so the holes were cut out with the dimensional accuracy about half of spot size i.e. \u00b1(10/15) m.Obtaining the highest precision of the cutout was possible thanks to the use of ablative process based on micromachining with nanosecond laser pulses of a Nd:YAG laser with wavelength of 355 nm and pulse width 30 ns (Fig. 9a and b). Experiments with use of the fiber laser with wavelength of 1070 nm, pulses duration of 60 ns, average power 9 W, frequency 65 kHz and the scanning speed 400 mm/s showed that also in this case it was possible to", + "F , width d\na (\nw F\nl m m e s\n(\nm\nc\np\nt\nc a w fi\nc n c o i l f m s a e\n4\na i\nig. 11. Temperature distribution in model of conducting element (Cu/Ag foil strip ue to current flow of 4 A.\nchieve good accuracy and satysfying quality of the hole edge Fig. 9c).\nTemperature distribution in a model of conducting elements ith holes of geometry like in Fig. 9 has been shown in\nigs. 10 and 11. It was achieved with a rapid rise of temperature simiar to that characterizing elements with holes made by laser elting\u2013evaporating (Fig. 7). At the same time the volume of aterial, that should be melted and evaporated when the current xceeded rated value, was significantly lower than the correponding volume for elements with holes of toroid-shaped edges\nFig. 7). Investigation into switching properties of conductive eleents with the holes made by nanosecond laser micromachining onfirmed their excellent features. Results of simulation of a temerature distribution due to a 4 A current were verified through hermographic analysis.\nProcessing time of one fuse element was about one second and it ould be optimized further. Moreover the edge quality of holes was dequate, because parameters of the laser beam for the fiber laser ere very stable. Details cutting by this method did not require any nishing treatment, because of small burr-effect.\nThe conductive elements with narrowing for fuse links are ommonly created by mechanical punching. However of this techology is restricted due to a decrease in mechanical durability of onductive elements. In the case of very thin and narrow foil strips r when high grade of narrowing is needed, the use of this mechancal technology is simply impossible. On the other hand precise aser cutting produces excellent results and is highly recommended or microfabrication of parts with dimensions below few hundreds\nicrometers To use this method in mass production, the cutting ystem should be equipped with automatic subsystem for fixing nd transporting of the foil strip under scan-head. It can also be asily coupled with a computer control system.\n. Conclusion\nThis papers presents examples of laser microtechnologies, pplicable to electronics and electrical engineering. Manufacturng of the high resistivity area in thin wires was achieved by laser\n[ [\n620 m, thickness 50 m) with hole cut by use of fiber laser ( = 1070 nm; 60 ns)\nremelting of a two-layered FeNi/Cu wire or by removing silver layer of a two-layered Ag/kanthal. Determination of proper conditions for such laser microtreatment was possible through use of computer modelling. This modelling took into account proper spatial shaping of the beam, physical and thermal properties of materials in solid and liquid state as well as geometry of elements. Solutions to discussed problems were confirmed experimentally through creation of conducting micro-elements with help of a laser. These micro-elements were later used as miniature fuse-links. Well known operations such as laser melting and evaporation proved to be inadequate for production of holes in thin metallic foils (Nd:YAG laser with pulse of duration of milliseconds). On the other hand, use of nanosecond pulse laser microtreatment produced excellent results. Data from simulations of thermal state of elements with holes confirmed desirability of their features for the purpose of their use as fuse links.\nReferences\n[1] R. Pawlak, F. Kostrubiec, M. Tomczyk, M. Walczak, Proceedings of 24th International Congress on Applications of Lasers and Electro-Optics ICALEO, Miami, 2005, pp. s395\u2013403. [2] R. Pawlak, K. Cwidak, J. Sulikowski, M. Walczak, 10th International Conference Switching Arc Phenomena, Lodz, Poland, 2005, pp. s314\u2013315. [3] J. Sulikowski, K. Cwidak, W. Gregorczyk, Proceedings of XXVI Conference of International Microelectronics and Packaging Society (IMAPS), Warsaw 25\u201327 September, 2002. [4] A. Ehrhardt, M. Presia, K. Mertens, Elektrie 55 (10\u201312) (2001) 478\u2013487. [5] R. Pawlak, F. Kostrubiec, M. Walczak, M. Tomczyk, 11th NOLAMP Conference\nin Laser Processing of Materials, Acta Universitatis Lappeenrantaesis, Finland, 273, 2007, pp. 271\u2013281 (ISBN 978-952-214-412-6). [6] R. Pawlak, F. Kostrubiec, M. Tomczyk, M. Walczak, Proceedings of SPIE, vol. 5629, 2005, pp. 349\u2013360, ISBN 0-8194-5584-9. [7] J. Mazumder, Optical Engineering 30 (August) (1991) 8. [8] V. Alexiades, A.S Solomon, Mathematical Modelling of Melting and Freezing\nProcesses, Hemisphere Publishing Corporation, Washington, Philadelphia, London, 1993 (ISBN 1-56032-125-3). [9] M. Walczak, Modelling of phenomena occurring during laser treatment of\nconducting materials (in Polish) \u2013 Sympozjum \u201cWsp\u00f3\u0142czesne Problemy Fizyki Materia\u0142\u00f3w I Elektrotechnologii\u201d Zeszyty Naukowe Instytutu Elektrotechniki Teoretycznej, Metrologii I Materia\u0142oznawstwa Politechniki \u0141\u00f3dzkiej, No. 1/2004, 101-108, ISBN 83-921172-0-4, \u0141\u00f3dz\u0301.\n10] M. Walczak, Doctor\u2019s Thesis, \u0141\u00f3dz\u0301, 1992 pp. 19\u201327. 11] B.N. Buszmanov, J.A. Chromov, Fizyka cia\u0142a sta\u0142ego, WNT, Warszawa, 1973." + ] + }, + { + "image_filename": "designv11_33_0002933_icelmach.2018.8507125-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002933_icelmach.2018.8507125-Figure8-1.png", + "caption": "Fig. 8. Mode shapes and frequencies (a) 0th mode; (b) 2nd mode; (c) 3rd mode; (d) 4th mode.", + "texts": [ + " We need to focus on low-order mode shapes, including the zeroth (also called breathing mode), second, third and fourth circumferential mode, because their natural frequencies are low and generally concerned with significant vibration. The natural frequencies are 1160.4Hz, 278.03Hz, 730.04Hz, and 1547Hz, respectively. The zeroth mode is concerned with deformation that is uniform around the circular space. The frequency of zeroth mode always occurs in high frequency band according to Hoppe\u2019s formulas. Mode shapes of the system are shown in Fig. 8. To clearly see the details on teeth, the concentrated windings are hidden. With contribution of windings and frame, the mode frequencies are higher compared with bare stator core. B. Vibration Response and Spectrum Twenty-four test nodes are selected in the stator mechanical system along the circumferential frame surface, as shown in Fig. 9, in order to observe the vibration shapes and acceleration distribution. Through the analysis in Section IV, Part B, we can figure out that the radial force distribution in the region where the four unit motors are located is no longer circumferentially periodical due to electrical asymmetries" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002793_ilt-07-2017-0208-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002793_ilt-07-2017-0208-Figure1-1.png", + "caption": "Figure 1 Physical dynamic model of a spiral groove face seal and the geometry of the rotor", + "texts": [ + " The aim of the present study is to propose a closed-form contact model for a gas face seal under a nonparallel-plane contact condition. The nonparallel-plane contact condition corresponds to the local face contact of sealing rings arising from disturbances during the opened operation. As the modeling idea comes from the typical rotor dynamics, the closed-form model also suits for the axial rub-impact of rotor dynamics. The closed-form model is then demonstrated by the direct numerical contact model regarding Gaussian sealing surfaces. Figure 1 shows the simplified dynamic model of a spiral groove face seal (Green, 2002; Miller and Green, 2000, 2003; Hu et al., 2016b, 2017d). Then, the rotor tilt angle g r, the stator tilt angle g s and their relative tilt angle g rel can be decomposed into components along the x and y axes as: cr \u00bc g rcos v t\u00f0 \u00de; g rsin v t\u00f0 \u00de\u00bd cs \u00bc g x t\u00f0 \u00de; g y t\u00f0 \u00de crel \u00bc g x t\u00f0 \u00de g rcos v t\u00f0 \u00de; g y t\u00f0 \u00de g rsin v t\u00f0 \u00de ; (1) where v is the shaft speed and t means the time. A thin gas film exists between the coupled rings" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003852_j.mechmachtheory.2019.103628-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003852_j.mechmachtheory.2019.103628-Figure2-1.png", + "caption": "Fig. 2. The machine setting model of Formate\u00ae face hobbing process.", + "texts": [ + " By controlling the position of the conical grinding wheel, the generator of the grinding wheel can be tangentially in contact with the tooth profile curve of the FFHHG; (4) Sweep the generator of the grinding wheel along the root curve of the tooth to calculate the tool locations of tooth grinding; (5) Check interference for no-working side of grinding wheel; (6) Avoid interference by tilting the axis of the grinding wheel; and (7) Update the grinding wheel parameters and the tool locations of tooth grinding at the same time. The machine setting of Formate\u00ae face hobbing process is shown in Fig. 2. Firstly, the coordinate systems are established for cradle, cutter and gear separately. The coordinate system r = { O, i r , j r , k r } is rigidly connected with virtual cradle, in which O is the center of cradle, i r O j r is the plane of cradle, k r is the axis of cradle, and its positive direction points to inner of cradle. The coordinate system \u2032 c = { O c , i \u2032 c , j \u2032 c , k \u2032 c } is rigidly connected to the cutter, in which O c is the center of cutter, i \u2032 c O c j \u2032 c is the plane where the blade pitch point is located, k \u2032 c is the axis of cutter whose positive direction opposites to tip of cutter, and j \u2032 c is the projection vector of k \u2032 c on the cradle plane" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000253_978-3-642-00644-9_48-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000253_978-3-642-00644-9_48-Figure3-1.png", + "caption": "Fig. 3 Proe drawings of the mecanical part", + "texts": [ + " Despite the long-term purpose to design a fully distributed system, this first attempt has been based on a centralized approach, at least at the structural level. A PC has the control of the whole setup, taking images from the camera, extracting the position of the glasses in respect to the table shape and controlling the glasses based on the feedback they give (see figure 2). The glasses robjects are still in charge of all sensor pre-processing, detection of events and motor control. The robject has the following configuration: It is a glass fixed on a small cylindrical mobile robot having 10cm of diameter and 3.4cm high (figure 3). The robot has two differential wheels with a diameter of 2.6cm. The part of the robot that is glued to the glass can move around an axis and acts on a force sensor as a balance. The same mechanical part protects the electronic of spatter. Figure 4 shows the mechanism of the balance. A plexiglas frame (blue part in figure 3) fixes the battery, motors and the print circuit board. Wheel are actuated by stepper motors with 20 steps by revolution and a 50:1 gearbox. The torque is sufficient to move the robot with a payload of one kilogram. The electronics of the robot is based on the electronics of the e-puck robot1. The microcontroller used for controlling the robot is a DSPIC30F by Microchip running at 15 MIPS. It has 68 I/O pins of which 12 can perform analogical/digital conversion and it support two UART communications channels" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000395_978-1-4419-7344-3_3-Figure3.4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000395_978-1-4419-7344-3_3-Figure3.4-1.png", + "caption": "Fig. 3.4 The heat conducted per unit time, Q\u0307, through a flat plate of cross-sectional area, A, and thickness d, subject to a temperature gradient, (Th \u2212 Tc)/d. Joseph Fourier used such an arrangement to formulate his laws of conduction heat flow", + "texts": [ + " Joseph Fourier propounded the laws of heat conduction and developed the first analytic theory of heat in continuous bodies, his first discourses on these subjects given to the French Academy of Sciences being published in 1807-1808 [3]. His laws are based both on his quantitative measurements of conductive heat transfer through a solid material, and on his prodigious mathematical skills. The experimental observations were based on an iron plate steadily transferring heat through a uniform thickness, d, measured between its two broad surfaces, which were maintained at different uniform temperatures, Th and Tc, where Th > Tc, as is sketched in Fig. 3.4. Fourier noted that the rates of heat transfer per unit area through the plate, Q\u0307, often referred to figuratively as \u2018heat flow\u2019\u2014although nothing actually flows\u2014were proportional to the temperature difference established across the plate, Th \u2212 Tc, and were inversely proportional to its thickness, d. His observations may be expressed mathematically as the linear response function, Q\u0307 = k A (Th \u2212 Tc) d , (3.22) where k, a property specific to the plate\u2019s material, is the constant of proportionality, and bears the modern units of W/m-K" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002089_0954406217718857-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002089_0954406217718857-Figure4-1.png", + "caption": "Figure 4. Lateral leakage components between the gear\u2019s lateral side and the floating plate.", + "texts": [ + " The flow through the fixed orifice yields Q 1\u00f0 \u00de out \u00bc Qori \u00bc CdAori ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2p= p\u00f0 \u00de p \u00f03\u00de The flow through the triangular groove into the gear shaft\u2019s TSV yields (Figure 2) Qs tri \u00bc CdAtri ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 pi\u00fe1 pi\u00f0 \u00de= p\u00f0 \u00de p \u00f04\u00de The flow through the triangular groove into the ring gear\u2019s TSV yields the same as equation (4), and thus the flow through the triangular grooves yields Q 2\u00f0 \u00de out \u00bc Qs tri \u00feQr tri \u00f05\u00de The internal leakage consists of two parts: from the lateral clearance between the gears\u2019 lateral sides and the floating plates (Figure 4), and from the radial clearance between the tooth tips and the fillers (Figure 2). Due to the complexity of the gear geometry, the former part needs to be divided into additional two parts: from the tooth root and from the tooth side (Figure 4). The lateral leakage from the gear shaft\u2019s tooth root yields Qs1 l \u00bc b1 3 l 12 p\u00f0 \u00de l1 p \u00f06\u00de where is the fluid dynamic viscosity, a function of the pressure.29 An equivalent rectangular area with the length l2 equal to 0.75 times the tooth height is applied to evaluate the lateral leakage from the gear shaft\u2019s tooth side8 Qs2 l \u00bc b2 3 l 12 p\u00f0 \u00de l2 p \u00f07\u00de The lateral leakage from the ring gear\u2019s lateral sides yields the same as equations (6) and (7), and thus the lateral leakage yields Ql \u00bc 2 X Qs1 l \u00fe X Qs2 l \u00fe X Qr1 l \u00fe X Qr2 l \u00f08\u00de The radial leakage from the radial clearance between the gear shaft\u2019s tooth tip and the big filler yields (Figure 2) Qs r \u00bc b 3r pi\u00fe1 pi\u00f0 \u00de 12 p\u00f0 \u00de l3 b r" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001127_j.ijepes.2012.08.004-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001127_j.ijepes.2012.08.004-Figure4-1.png", + "caption": "Fig. 4. Slice of core iron and winding bounded by planes at mid-slot, mid-tooth.", + "texts": [ + " (c) The value of the potential function in each element is determined by the order of the finite element. The order of the element is the order of polynomial of the spatial co-ordinates that describes the potential within the element. The potential varies as a quadratic function of the spatial coordinates on the faces and within the element. 4.2. Boundary conditions The details of the induction motor are shown in Fig. 3. In this analysis, the two-dimensional domain of core iron and winding chosen for modeling the problem is shown in Fig. 4 and the geometry is bounded by planes passing through the mid-tooth and the mid-slot. The temperature distribution is assumed symmetrical across two planes, with the heat flux normal to the two surfaces being zero. From the other two boundary surfaces, heat is transferred by convection to the surrounding gas. It is convected to the air\u2013gap gas from the teeth, to the back of core gas from the yoke iron. The boundary conditions may be written in terms of dT/dn, the temperature gradient normal to the surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001516_iros.2011.6094416-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001516_iros.2011.6094416-Figure4-1.png", + "caption": "Fig. 4. The musculoskeletal leg robot.", + "texts": [ + " Table V presents the inner products of each subject\u2019s PC vectors of A-A ratios averaged and normalized at all run speeds. In all cases, most cosine values are close to 1.0. Muscle coordination 1 and 2 (pr1 and pr2), extracted from the A-A ratios, contribute to the joint angles and contain the kinematic meanings independent of each other. To confirm the kinematic meaning of these two patterns of muscle coordination in human running, we transferred their patterns to our human-like musculoskeletal leg robot. Figure 4(a) presents the human-like musculoskeletal leg robot. The robot consists of a skeletal model (Avice, Inc.) and eight pneumatic artificial muscles (PAMs) (Kanda Tsushin Kogyo Co., Ltd.) corresponding to the examined muscles, as illustrated in Fig. 2. The robot body parameters (segment length, segment mass, moment of inertia) are reconstructed from human-body parameters [9], [10] by attaching metal sheets to the robot. The muscle-attaching locations were decided by referring to [11]. The robot has three degrees of freedom, with the rotation of the hip, knee, and ankle joints in the sagittal plane", + " Figure 6 illustrates the leg robot\u2019s movements generated by the first or second pattern of muscle coordination in the case of subject C at 7km/h. As seen in Fig. 6, based on a hip joint, muscle coordination 1 creates a rotary motion in the limb of a toe, and muscle coordination 2 creates a bending and stretching motion of the toe. The same results were obtained for different subjects at different running speeds. This result implies that human running is described by a combination of units of motor function that contribute to the argument \u0398 and the moving radius L of a toe, based on the hip joint (Fig. 4(b)). This is an interesting result that makes us anticipate a new method of controlling a musculoskeletal robot by using patterns of muscle coordination. Ivanenko et al. achieved a similar result, although the analysis was performed in the joint space [3]. They decomposed some motions of a limb into two principal components by using the joint angles [\u03b81, \u03b82, \u03b83] T , based on the direction of gravitational force (Fig. 4(b)). According to their study, we analyzed the joint-angle space [\u03b81, \u03b82, \u03b83] T based on the direction of gravitational force using PCA. \u03b8 = [\u03b81, \u03b82, \u03b83] T can be expressed as follows: \u03b8 = 3 \u2211 i=1 w\u03b8i(t)p\u03b8i + \u03b8\u0304 (7) where w\u03b8i(t) is i-th PC score, p\u03b8i = [p\u03b8i1, p\u03b8i2, p\u03b8i3] T is the i-th PC vector, and \u03b8\u0304 = [\u03b8\u03041, \u03b8\u03042, \u03b8\u03043] is the mean vector. Figure 7 depicts the result of simulation for two principal components in the case of subject C at 7km/h as follows: \u03b8(t) = { w\u03b81(t)p\u03b81 + \u03b8\u0304 w\u03b82(t)p\u03b82 + \u03b8\u0304 (8) It seems that the motions of the musculoskeletal leg robot (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000104_s10895-010-0700-7-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000104_s10895-010-0700-7-Figure1-1.png", + "caption": "Fig. 1 a Scheme of the fluorescence measuring set-up [31]. b Fluorescence image of TEOS/CHIT/CPO/AChE spots", + "texts": [ + " After drying, spots were treated with glutaraldehyde solution (2.5% in potassium phosphate buffer, 50 mM, pH 7.0) for 2 min and washed with distilled water and phosphate buffer solution, respectively. Apparatus and the detection system setup BioRobotics MicroGrid, A High Throughput Automated Microarrayer (England) was used for spotting of sol-gel/ chitosan solution on the surface of the glass slides. For the fluorescent intensity measurement of the spots, Sensovation AG, Stockach (Germany) control system (Fig. 1a) was used [31]. Absorption measurements were performed with a UV\u2013Vis spectrophotometer (Perkin Elmer- UV/VIS Spectrometer-Lambda 2, (Germany)). Fluorescence spectral data were recorded on Horiba JOBIN YVON Fluoromax-4 Spectrofluorometer (Germany). The pH values of buffer solutions were adjusted with VWR digital pH-meter (Germany) calibrated with Merck pH standards of pH 7.00 and pH 4.01. The measuring arrangement\u2013 a fluorescence reader (Sensovation AG, Stockach) (Fig. 1a)\u2014consists of a lamp, a shutter, filters for excitation and emission and a 3-stage Peltier cooled CCD image sensor with integrated microlenses and a high quantum efficiency (up to 90%) for detecting the fluorescence signal. The chip array is located in front of the fluorescence camera which is gathering as much fluorescence light as possible. The excitation light, passing the filters for wavelength selection, is directed onto the front of the measuring cell via mirrors. Additional filters in front of the camera enable a reduction of the stray light and a preselection of the detection wavelength region [31]. The excitation light can be also directed to the backside of the cell for adsorption measurements. Such a modular set-up made the measuring console very flexible to various measuring regimes. The emitted radiation is taken by a deep cooled CCD camera. The image detected by means of the digital camera (Fig. 1b) consists of data regions from the spotted area together with a huge background region around. Technological spotting of the indicator volume onto the slide does not yield a uniform centred spotting but the spots jitter around the centre position. To reduce the background signal contributing to the averaged signal intensity taken from the spotted area, a spot recognition software module (Iconoclust, Clondiag GmbH Jena) was used to accumulate the signals in the spotted area only. Measuring procedure The TEOS/CHIT/CPO/AChE arrays were subjected to plain phosphate buffer solutions and buffer solutions containing different concentrations of ACh for 3 min and emissions were collected at 540 nm under 460 nm excitation wavelength", + " The photostability test of CPO dye in TEOS/CHIT matrix was performed with a steady state spectrofluorimeter in mode of time-based measurements. The data in Fig. 2b, acquired at 545 nm emission after 1 h of monitoring, reveals the excellent photo stability of CPO in TEOS/CHIT matrix. The fluorescence images of TEOS/CHIT/CPO/AChE spots were obtained by collecting the emission intensity at 540 nm under 460 nm excitation wavelengths with the measuring arrangement\u2013 a fluorescence reader (Sensovation AG, Stockach) shown in Fig. 1a (Fig. 1b). Statistic and dynamic quenching rate constants In the case of energy quenching by encounter complexes, the dynamical quenching, the fluorescence lifetime follows the Stern-Volmer relation [32] (Eq. 2): I0=I \u00bc 1\u00fe tFkq Q\u00bd \u00f02\u00de where I0 is the initial fluorescence intensity and I is the fluorescence intensity in the presence of quencher, \u03c4F fluorescence lifetime, kq rate constant and [Q] the quencher concentration. Diffusion-controlled (or diffusion-limited) reactions are reactions that occur so quickly that the reaction rate is the rate of transport of the reactants through the reaction medium (usually a solution)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003183_0954406219843954-Figure10-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003183_0954406219843954-Figure10-1.png", + "caption": "Figure 10. Radii of curvature (left) and transverse path of contact (right) of the virtual cylindrical gear according to FZG/Klein et al.9", + "texts": [ + " The virtual cylindrical gear is derived from the reference cone of the bevel gear or rather the hypoid gear. The transverse path of contact between pinion and wheel of the virtual cylindrical gear is divided in a number of sections (e.g. n\u00bc 10) to determine the local scuffing safety factor for n\u00fe1 points of contact. To locate the contact points, the coordinate gY is established, which originates in the pitch point C (\u00bc design point P of the bevel gear set). Towards the pinion tip, gY is defined as positive, and towards the pinion root, it is defined as negative. Figure 10 shows an exemplary path of contact including the coordinate gY. 9 In Figure 11, safety factors regarding scuffing calculated by use of the local calculation method based on a loaded tooth contact analysis are compared with safety factors determined by use of the standardized calculation method. For the bevel gear variant \u2018\u2018G0\u2019\u2019 as well as for the hypoid variant \u2018\u2018G31,75\u2019\u2019 a good qualitative and quantitative accordance can be shown. The standardized calculation method represents the real contact conditions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003222_ffe.12997-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003222_ffe.12997-Figure5-1.png", + "caption": "FIGURE 5 Theoretical distribution of the normal force F bti acting o gear (\u03b5\u03b1 < 2) and B, high\u2010contact ratio gear (2 < \u03b5\u03b1 < 3)5", + "texts": [ + " Listed parameters are necessary for subsequent definition of the contact model and the loading model. Principles and underlying mathematical expressions used for the calculation of points defining resulting gear profile and resulting parameters are presented and explained in the previous work.5 During the loading cycle, amplitude of normal load acting on a single tooth flank changes constantly. Theoretical load distribution and load sharing on a gear tooth for the low\u2010contact ratio (LCR) and high\u2010contact ratio (HCR) gearing are shown in Figure 5. In order to accurately determine values of loading that act on the tooth during the meshing cycle, the overall tooth deformation and deflection comprising tooth bending, tooth foundation deformation, and local deformation in the vicinity of contact must be taken into account. For the calculation of the load sharing between gear teeth in simultaneous contact, a number of methods mostly based on the deflection equality of teeth pairs in simultaneous contact were proposed in the literature.21,22 In the presented model, the load sharing, ie, change of the load acting on the flank of an individual tooth during the loading cycle, was determined based on values of stiffness of each tooth23 according to the procedure for determining total tooth deflection" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000028_s12206-009-1166-x-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000028_s12206-009-1166-x-Figure2-1.png", + "caption": "Fig. 2. Geometrical model of the 5-DOF parallel-serial hybrid robot.", + "texts": [ + " First, this hybrid robot has a complicated structure, and second, the general inverse kinematic analysis cannot calculate the 5-DOF spherical robot motion which lacks the 1-DOF. Therefore, a new inverse kinematic analysis method of the 5-DOF parallel-serial hybrid robot is needed. In this paper, the structure of this hybrid robot is shown. Then, the kinematic analysis methods are proposed. Also, the continuous path control algorithm is proposed. To verify usefulness, a prototype hybrid robot is tested using the positioning accuracy and the pose accuracy. The geometrical model of this hybrid robot is shown in Fig. 2 [1-3]. The origin point is defined as Op. The parallellink arm section consists of the constant vectors C1, C2, the linear motion vectors L3, L4, and the polar vectors P5, P6. \u2020 This paper was presented at the ICMDT 2009, Jeju, Korea, June 2009. This paper was recommended for publication in revised form by Guest Editors Sung-Lim Ko, Keiichi Watanuki. *Corresponding author. Tel.: +055 220 8452, Fax.: +055 200 8452 E-mail address: g08mm006@yamanashi.ac.jp \u00a9 KSME & Springer 2010 The serial-link arm section consists of the linear motion vector L7, the constant vectors C8, G9 with the rotation \u03b84 around the k-axis, and the tool center point (TCP) rotation \u03b85 around the w-axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002709_978-1-84996-220-9_5-Figure5.4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002709_978-1-84996-220-9_5-Figure5.4-1.png", + "caption": "Figure 5.4 (a) An object (wok) with predefined grasping positions for two arms of ARMAR-III (b) The 3D projection of the reachability spaces for both arms of ARMAR-III", + "texts": [ + " For each object to be grasped or manipulated by the robot, a collection of feasible grasps is stored in a database. This collection of feasible grasps holds information about the transformations between the end effector and the final grasping position, the type of grasp, the preposition of the end effector and some grasp quality descriptions. These database entries can be generated automatically (e.g., with GraspIt! [22] or GraspStudio [29]) or, like in the following examples, by hand. A wok with 15 feasible grasps for each hand of the humanoid robot ARMAR-III can be seen in Figure 5.4(a). To grasp an object o (located at position Po) with the end effector e by applying the grasp gk of the feasible grasp collection gce o, the IK problem for the pose Po k has to be solved: Po k = T\u22121 k \u2217Po. (5.4) To grasp a fixed object with one hand, the IK-problem for one arm has to be solved. In case of the humanoid robot ARMAR-III, the operational workspace can be increased by additionally considering the three hip joints of the robot. This leads to a 10 DoF IK problem. Typically, an arm of a humanoid robot consists of six to eight DoF and is part of a more complex kinematic structure", + " Since this IK-solver is used within a probabilistic planning algorithm, this approach fits well in the planning concept. The use of a reachability space can speed up the randomized IK-solver. The reachability space represents the voxelized 6D-pose space where each voxel holds information about the probability that an IK query can be answered successfully [4, 11]. It can be used to quickly decide if a target pose is too far away from the reachable configurations and therefor if a (costly) IK-solver call makes sense. The reachability space of the two arms of ARMAR-III is shown in Figure 5.4(b). Here the size of the 3D projections of the 6D voxels is proportional to the probability that an IK query within the extent of the voxel can be answered successfully. The reachability space is computed for each arm and the base system is linked to the corresponding shoulder. The reachability spaces can be computed by solving a large number of IK requests and counting the number of successful queries for each voxel. Another way of generating the reachability space is to randomly sample the joint values while using the forward kinematics to determine the pose of the end effector and thus the corresponding 6D voxel [4]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003243_1.4043114-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003243_1.4043114-Figure1-1.png", + "caption": "Fig. 1 Description of the aircraft states, inputs, coordinates, and forces", + "texts": [ + " The mass of the UAV remains constant while seeking solutions to optimization problems, as the change in fuel weight in the period of interest is small relative to the gross weight of the UAV. The differential equations governing the dynamics of the UAV (can be found in Ref. [40]) and are given by _x \u00bc V cos c cos v (1) _y \u00bc V cos c sin v (2) _z \u00bc V sin c (3) _c \u00bc g V n cos / cos c\u00f0 \u00de (4) _v \u00bc g V n sin / cos c (5) _V \u00bc T D m g sin c (6) where x, y, and z are the position of the aircraft center of gravity in the earth reference frame, c is the flight-path angle, v is the heading angle as illustrated in Fig. 1, V is the aircraft speed, g is the acceleration due to gravity, and m is the mass of the aircraft. T is the thrust, / is the bank angle, n is the load factor, and D is the drag. The load factor (n) is given by n \u00bc L mg and L \u00bc 1 2 qS CL0 \u00fe CLaa\u00f0 \u00deV2 (7) where L is the lift, q is the density of air, S is the planform area of the UAV wing, a is the angle of attack, and CL0 and CLa are the aerodynamic coefficients of lift. The expression for lift can be found in Ref. [41]. The expression for drag (D) is also taken from the literature [41] and is given by D \u00bc 1 2 qS CD0 \u00fe CL0 \u00fe CLaa\u00f0 \u00de2 peAR V2 (8) where e is the Oswald efficiency factor, AR is the wing aspect ratio, and CD0 is the constant parasitic drag" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002780_1.4041028-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002780_1.4041028-Figure1-1.png", + "caption": "Fig. 1 The motion of each cross section of the rod at length s and time t is determined by tracking the transformations of the body-fixed frame a\u0302 i (s; t) with respect to the inertial frame of reference e\u0302 i", + "texts": [ + " Section 4 presents some case studies to compare the proposed strategy with existing (but non user-friendly) approaches. The mathematical formulation of the dynamic rod model that we use [2] employs the classical approach of the Kirchhoff [36], which assumes each cross section of the rod to be rigid. The position and orientation of each cross section are determined in space s and time t by tracking the transformation of a body-fixed frame a\u0302i\u00f0s; t\u00de with respect to an inertial frame of reference e\u0302i that are shown in Fig. 1, where subscript i\u00bc 1, 2, 3. Vector R\u00f0s; t\u00de defines the position of the body-fixed reference frame a\u0302i\u00f0s; t\u00de relative to the inertial frame of reference e\u0302i. The spatial derivative of R\u00f0s; t\u00de, which we denote by r\u00f0s; t\u00de, is in the tangential direction along the centerline. Its deviation from the unit normal of the cross section determines the shear and stretch at the cross section. The change in magnitude of r\u00f0s; t\u00de quantifies the extension or compression along s, and the change in its orientation with respect to the body-fixed frame a\u0302i\u00f0s; t\u00de quantifies shear", + " (24) represents the skew-symmetric tensor associated with the vector m generated as follows from its components: em \u00bc 0 3 2 3 0 1 2 1 0 24 35 (25) and the Jws is given below: Jws \u00bc @ @w @s @w @j0 dj0 ds @j (26) Equation (20) is linearized around a guessed solution. This means that the terms AY; BY, and K\u0302\u00f0@Y=@s\u00deij 1 are calculated using the guessed solution so that Eq. (20) rendered integrable with respect to Yi j in space AYYi j \u00fe BYYi j 1 \u00bc H (27) The matrix H contains all of the known terms from previous time-step (i 1) as well as the linearization terms that depend on the guessed solution. In most scenarios, the boundary condition contains partial information on Y at one end (s\u00bc 0) and the rest is known at the other end (s\u00bc L). For example, in Fig. 1 the left-hand side of the rod is fixed by a clamp which imposes v\u00f00; t\u00de \u00bc 0 and x\u00f00; t\u00de \u00bc 0, while on the right-hand side, the external forces and moments, f\u00f0L; t\u00de and q\u00f0L; t\u00de, are prescribed. We use the shooting method at each time-step as explained by Sun et al. [43] to start integration from one end and match the boundary conditions at the other end. Alternatively, an assembled matrix approach can also be used to match the boundary conditions at both the ends simultaneously. Figure 2 shows the algorithm of how shooting method is applied to this problem" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002466_tvt.2018.2800777-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002466_tvt.2018.2800777-Figure3-1.png", + "caption": "Fig. 3 Experiment system of LSMM", + "texts": [ + " X, X 2017 2 Japan established equivalent 1/6 and 1/3 mass models of planet surface rovers in order to imitate the gravity of the Moon and Mars, respectively, as shown in Fig. 2. 2) Hanging compensation method The hanging compensation method uses the following suspension system in order to hang the vehicle and compensate for 5/6 of the lunar rover\u2019s vertical loads. This method was first used by NASA in the testing of a local scientific survey module (LSMM). The hanging compensation system used by NASA contained two subsystems; namely, a trailer platform following the rover, and a suspension platform providing the hanging force, as shown in Fig. 3. The purpose of the equivalent mass method was to diminish the mass of the vehicle in order to obtain vertical loads that were identical to those that would be experienced on the surface of a different planet. However, this method can also alter the dynamic properties of the vehicle. Additionally, the hanging compensation method was constrained by the range of the experimental suspension system. The ultimate objective of these gravitational imitation methods is to produce the same mechanical relationship between the wheel and soil; therefore, it was important to establish an accurate mechanical model describing the behavior of the rover\u2019s wheels [16]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000336_6.2010-2790-Figure13-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000336_6.2010-2790-Figure13-1.png", + "caption": "Figure 13. Details of wing clamp and representative distribution of scan points distribution. (A) A typical user specified grid of 250 wing scan points. Each point is sampled by the scanning laser vibrometer as the wing base is subjected to broadband a excitation. (B) A schematic of the clamp and wing in the open configuration shows (C) A schematic of a clamped wing in its \u201csandwiched\u201d state.", + "texts": [ + " We employed this system to extract the natural frequencies and corresponding modeshapes from specimens of hawkmoth forewing. The specific details of the system and the methodology it employs to perform system identification are outside the scope of this paper but can be found at their website (www.polytec.com/usa/). The basic experimental arrangement (Figure 12) is comprised of a shaker head that excites the wing\u2019s vibrational modes, a scanning laser vibrometer that measures the velocity response of the wing at predefined locations (Figure 13A), a reference laser vibrometer to measure the actual velocity output of the shaker head (the input excitation to the wing) and a control unit that enables the user to modify operating parameters of the vibrometers as well as select excitation signal waveforms for the shaker head. Of course, Polytec\u2019s interface computer and accompanying software is an integral part of the setup and is the primary user interface for interacting with each hardware component, defining test parameters, triggering test execution, and visualizing results", + " diately following wing liberation as described above, we mount shaker head rigidly mounted to optical table, and (E) close-up of shaker head with mounted sample wing. The red dot illustrates only one of hundreds of points on the wing where the laser vibrometer collects frequency response data. The green laser dot is held fixed on a rigid bolt threaded into the shaker providing a reference for the forcing input. Note that the red and green lines/dots on this figure are for illustrative purpose only. ibr \u2019s p 14 American Institute of Aeronautics and Astronautics to o subs tructures where the affect can most often be igno ne side. As illustrated in Figure 13B/C, the sides with foam face one another and serve to sandwich the root of the wing when the plates are drawn together with a securing locknut. This \u201csandwiching\u201d methodology, used across all clamp variants, ensures a tight fit of the wing and approximates a clamped condition. Early trial tests showed that using rigid plates without insulating foam damaged the veins at the wing root and induced deformation throughout the wing by forcing the veins at the root to unnaturally lie in a plane, effectively \u201csqueezing\u201d the natural camber out of the wing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001243_s11044-012-9325-8-Figure10-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001243_s11044-012-9325-8-Figure10-1.png", + "caption": "Fig. 10 Example of the balance of forces for the flexor digitorum profundis (FDP) at the direction change imposed by the MCP joint of a finger: (a) configuration of the joint with FDP tendon, the reference frame is fixed with the metacarpus bone; (b) balance of forces showing the tendon forces and reactions (on the tendon), in this case the reaction is applied in the structure by means of the synovial sheath", + "texts": [ + " 9, where the muscle is represented by three components: a contractive element in series with a spring and a second spring in parallel with the first two. With this representation both passive (using the springs) and active behaviors are modeled. The values for the springs and the expressions for the contractile element can be found in detail in [26].4 The muscles forces that are transmitted to the structure by the tendons are, in this work, modeled using analytical expressions derived from the balance of forces at each change in direction of the tendon (the direction change can be imposed by a joint or by the presence of ligaments), Fig. 10. The boundary condition that enables the solution of the balance of forces is that the force remains constant along the tendon5: \u2016F\u2032 FDP\u2016 = \u2016FFDP\u2016. With the reaction vector determined, it is applied to the structure by means of the \u03b2\u0302 of the link using Eq. (21) and Eq. (23) (where X\u0302ef is the spatial transformation from the base of the link to the point of application of the external force (ef) and F\u0302ef is the external spatial force). \u03b2\u0302 = \u2212( X\u0302ef )T F\u0302ef (23) This procedure, although applicable in most cases, has some special cases that need special care" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003297_j.wear.2018.12.040-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003297_j.wear.2018.12.040-Figure3-1.png", + "caption": "Fig. 3. Test rig assembly.", + "texts": [ + " The samples were cut using wire Electric discharge machining with dimension of 10mm x 10mm. Later the samples were mounted and polished using SiC sand paper with grit of 120, 240, 400, 600, 800 dan 1200 for ten minutes each, and finally with diamond paste of 1 \u00b5m. The optical microscope was used to observe and the microstructure at 20\u00d7 magnification. In order to obtain the friction coefficient of the fabricated dimples, tribological test was conducted using a linear reciprocating tribometer model TR 282, as shown in Fig. 3. The size of two samples for the tribological testing is a 10mm\u00d710m, and referred as lower and upper plates. The lower smooth plate is fixed in a lubricated condition and it slides in a reciprocating manner with the aid of a slider \u2013 crank mechanism. The upper plate is a textured specimen, which is fixed with Nomenclature DATT dynamic-assisted tool \u03b3 clearance angle \u03b2 rake angle r\u03b5 nose radius V cutting speed (m/min) f feed rate (mm/rev) amp amplitude (mm) fq frequency (Hertz) ar area ratio Ra surface roughness (\u00b5m) \u03b1 Inclination angle v direction of movement COF coefficient of friction arm and keeps stationary during the test" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000777_s0025654411040017-Figure12-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000777_s0025654411040017-Figure12-1.png", + "caption": "Fig. 12. Fig. 13.", + "texts": [], + "surrounding_texts": [ + "By setting \u03b4 = 0 in Eqs. (3.6) and (3.7) and by introducing new variables z1 and z2 by formulas z1 = z + a1\u03d5 and z2 = z \u2212 a2\u03d5, we obtain the equations z\u2022\u20221 + a11(c1z1 + \u03b31z \u2022 1) + a12(c2z2 + \u03b32z \u2022 2) = 0, z\u2022\u20222 + a21(c1z1 + \u03b31z \u2022 1) + a22(c2z2 + \u03b32z \u2022 2) = 0, (4.1) MECHANICS OF SOLIDS Vol. 46 No. 4 2011 a11 = 1 + a2 1 \u2212 a1f1, a12 = 1 \u2212 a1a2 \u2212 a1f2, a21 = 1 \u2212 a1a2 + a2f1, a22 = 1 + a2 2 + a2f2. (4.2) The dimensionless variables ak and fk (k = 1, 2) in formulas (4.2) are the ratios of the corresponding variables introduced in the preceding sections to the radius of inertia \u03c1; i.e., we assume that the radius of inertia \u03c1 is equal to unity. The variables z1 and z2 have the following physical meaning. These are small deviations of the points O1 and O2 from their equilibrium positions described in Sec. 3. Equations (4.1) and (4.2) and relations (1.1) with the use of the Lyapunov theorem permit solving the above stability and instability problem for any fixed v. Let us write out the characteristic polynomial of the linear system (4.1) as p4 + p3(a11\u03b31 + a22\u03b32) + p2(a11c1 + a22c2 + \u03b31\u03b32\u0394) + p(\u03b31c2 + \u03b32c1)\u0394 + c1c2\u0394 = 0, (4.3) where \u0394 = a11a22 \u2212 a12a21. Applying the Routh\u2013Hurwitz theorem to the polynomial (4.3), we arrive at the following result. Proposition 1. For the asymptotic stability of the equilibrium in the case of rectilinear and uniform motion of the vehicle, it is necessary and sufficient that the rolling friction coefficients f1 and f2 (and hence the velocity v of the vehicle translational motion determined by (1.1)) satisfy the inequalities a11\u03b31 + a22\u03b32 > 0, a11c1 + a22c2 + \u03b31\u03b32\u0394 > 0, \u0394 > 0, (4.4) (a11\u03b31 + a22\u03b32)[a11\u03b32c 2 1 + a22\u03b31c 2 2 + \u03b31\u03b32\u0394(\u03b31c2 + \u03b32c1)] \u2212 (\u03b31c2 + \u03b32c1)2\u0394 > 0, \u0394 = a11a22 \u2212 a12a21 = a(a + f2 \u2212 f1), a = a1 + a2. (4.5) The dimensionless parameters aij (i, j = 1, 2) are determined by formulas (4.2). Consider several particular cases where Proposition 1 applies. Proposition 2. If the system is completely symmetric, i.e., if \u03b3k = \u03b3 = 0, ck = c, fk = f, k = 1, 2, a1 = a2 = a0, (4.6) then inequalities (4.4) hold for each f ; i.e., the asymptotic stability of the vehicle equilibrium considered above is preserved for any wheel rolling friction coefficients and hence for any velocities of its translational motion. The proof is by a straightforward verification of inequalities (4.4) under conditions (4.6). In this case, it follows from (4.5) that \u0394 = a2 > 0 and the first two inequalities in (4.4) are satisfied, because they are reduced to the following ones: 2\u03b3(1 + a2 0) > 0, 2c(1 + a2 0) + \u03b32a2 > 0. The last inequality in (4.4) has the form c2(1 \u2212 a2 0) 2 + 4\u03b32c2a2 0(1 + a2 0) > 0. Thus, the stability conditions (4.4) are satisfied for any f \u2265 0, and the proof of Proposition 2 is complete. Proposition 3. Assume that all conditions (4.6) except for the last are satisfied; i.e., a1 =a2 (the case of incomplete symmetry; i.e., the car body center of mass is displaced). Then, for a1 < a2, the stability conditions (4.4) are satisfied for every f \u2265 0, and for a1 > a2, the stability is violated for f > f0 = 1 a1 \u2212 a2 [ 2 + a2 1 + a2 2 + \u03b5(a1 + a2)2 \u2212 \u221a \u03b52(a1 + a2)4 + 4(a1 + a2)2 ] > 0, \u03b5 = \u03b32 c . (4.7) Remark. The quantity f0 in inequality (4.7) can be arbitrarily small for some values of the system parameters. For example, as \u03b5 \u2192 0 (small damping or large rigidity of the wheel) and for a1a2 = 1 and a1 > 0, we have f0 = (a1 \u2212 1)(1 + a2 1)/(a1 + a2 1), which can be arbitrarily small as a1 \u2192 1. MECHANICS OF SOLIDS Vol. 46 No. 4 2011 The proof is by a straightforward verification. Under the conditions given in Proposition 3, the instability inequalities (4.4) become \u03b3[2 + a2 1 + a2 2 + (a2 \u2212 a1)f ] > 0, c[2 + a2 1 + a2 2 + (a2 \u2212 a1)f ] + \u03b32(a1 + a2)2 > 0, \u0394 = (a1 + a2)2 > 0, x2 + 2x\u03b5(a1 + a2)2 \u2212 4(a1 + a2)2 > 0, x = 2 + a2 1 + a2 2 + (a2 \u2212 a1)f. (4.8) For a2 > a1, these inequalities are necessarily satisfied for any f > 0. (For the first three inequalities in (4.8), this is obvious, and for the last, this follows from the inequality x > 2 + a2 1 + a2 2 > 2(a1 + a2).) But if a2 < a1, then the first inequality in (4.8) is violated for f > f1 = (2 + a2 1 + a2 2)/(a1 \u2212 a2), the second is violated for f > f2 > f1, the third is always satisfied, and the last is violated for f \u2208 [f0, f00], where f0 is given by the formula in (4.7) and f00 is given by the same formula with the plus sign at the radical. It is easily seen that 0 < f0 < f1 < f00, which implies that all inequalities in (4.8) are violated for f > f0. The proof of Proposition 3 is complete. Proposition 4. Let f2 = 0, \u03b30 = 0, and f1 = f > 0; i.e., assume that the fore (driving) wheel does not experience any rolling resistance and there is no dissipation in it. In this case, if a1a2 < 1, then the instability arises for f > a1, and if a1a2 > 1, then the instability arises for f > (a1a2 \u2212 1)/a2. Proposition 5. Let f1 = 0, \u03b31 = 0, and f2 = f > 0; i.e., assume that the rear (driven) wheel does not experience any rolling resistance and there is no dissipation in it. In this case, if a1a2 < 1, then the instability arises for f > (1 \u2212 a1a2)/a1, and if a1a2 > 1, then the system is stable for any f > 0. The proof of Propositions 4 and 5 is by a straightforward verification of inequalities (4.4) in Proposition 1 under the assumptions stated above. Let \u03b31 > 0, \u03b32 = 0, and \u03b32 > 0. Then the following assertion holds. Proposition 6. Under the above conditions, the asymptotic stability domain D1 in the plane of the parameters (f1, f2) is given by the following inequalities. If 0 < a1a2 < 1, then 0 < f2 < 1 \u2212 a1a2 a1 , 0 < f1, c1a1f1 \u2212 c2a2f1 < c0, f1 \u2212 f2 < a1 + a2. (4.9) If 1 < a1a2 < \u221e, then 0 < f1 < a1a2 \u2212 1 a2 , 0 < f2, c1a1f1 \u2212 c2a2f2 < c0, (4.10) c0 = c1(1 + a2 1) + c2(1 + a2 2). (4.11) The proof of Proposition 6 is by a straightforward verification of inequalities (4.4) under the restrictions given in the statement. Note that if one of the inequalities in conditions (4.10), (4.11) is not satisfied, then the vibrations under study are unstable. Now let \u03b31 = \u03b32 = 0; i.e., assume that there is no dissipation in the wheels. In this case, Eq. (4.3) is biquadratic, the stability of system (4.1) can be only nonasymptotic, and the conditions for such stability are equivalent to the conditions that this biquadratic equation has only pure imaginary roots. These considerations lead to the following result. Proposition 7. For \u03b31 = \u03b32 = 0, the stability domain D0 in the plane of parameters (f1, f2) is given by the following inequalities: f1 > 0, f2 > 0, c0 > c1a1f1 \u2212 c2a2f2, a1 + a2 > f1 \u2212 f2, [c0 \u2212 (c1a1f1 \u2212 c2a2f2)]2 > 4c1c2(a1 + a2)(a1 + a2 \u2212 f1 + f2). (4.12) The proof of Proposition 7 follows from well-known necessary and sufficient conditions for the existence of pure imaginary roots of a biquadratic equation. Proposition 8. The inclusion D1 \u2282 D0 holds. MECHANICS OF SOLIDS Vol. 46 No. 4 2011 The proof of Proposition 8 is by a straightforward substitution of the extreme values in the linear inequalities (4.9) or (4.10) into inequalities (4.12) and verification that the latter are satisfied. Remark. Proposition 8 means that this nonsymmetric addition of dissipation narrows down the stability domain in the plane of the parameters (f1, f2). Graphically (schematically), the domains D1 and D0 in various possible cases are shown in Figs. 2\u201315, where digit 1 corresponds to the line f1 \u2212 f2 = a1 + a2, digit 2 corresponds to the line c1a1f1 \u2212 c2a2f2 = c0 = c1(1 + a2 1) + c2(1 + a2 2), digit 3 corresponds to the line f2 = |1 \u2212 a1a2|/a1, digit 4 corresponds to the line f1 = |1 \u2212 a1a2|/a2, digit 5 corresponds to the line f1 \u2212 f2 = \u2212|1 \u2212 a1a2|(a1 + a2)/(a1a2), and digit 6 corresponds to the parabola (c0 \u2212 c1a1f1 + c2a2f2)2 \u2212 4c1c2(a1 + a2)(a1 + a2 \u2212 f1 + f2) = 0. Figures 2, 4, 6, 8, 10, 12, and 14 correspond to the case of \u03b31 = \u03b32 = 0, while Figs. 3, 5, 7, 9, 11, 13, and 15 correspond to the case of \u03b31 = 0, \u03b32 = 0 or \u03b31 = 0, \u03b32 = 0 for all possible distinct values of the parameters a1, a2, c1, and c2. Now consider the case of weighted wheels, m1 = 0 and m2 = 0. Since we assume that the car body mass is m = 1, the quantities m1 and m2 are the respective dimensionless ratios of the wheel masses to the car body mass. Simple calculations show that the form of the vibration equations (4.1) remains the same and the coefficients aij (i, j = 1, 2) are determined by the functions a11 = \u03c12 \u2212 (a1 + am2)(\u2212a1 + f1) \u03c12 0 , a12 = \u03c12 \u2212 (a1 + am2)(a2 + f2) \u03c12 0 , a21 = \u03c12 + (a2 + am1)(\u2212a1 + f1) \u03c12 0 , a22 = \u03c12 + (a2 + am1)(a2 + f2) \u03c12 0 , where \u03c12 0 = m0\u03c1 2 and m0 = 1 + m1 + m2. Then one can readily compute that \u0394 = a11a22 \u2212 a12a21 = a(a + f2 \u2212 f1)/\u03c12 0. Proposition 1 remains valid under the above variations in aij , i, j = 1, 2, and \u0394. In a similar way, one can restate Propositions 2\u20138. Equations (4.1) and (4.2) describing the vehicle vibrations under the action of the rolling friction forces admit the following analogies (interpretations) to well-known systems in mechanics. Each of the coefficient matrices multiplying zk and z\u2022k (k = 1, 2) in (4.1) with (4.2) taken into account consists of MECHANICS OF SOLIDS Vol. 46 No. 4 2011 two parts. The first part (it is said to be unperturbed) does not contain the friction coefficients f1 and f2. The second (perturbed) part depends on these coefficients linearly. As follows from the above derivation of the equations of motion, the origin of the perturbed part significantly depends on the following two external causes. One of them is the interaction between the deformable wheels and the rough road, which causes the rolling resistance torques Mk1 and Mk2; the second cause is the torque M1 from the engine (external energy source) to the driving wheels ensuring their (and the driven wheels) permanent rolling if there are resistance torques mentioned above. If these causes are absent, then the system motion occurs for f1 = f2 = 0, and Eqs. (4.1) contain only the first parts of the above-mentioned matrices for which the solutions have the property of (asymptotic) stability (a conservative system MECHANICS OF SOLIDS Vol. 46 No. 4 2011 under the action of some forces with complete or incomplete dissipation). But if f2 1 + f2 2 = 0, then these causes result in the appearance of additional terms in Eqs. (4.1), which, according to [6], can be interpreted as positional (circular) forces (for the corresponding terms with z1 and z2); gyroscopic and nonconservative (depending on the velocities) forces (for the corresponding terms with z\u20221 and z\u20222); conservative positional forces (for the corresponding terms with z1 and z2); and dissipative (Rayleigh) MECHANICS OF SOLIDS Vol. 46 No. 4 2011 MECHANICS OF SOLIDS Vol. 46 No. 4 2011 forces (for the corresponding terms with z\u20221 and z\u20222). In the general case, such a decomposition of forces was performed in [7]. Thus, the existence of the above-mentioned external causes leads to the appearance of additional forces of the described structure, whose destructive character is well known in mechanics of gyroscopic systems, structural mechanics (in particular, in mechanics of bridges), and other adjacent fields; it is described in the literature [8\u201310]. Even if there forces do not actually lead to the vehicle instability, they still significantly change the frequencies of its vibrations in the vertical plane, and this fact cannot be ignored when solving the corresponding problems. These facts were already considered earlier by the authors in the paper [11]." + ] + }, + { + "image_filename": "designv11_33_0002647_s12283-018-0276-z-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002647_s12283-018-0276-z-Figure1-1.png", + "caption": "Fig. 1 Set-up", + "texts": [ + " Once the true event is determined, the algorithm resets the initial conditions and surveys the next combination of candidate collisions. The next section will describe the event-driven approach in detail. The \u201cResults\u201d section determines the standard deviations in the launch conditions for different launch heights and the reasons behind the differences. The last section summarizes the findings. Let an event represent a trajectory that ends in a collision. The coordinates x and z of the ball during one step in a collision sequence are as follows (see Fig.\u00a01): The ball undergoes planar motion and the lateral coordinate y is neglected. At the beginning of the step, t0 is time, x0 and z0 are coordinates, vx0 and vz0 are velocity components, and g is the acceleration of gravity. The collision events with the ball satisfy the event constraints: (1)x = x0 + vx0(t \u2212 t0) z = z0 + vz0(t \u2212 t0) \u2212 g 2 (t \u2212 t0) 2. (2) Front hoop: (RC + R)2 = (x \u2212 RH) 2 + z2 Back hoop: (RC + R)2 = (x + RH) 2 + z2 Backboard: x = \u2212W, 0 \u2a7d z \u2a7d H Fictitious plane: \u2212 RH \u2a7d x \u2a7d RH z = \u2212R" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001677_iros.2011.6095023-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001677_iros.2011.6095023-Figure9-1.png", + "caption": "Fig. 9. An overview of the information processing in a multi-joint system with mechanically-coupled springs. The input signals multiplied with input weights win are applied to some springs around the upper joint as external forces. The output of the system is computed as a sum of all length of springs around the lower joint multiplied with output weights wout.", + "texts": [ + " If the passive part can emulate nonlinear combination and temporal integration of the indirect input signal, it can serve to generate an input-dependent signal to an adjoining part which should be actuated. This might allow us to regard passive elements in a robotic system as a local controller. Therefore, we investigated whether a spring network that receive the indirect input through mechanical connection can emulate nonlinear filters. We used a two-joint model in which N (= 16) springs are placed with equal spacing around the perimeter of the links connected by each joint (Fig 9). While the parameters of the springs around lower joint are randomly set at the same range as the ones on the previous experiments, the parameters of the springs around upper joint are set so that they are affected by input signals easily: k1 and d1 are randomly drawn from [1, 0 \u00d7 103, 1.0 \u00d7 105] and k3 and d3 are randomly drawn from [5.0 \u00d7 1011, 1.0 \u00d7 1012]. The same input signal shown by Eq. (6) was applied for randomly-chosen four springs around the upper joint with input weights. Since it turns out that the computational performance depends much on the distribution of the input weights, weight win,i for i-th spring was randomly drawn from: win,i \u2208 [0.0, 1.0] if N/2 \u2265 i win,i \u2208 [\u22121.0, 0.0] otherwise. The output of the system was computed as a sum of the weighted lengths of all springs around the lower joint (blue spring in Fig 9). Note that the input signal is not directly applied to springs that are used to compute the outputs of the system. The input information was transferred to the springs around the lower joint through the physical and mechanical connection. The passive movement of the springs was used to emulate the nonlinear filter. We ran 400 simulations, following the procedure described in section III-A. The MSEs between desired outputs and outputs of the adapted readout units were computed for the 2nd and 10th order system tasks" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002649_0142331218775490-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002649_0142331218775490-Figure3-1.png", + "caption": "Figure 3. Experiment equipment. (a) Test slewing bearing. (b) Test platform.", + "texts": [ + " (3) Perform iterative calculation: calculate and update the fuzzy partition matrix U (k + 1) as well as cluster center V (k + 1). When it meets the condition V (k + 1) V (k) e, the iteration stops. Then, obtain each fuzzy clustering center. (4) Recognize the life state: calculate the subjection to normal state mnormal as the degeneration index to recognize the life state of the slewing bearing. To verify the feasibility of the proposed method, the slewing bearing for the experiment is shown in Figure 3(a), and the main parameters are as follows: the average diameter is 730 mm; the ball diameter is 22.225 mm; the inner ring diameter is 621.6 mm; the outer ring diameter is 811 mm; the number of ball is 91; and the contact angle is 45 . The test slewing bearing was put on the home-made test platform for the full life test, which is shown in Figure 3(b). The parameters of the test platform can be found in Table 2. Also, the test platform is combined with mechanical system, hydraulic system, measurement system and control system. Moreover, the test platform is driven by the hydraulic motor to rotate the pinion. Also, the accompanying slewing bearing is rotating along with the pinion. Then, the accompanying slewing bearing drives the testing slewing bearing to rotate by bolts and switching ring. In order to accurately acquire the force of slewing bearing at different parts, four accelerometers intervals of 90 are arranged on the slewing bearing, which are shown in Figure 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003130_icdsp.2018.8631550-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003130_icdsp.2018.8631550-Figure4-1.png", + "caption": "Fig. 4 Experimental bench and bearing faults", + "texts": [ + " According to aforementioned theory and principles, the framework of proposed method is presented in the Fig. 3: III. EXPERIMENTAL STUDY: USED FOR LOW SPEED ROLLING BEARING FAULT SIGNALS For verifying the effectiveness of the proposed method, this paper states the application in low speed bearing fault diagnosis and analyzes the result of the experiment. This paper will validate by the analysis of three rotating machinery fault signal of outer race fault, inner race fault and roller fault and normal condition signal. As shown in Fig. 4, the experimental bench is for condition diagnosis test, it includes a servo motor and rotor system. The normal bearing, outer race fault, inner race fault and roller fault type is NU312. As shown in Fig. 4, three different types fault which is outer race fault, inner race fault, and the roller fault was artificially made by using electro discharge machining which fault width is 5.0 mm, depth is 0.03 mm. Each state original vibration signals was measured by accelerometer (PCB MA352A60, PCB Piezotronics Inc., New York, NY, USA) with 100,000 Hz sampling frequency. The accelerometer was fixed on vertical direction of the bearing. While obtaining the vibration signals, the speed of servo motor was 100 RPM" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003466_j.mechmachtheory.2019.06.017-Figure10-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003466_j.mechmachtheory.2019.06.017-Figure10-1.png", + "caption": "Fig. 10. Contact lines on the tooth surfaces of: (a) the face-gear; (b) the tapered involute shaper.", + "texts": [ + " Meshing equation between the face-gear and the tapered involute shaper is expressed by f s 2 = 0. v s (s 2 ) is the relative velocity and n s is the normal vector to the shaper surfaces s ; both of them are represented in coordinate system S s . A contact line of the face-gear drive can be obtained by application of Eq. (21) if the rotation angle \u03c8 s of the tapered involute shaper is given. The contact lines which entirely cover the tooth surfaces of the face-gear and the tapered involute shaper will be determined with a series of corresponding rotation angles \u03c8 s of the shaper ( Fig. 10 ). Due to the helix angle, the contact lines are asymmetric on different sides of the tooth surfaces. Generally, there are four limiting conditions for the generated tooth surfaces of the face-gear, i.e. undercutting, involute interference, pointing, and fillet intersection, as Fig. 11 shows. Singular points on the face-gear surfaces should be eliminated in order to avoid undercutting. The singularity on 2 occurs when the following equation is satisfied [30] : v ( 2 ) r = v ( s ) r + v ( s 2 ) = 0 (25) where v r ( 2 ) and v r (s) are the velocities of the contact point on the face-gear tooth surfaces 2 and the shaper surfaces s , respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003921_b978-0-12-803581-8.11722-9-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003921_b978-0-12-803581-8.11722-9-Figure6-1.png", + "caption": "Fig. 6 Schematic of a drivetrain with a hydraulic clutch.", + "texts": [ + " The ER fluid flowing between the rotor blades, connected alternatingly with the terminals of the electric high voltage power supply, is exposed to the electric field which causes an increase in shear stresses in the ER fluid and an increase in resistance of the fluid flow, consequently decreasing the transferred torque. One of the parts of the design process of the clutches and brakes with ER fluid is the creation of mathematical models allowing to determine how different construction parameters influence the performance, as well as allowing to optimize the construction. The basic equations of the mathematical model of a drivetrain of a machine with a hydraulic clutch is created after dividing the drivetrain into two parts connected with a working fluid, Fig. 6. For simplification purposes, the movement resistance of bearings and sealings are omitted, and it is assumed that the shafts are inflexible. With these assumptions, the equations can be formed as follows: Me \u00bcM\u00fe J1 do1 dt Mr \u00bcM J2 do2 dt \u00f02\u00de where: Me \u2013 torque of the drive engine, Mr \u2013 torque of the machine\u2019s movement resistance, M \u2013 torque transmitted by the clutch, t \u2013 time, J1 \u2013mass moment of inertia for rotating masses connected to input shaft of the clutch, J2 \u2013 mass moment of inertia for rotating masses connected to output shaft of the clutch. The Eq. (2) describing the unsteady motion of the transmission system of the machine containing the hydraulic clutch are nonlinear differential first-order equations, solving which demands practical use of numerical methods. According to the schematic shown in Fig. 6 the torques Me and Mr affect the hydraulic clutch\u2019s shafts from the outside, while the torque M works inside the clutch. Depending on the modeling aim, the outer moments Me and Mr which extort the unsteady motion are established as time functions t or shafts\u2019 angular velocity functions o1 or o2. However, the value of the inner torque M depends on the build of the clutch and is established as an angular velocity function of shafts o1 or o2. If the drivetrain in question contains a hydraulic clutch with ER fluid, the cause of changes resulting in movement can be (apart from changes of outer torques Me and Mr) changes in torque M which is additionally dependent on voltage U supplying the electrodes of the clutch" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002336_iccas.2017.8204221-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002336_iccas.2017.8204221-Figure5-1.png", + "caption": "Fig. 5 Apperance of apparatus", + "texts": [ + " Note that, in order to realize compact and simple mechanism, we apply one artificial muscle for each of right and left arm-lower back assistance independently. The compressive force generated by the McKibben artificial muscle rotates the shoulder pulley and the leg one through the wire. And then the wearer's arms and lumbar assistance for lifting loads are realized via the arm frame and the leg one connected to each of the shoulder pulley and the leg one. The wearer balances the assistance power for the arm and the waist by himself. The diameter of the shoulder pulley and the leg one were empirically set at 50 mm. Fig.5 shows the entire figure and axes of the arm and lower back assistant muscle suit. The axes (i) and (ii) are perpendicular to the longitudinal direction of the arm. These axes are passive ones and correspond to elbow and shoulder position respectively. The axes (iii), (iv) and (v) are actively actuated by the artificial muscle, whereas (iv) and (v) are guides of the same wire. The distance between each axis was determined empirically. The main body is worn on the body by the shoulder belt (a) and the waist belt (b)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002131_metroaerospace.2017.7999605-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002131_metroaerospace.2017.7999605-Figure8-1.png", + "caption": "Fig. 8. Simplified representation of the load cell-chain system.", + "texts": [ + " In the following subsections the measurements provided by \u201cDronesBench\u201d are described. In this subsection, the mathematical model used for evaluating the module of thrust force given by the VTOL-RPAS is reported. In this analysis, the coordinate system has been considered according to Fig. 7. In particular, the chosen origin corresponds to the reference plane center and the points S1, S2, and S3 represent the load cell coordinates. By using the three load cells, the system allows to measure the forces applied by RPAS to the accommodation plane (see Fig. 8). Each load cell measures the force F s i . The corresponding reaction force of the chain Ri can be evaluated as follows: Ri = F s i cos(\u03b1i) (2) where \u03b1i corresponds to the elevation of vector distance \u2212\u2212\u2192 SiTi respect to its projection on the plane defined by the coordinates of the three load cells S1, S2, and S3. This plane is given by: z = Ax+By + C (3) where A, B and C are estimated using the distance measurements provided by SHARP sensors and the pitch and roll angles evaluated by GY801 module" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002149_gt2017-63495-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002149_gt2017-63495-Figure1-1.png", + "caption": "Figure 1 Rotor-bearing configuration and the coordinate system", + "texts": [ + " The other assumptions are, (1) two bearings are concentric, and (2) entire rotor is a rigid body with four degrees of freedom (4 DOF) motions. Orbit simulations are carried out by timeintegration of the equations of the rotor 4 DOF motions, equation for bump foil deflection, and transient Reynolds equation for each bearing simultaneously. Similar approach using 4 DOF rigid rotor model was used in the other previous studies [24-26]. The 4 DOF motions of the rotor are translation motions along the X and Y directions, and rotational motions about X and Y axes ( and , respectively). Figure 1 depicts the rotorbearing system configuration and the global coordinate system. The equations of motions for the rotor are: _ _ _ _ _ _ _ _ R brg X imb X R R brg Y imb Y T P brg imb T P brg imb m X F F m g m Y F F I I M M I I M M (1) , where Rm is the rotor mass, ,p TI I are the polar and transverse moments of inertia of the rotor. _ / _ /,brg X Y imb X YF F are the bearing reaction forces and the imbalance forces, and _ / _ /,brg imbM M are the reaction moments from the bearings and external moments from imbalance forces, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001046_wcica.2012.6358389-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001046_wcica.2012.6358389-Figure4-1.png", + "caption": "Fig. 4: Water pressure distribution along the planing hull Later, Garme changed the reduction length into a nondimensional form:", + "texts": [ + " Shuford recommended a factor of one-half to obtain a correct buoyancy value for each cross section in his work on steady-state planing. Total water forces and moment The total water force acting on the flying boat is obtained by integrating these 2D hydrodynamic forces estimated by Eqs(2) -(21) over the wetted length of the flying boat ( ) as showed in Eq kL (22). cos k w k k w s cfd bL bf b b L w s cfd b b b bL L N f f dx F f dx bM f f x dx f x dx (22) Water pressure distribution along the bottom of planing boats is showed in Fig 4. However, as the 2D strip theory does not consider the influence of water separation at the transom of the flying boat, the water pressure estimated by this method has no such a similar distribution curve, which makes the trimming angles are bigger than actual values. In order to solve this problem, Garme formulated a correction to the 2-dimensional strip theory basing on the assumption that the pressure at the dry stern is atmospheric pressure. It is expressed by Eq(23): 2.5tanhtr b bstrC x a x (23) where bstrx = the hull stern position in bX ; is a reduction length" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000970_tme.1963.4323041-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000970_tme.1963.4323041-Figure2-1.png", + "caption": "Fig. 2.", + "texts": [ + " The position error component AR is a straight-line continuation of the great circle path from launch point to target and since AR is small compared with the missile flight path it can be considered a continuation of the great circle path. From Fig. 1, it can be seen that 3R= AR: where a= earth's radius y= the range angle S=actual location of the shipboard navigator Sc=the computed location of the shipboard navi- gator T=the target IP=the missile impact point. The position error component AL is orthogonal to the great circle path from launch point to target. This is (2) shown in Fig. 2. The terms Sc, S, a, IP, T, and y retain their same definitions, and the following additional terms are defined: Great Circle Path 1 Correct firing path heading for actual location if missile firing is from computed location. Great Circle Path 2-Correct firing path heading for computed location if missile firing is from computed location. Great Circle Path 3 Correct firing path heading for actual location if missile firing is from actual location. Great Circle Path 4 -Correct firing path heading for computed location if missile firing is from actual location", + " Using small angle approximations (assuming S, Sc, and T lie in the same plane) which are certainly valid here, 3L2j = 5L1j . This implies Aa1-Aa2 which is valid under our assumption and the equivalence of 8L1 and UL2 follows from the law of sines. At this point, it is important to note that Aa contains the error in azimuth firing angle arising from a system azimuth error. This system azimuth error is expressible as 4,=4/k -AA sin 0 and if we assume there is no angular difference between the computer and platform axes, f= -AA sin 0. From our assumption of small angles and applying the law of sines to Fig. 2, sin@1 1 _- -=1 AL sin 900 a AL a 46 January Benso and duPlessis: Navigation System Position and Azimuth Errors on Radial Miss With f11 defined, the lateral miss distance can be obtained sin /1 1 1 SL2 sin (90-y) cos y a =L2 a sin /1 cos y= AL cos y. However, as 5L2 is equal in magnitude and diametrically opposite to the actual impact point, lateral miss distance as a function of lateral position error must be expressed as 3L = - AL cos y. Finally, lateral miss can be caused by an azimuth error in the autonavigator", + " Also, the computation time for the digital computer program is assumed to be zero, which is a close approximation to the actual case. Fig. 1 shows a block diagram of the system under consideration. The digital computer program is simply a gain factor T/R, where T is the sampling interval and R is the distance to earth center. The justification of this computer program can be easily seen if the discrete summer block is combined with the digital computer block to form the pulse transfer function denoted by D(Z) in Fig. 2. D(Z) represents a rectangular integration of the recursion form CN = CN-1+ (T/R)RN where CN and RN are general output and input symbols, respectively. It is obvious that D(Z) should represent an integra- 56 January" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001385_isci.2012.6222678-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001385_isci.2012.6222678-Figure4-1.png", + "caption": "Fig 4: Block Diagram of MFS", + "texts": [ + " Angular misalignment is sometimes referred as gap or face, is the difference in the slope of one shaft, and usually occurs in the moveable machine, as compared to the slope of the shaft of the other machine, usually in the stationary machine. In this paper, various tests were carried out on a machine fault simulator with fluid film journal bearings to observe both phenomena. In this work, we use data for both the shaft displacement from bearing as well as current samples. Experiment is conducted on the machine fault simulator is shown in fig 3. The experimental set \u2013up is placed in Dynamics and Machinery lab at Indian Institute of Technology Patna. Block diagram for machine fault simulator are also shown in fig. 4. Total experimental setup comprises of three major subsystems: Machine fault simulator, NV gate Oros data acquisition system and a computer, as shown in fig 5. Fig.5: Side view of MFS with Complete Experimental Setup Machine Fault Simulator with a 3-phase, 0.75 HP induction motor was used for experiment. The induction motor with a wiring enclosure are attached on the left side of the machine fault simulator and that wiring enclosure of the motor support permits access to the power supply and motor supply leads" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002321_j.compstruc.2017.11.003-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002321_j.compstruc.2017.11.003-Figure1-1.png", + "caption": "Fig. 1. Industrial stator vane model.", + "texts": [ + " Mistuning also implies a noticeable effect on the vibratory behavior of bladed disks by concentrating a large proportion of the energy on a reduced number of blades [10,11,35], thus increasing the vibratory level [29]. In this paper, the stator vanes of a high-pressure compressor will be considered; their role is to straighten the air flow arriving from the rotors located upstream. To reduce mass and simplify both manufacturing and assembly, some aircraft engine manufacturers have decided to segment stator vanes into several multiblade clusters (see Fig. 1). This segmentation induces a loss of cyclic symmetry properties even when mistuning is not taken into account. This revised architecture implies a high modal density area that is extremely sensitive to mistuning, therefore making the vibratory response of a stator vane cluster very difficult to predict [30]. Pichot and Laxalde [28,18] focused on mistuning identification through relevant experimentation; this paper however will propose an alternative approach based on the numerical mistuning characterization", + " The purpose of this paper is to devise a strategy that allows predicting the vibratory behavior of a mistuned stator vane cluster by taking uncertainties into account. In the first part, the dynamic behavior of a tuned stator vane sector will be detailed through establishing a model that takes uncertainties into account. The second part will focus on theoretical aspects of the implemented method, and the third and fourth parts will apply the methodology to an academic model and an industrial model, respectively. Fig. 1 illustrates the stator vane cluster of a high-pressure compressor; it is composed of eleven blades linked by both inner and outer platforms. A modal analysis has been conducted on the tuned case. Fig. 2 presents the evolution in the normalized eigenfrequencies vs. the mode number. Results have been normalized with respect to the 20th eigenfrequency. Fig. 2 indicates a high modal density area, with 15 modes being clustered within a 20% frequency range. This area is primarily composed of first flexural (1F) and torsional (1T) blade modes (Fig", + " This statement means that Ts\u00f0a\u00de differs depending on the mistuning. Moreover, this matrix would need to be computed for all deterministic computations, which could prove to be a critical factor with respect to computational costs. It is thus assumed that transform matrix remains constant regardless of stator vane mistuning: Ts\u00f0a\u00de \u00bc Ts\u00f00\u00de \u00bc Ts: \u00f026\u00de Thus, reduced matrices are defined by: Mr \u00bc Tt sMCBTs; \u00f027\u00de Kr\u00f0a\u00de \u00bc Tt sKCB\u00f0a\u00deTs: \u00f028\u00de In order to validate this assumption, 50 eigenfrequencies of the considered stator vane (Fig. 1) were initially computed for 100 mistuning cases, without exceeding a 5% variability and in maintaining the same tuned case transform matrix Ts for all detuned trials. The double modal synthesis retains 200 internal platform modes, 50 internal blade modes and 1000 static modes, in dividing the basis size by roughly 20 and by more than 330 when compared to the classical Craig-Bampton method and original model, respectively. These results have been compared to those of both the Samcef industrial code (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000776_msec2010-34325-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000776_msec2010-34325-Figure4-1.png", + "caption": "FIGURE 4. NORMAL AND TANGENTIAL FORCES ON THE CUTTER", + "texts": [ + " F Kbh= , (4) where F is the cutting force, b is the axial depth of cut, and h is the chip thickness. These parameters are shown in Figure 3, which shows also the radial depth of cut (DOCr). The chip thickness varies with time and can be approximated using equation (5). ( )sinth f \u03b8= (5) where ft is the feed per tooth and \u03b8 is the angle of rotation of cutting edge. Equation (1) is expanded to normal and tangential components of the cutting force, Fn and Ft respectively, by substituting kn and kt for K. Figure 4 illustrates the normal and tangential cutting forces on the cutter. The accuracy of the model is improved by the addition of edge effects to the basic equation. Edge effects are due to the tool rubbing the workpiece and are only dependent on the depth of cut. These effects include friction stresses at the tool/workpiece interface and plastic deformation occurring in the shear zone. The improved model equations are shown below. n n neF k bh k b= + (6) t t teF k bh k b= + (7) where k is the specific cutting force and the subscript n denotes \u2018normal\u2019, t denotes \u2018tangential\u2019, and e denotes \u2018edge effects\u2019" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001649_jfm.2012.1-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001649_jfm.2012.1-Figure1-1.png", + "caption": "FIGURE 1. (a) The mid-line of the flexible swimmer is represented by curve C , parameterised by the Lagrangian coordinate s and time t. The body moves with speed, \u2212W (t), relative to the fluid. \u03b8(t) is the tangent angle made by the body at P\u2032, the reference point on C at a Lagrangian distance so from the nose. The body frame of reference {ex(t), ey(t)} is defined at the fish centre of mass, P, at position vector Ro(t) with respect to the fixed origin, O. (b) A close-up view near P and P\u2032 indicates the prescribed deflection, x1(s, t), of a point s on the curve defined with respect to P\u2032, and with position vector X(s, t). At s the local tangential basis, {\u03c4 (s, t),n(s, t)}, makes an angle \u03c7(s, t) with the fixed basis.", + "texts": [ + " Based on observations (Gray 1933), a test deflection function is proposed and applied to analyse the turning response to a lateral bend. In \u00a7 5 we use these models to find an optimal body depth from a manoeuvrability perspective. We then use the optimal configuration to examine the role of different initial conditions and the prior shed wake in enhancing the turning response of a self-propelled fish. In this section we consider the model used to represent the fish as it swims in a quiescent fluid. Curve C (s, t), in the sketch in figure 1, represents the fish, parameterised by the Lagrangian coordinate s and time t. The laboratory frame of reference, { i\u0302, j\u0302, k\u0302}, is fixed in space at point O and the body frame, {ex, ey, ez}, is fixed on the centre of mass of the fish (referred to as point P in the sketch). P\u2032 is a reference point on the body a Lagrangian distance, so, from the nose with respect to which the flexible-body deflections are defined. P\u2032 is defined so that when the fish centre-line is straight, P and P\u2032 coincide; ex is tangent to the body at P\u2032 making an angle \u03b8 with i\u0302", + " The fluid velocity relative to the body is given as W (t) = \u2212Vo(t) and makes an angle \u03c6 with i\u0302 where \u03c6(t) = tan\u22121(Vyo/Vxo) is the direction of motion. The position vector of the centre of mass, with respect to the origin O, is Ro(t)= Rx(t) i\u0302+ Ry(t) j\u0302 = \u222b t 0 (Vxo(t) i\u0302+ Vyo(t) j\u0302) dt. (2.7) The position vector of an arbitrary point s on C with respect to fixed point O is X(s, t)= X i\u0302+ Y j\u0302 = Ro(t)+ r(s, t), (2.8) where r, the position vector with respect to the centre of mass P, is defined in (2.10). With respect to figure 1 recall that P\u2032 and P are coincident on a rigid body. 2.2.2. Kinematics of the combined flexible- and rigid-body motion Determining the instantaneous position of the centre of mass is non-trivial for flexible, self-propelled swimmers. We must distinguish between the reference point on the body, P\u2032, (at position vector Xo, see figure 1), and the centre of mass, P (at Ro). For a flexible body the centre of mass will shift as the body undergoes unsteady, flexible deflections; it will not in general lie on C . Furthermore, the flexible motion is prescribed by defining the form of C , so we need an independent reference point, P\u2032, at s= so on C . The position vector of P with respect to P\u2032 is ro(t)= \u222b l 0 mb(s)x1(s, t) ds\u222b l 0 mb(s) ds , (2.9) where x1(s, t) = x1(s, t)ex + y1(s, t)ey is the vector distance of s defined with respect to P\u2032", + "1 we compare results from Model 1 with Kambe\u2019s results. The impulse response is then applied to a rectangular-shaped fish and the results from Models 1 and 2 are compared in \u00a7 4.2. In \u00a7 4.3 a flexible deflection is prescribed to mimic the C-turn action of fish and we compare the response of both models. 4.1. Rigid-body response to an impulse: elliptical fish Kambe (1978) considered a fish with an ellipse-like profile such that the longitudinal variation of the span about so (reference point P\u2032, see figure 1) is given by d(s)= do ( 1\u2212 ( s\u2212 so so )2 ) , (4.1) where do is the span at so. Here we define the reference point at mid-length (so = l/2) and so the shape is symmetric about ez = 0. In Kambe\u2019s model the action of the tail was modelled as a point force at s= l, implemented by applying an impulsive input to \u03b8\u0307 and R\u0307y at time, t = 0. The rigid-body response of the fish was then calculated. Thus initial conditions, (2.19) and (2.20), are Rx(0)= 0, Ry(0)= 0, \u03b8(0)= 0, (4.2) R\u0307x(0)=\u2212Uo, R\u0307y(0)= Io, \u03b8\u0307 (0)= 7Io", + " Finally, in this work we exclusively concentrate on the mechanical aspects of manoeuvres in fish; we have not accounted for biological aspects such as muscle action of the swimmer. We are currently pursuing this aspect following Wu (1971a), Cheng et al. (1998) and Pedley & Hill (1999). Acknowledgements The authors wish to acknowledge DAMTP, University of Cambridge and Fitzwilliam College for financial support of this work, and the anonymous referees for their valuable suggestions and feedback. In \u00a7 2.2.2 we outlined the solution formulation. Here we present the algorithm (also refer to figure 1). (a) At time instant t prescribe curve C , constrained by the length-preserving criterion, (2.11). This establishes position vector x1(s, t) of an arbitrary point on the body relative to reference point, P\u2032. (b) From (2.9) compute the position vector of the centre of mass, P, relative to P\u2032. (c) From (2.10) determine the position vector, r, of all points on the body with respect to P. (d) Determine the increment in position vector of the centre of mass, Ro(t), and orientation, \u03b8(t), over the previous time step by solving equations of motion, (2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000289_vppc.2009.5289719-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000289_vppc.2009.5289719-Figure4-1.png", + "caption": "Figure 4. Size of the prototype FRID EV used for evaluating the simulator.", + "texts": [ + " Dx and Dy are \u03bcij\u22c5(\u03c3xij/\u03c3ij) and \u03bcij\u22c5(\u03c3 yij/\u03c3ij), respectively. D. Validity of the Simulator for Vehicle Dynamics Analysis at the Time of Failure Before analyzing various behaviors occurring at the time of failure, it is necessary to examine whether the simulator obtained through the above procedures can reproduce operation of an actual vehicle. Here, the effectiveness of the simulator is checked using the front- and rear-wheel torques obtained when a prototype FRID EV with specifications shown in Fig. 4 and Table II actually experienced failure. Fig. 5 (a) shows the front and rear torques when the FRID EV failed during the rectilinear-propagation running on a dry road (friction coefficient \u03bc = 0.75) at 60 km/h. The simulator (KT of (11) is 250 N/mm; and \u03c1, Cd and A of (1) are 1.206 kg/m3, 0.3431 and 2.4 m2, respectively) with the same specifications as the FRID EV is created. The failure torque signals of Fig. (a) are input into the simulator, and the body speed is simulated. Fig. 5(b) shows the same vehicle speed as in an experiment is also obtained by the simulator" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002317_ecce.2017.8096018-Figure16-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002317_ecce.2017.8096018-Figure16-1.png", + "caption": "Fig. 16. Illustations of radial and torsional modes (m,n).", + "texts": [ + " For the sake of simplicity, the stator is here modelled without frame, in clamped-clamped boundary conditions (null displacement at both ends), and the stator windings are modelled as a distributed mass in the slots. The number of axial layers is 8 which makes it possible to catch longitudinal modes of orders up to . The stator lamination is M400-50A for which the density is 7650 kg/m3 and the orthotropic elastic properties are listed in Table II. After the numerical resolution, 500 modes are found under 1500 Hz but only a few of them will potentially have an influence on the radial vibration level. Some of the main structural modes of the stator are illustrated in Fig. 16, and their associated natural frequencies are given in Table III. The natural frequencies are very low due to the large size of the machine and can be excited by magnetic force harmonics due to the interaction between rotor salient poles and stator slots, saturation and PWM effects. Each mode is characterized by the nature of the deflection \u2013 either radial or torsional \u2013 and the pair ( ), where is the circumferential order and is the axial order of the deflection. The radial modes directly deflect the stator yoke while torsional modes may cause radial vibrations due to tooth bending modes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001673_sav-2010-0501-Figure11-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001673_sav-2010-0501-Figure11-1.png", + "caption": "Fig. 11. Optimization process: shape changes in the automobile part.", + "texts": [ + " The first natural frequency mode varies from 416 Hz to 439 Hz and the second mode from 1754 Hz to 1883 Hz, moving away from the frequency range of the PSD (Fig. 8). Predicted fatigue life of the part is about 5 \u00d7 107 to 108. Two other stages of material removal were carried out to reduce weight. Figures 9 and 10 show second and third stage optimization results. The part\u2019s mass decreases without affecting either its fatigue life or its first mode significantly. The mass decreases from 3.5 kg to 3.1 kg, which corresponds to a 11.42% weight loss. All critical elements are located in the embedded area. Figure 11 shows the geometrical evolution of the part during the optimization process. Numerical results are summarized in Table 4. The results of the optimization study are comparable to Haiba et al.\u2019s [1] work on weight optimization of a simple mechanical part. However, their fatigue study is based primarily on simple experimental method of the Wo\u0308hler curve extrapolation in the calculation of fatigue life. They took into account only one parameter: stress variation. Our approach takes into account two parameters: stress and strain, using the multiaxial criterion of the strain energy density equivalent to the uniaxial case" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001670_0142331211408638-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001670_0142331211408638-Figure1-1.png", + "caption": "Figure 1. PVTOL aircraft.", + "texts": [ + " In a general perspective, controller design of these vehicles is not an easy task because they have quite complex dynamics when the flexibility of wings, the internal dynamics of motors and many other uncertain variables are considered. However, simplified dynamic models having less states and control inputs than the real systems without sacrificing from their phenomena are commonly used for controller design (Hauser et al. 1992; Fantoni and Lozano 2002). In this paper also, global stabilization problem of PVTOL aircraft (Figure 1) whose normalized and simplified dynamic model is given by (2) is considered. Among the others, studies on PVTOL aircraft are often encountered in the literature. A nonlinear control of the PVTOL via approximate input/output linearization is introduced by Hauser et al. (1992). Stabilization of the PVTOL aircraft is examined as an example of the interconnection and damping assignment passivity-based control (IDA-PBC) method by shaping the kinetic and potential energies of the port-Hamiltonian system (A\u0306costa et al", + " In order to achieve that, a candidate Lyapunov function is constructed having a negative semi-definite derivative obtained by means of smooth and static feedback signals. Then, global asymptotic stability is proven with invoking La Salle\u2019s invariance set principle. This paper is organized as following. Section 2 introduces the dynamic equations of motion of the PVTOL aircraft. A controller design procedure and stability analysis are discussed in Section 3. Following, in Section 4, an illustrative simulation study is presented. Lastly, Section 5 concludes the paper. The normalized dynamic equations of motion of PVTOL aircraft (Figure 1) can be given by \u20acq1 \u00bc u1 sin\u00f0q3\u00de \u00fe Eu2 cos\u00f0q3\u00de, \u20acq2 \u00bc g\u00fe u1 cos\u00f0q3\u00de \u00fe Eu2 sin\u00f0q3\u00de, \u20acq3 \u00bc u2, \u00f01\u00de where q1 and q2 are horizontal and vertical positions of the aircraft\u2019s center of mass, q3 is the roll angle with the horizon, u1 and u2 denoting control inputs are the thrust directed out the bottom and rolling moment respectively and g is the gravitational acceleration. Here e is the coupling between rolling moment and the lateral acceleration (see Fliess et al. (1999) for the details of the model)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001127_j.ijepes.2012.08.004-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001127_j.ijepes.2012.08.004-Figure7-1.png", + "caption": "Fig. 7. Power circuit for a line reactor reduced voltage starter.", + "texts": [ + " Copper losses in the winding are determined from the length as well as the area required for the conductors in the slot. Iron loss in stator core per unit volume = 3.88708 10 5 W/mm3. Iron loss in stator teeth per unit volume = 3.92352 10 5 W/mm3 7.3.1. Stator copper A line reactor reduced voltage starter is used for comparatively high voltage or high current installations where resistors become bulky and heat dissipation is a problem. A reactor starter is preferred where emphasis is on getting a lower starting torque than on reduction in starting current. The power circuit for the starter is shown in Fig. 7. Here 50% tapping has been used. When contactor S is energized the whole of the reactor winding gets connected in circuit and thus 50% of the supply voltage gets applied across the motor terminals. Next, the run contactor R is energized then S is also energized which bypasses the reactor and connect the motor directly to the supply. In reactor starting of the induction motor, the equivalent circuit of which is shown in Fig. 6, we are interested to calculate the temperature distribution in the stator during the starting period" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001329_05698196408972042-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001329_05698196408972042-Figure3-1.png", + "caption": "FIG. 3. Backing shaft assembly", + "texts": [ + " i radial deflection of ring E;, EI error functi ons EI error A roller crowning and guidance facter v life exponent factor 0, k2 > 0, k3 > 0, \u03b31 > 0, \u03b32 > 0, \u03bd20 and \u03bd30 are obtained via sliding mode differentiator. The simulation figures are presented in Figures 2\u20134. Curves of output, reference signal and tracking error are presented in Figure 2, where the effectiveness of our proposed method is demonstrated. The control input signal is given in Figure 3. In Figure 4, the quantised control input is shown. 5. Conclusions In this paper, the problem of asymptotical tracking control for nonlinear system with unknown function and input quantisation has been investigated. The tracking controller has been proposed to guarantee that all the signals were uniformly ultimately bounded and the tracking error converged to zero asymptotically. An application example has been given to demonstrate the effectiveness of the proposed scheme" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003884_itsc.2019.8917468-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003884_itsc.2019.8917468-Figure1-1.png", + "caption": "Fig. 1 Tracked vehicle kinematic relationship, the left is kinematic model with sliding steering characteristics and the right is kinematic model with no sliding steering characteristics. When turning on the XOY plane, the vehicle can be regarded as a rigid body and its motion around its instantaneous centers of rotation co . lo and ro are instantaneous center of both sides of the track", + "texts": [ + " And it can also determine the relationship between control sequence under different weight coefficients and the control operation of the experienced driver. A weight coefficients design method is provided for the application of the MPC algorithm on tracked vehicles. The structure of the remaining articles is as follows: the \u2161 section is about the theoretical analysis and the objective function. In the \u2162 section, the simulation experiment and the real vehicle test are completed. The \u2163section is data analysis. The \u2164section is the conclusion and summary. We have established the following kinematics model of the tracked vehicle [8], as shown in Fig 1. The algorithm takes advantage of the global reference frame G (X, Y, W) and the vehicle reference frame L (x, y,\u03c9 ). Assuming that the tracked vehicle is moving on a 2D plane and the center of mass of the vehicle coincides with its geometric center. Fig. 1 shows the kinematic parameters and geometric relationships during the turning about two kind of the kinematic models with sliding steering characteristics and with no sliding steering characteristics. The frame G is XOY and frame L is xoy. Vehicle state vector x [ xv , yv , zw ] are forward, lateral and angular velocity, respectively. M and N are points on surfaces of the tracks that are in contact with terrain. Variables Mqv and Nqv are vehicle body velocities at points M and N, l sv and r sv are tracks\u2019 rolling speeds relative to the vehicle body" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000215_cefc.2010.5481322-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000215_cefc.2010.5481322-Figure3-1.png", + "caption": "Fig. 3. Contours of the density of eddy current density (Phase angle = 90 deg)", + "texts": [], + "surrounding_texts": [ + "Fig. 1 shows an axial-type magnetic gear used in this study, which mainly consists of a high-speed rotor, a low-speed rotor, and stationary pole pieces. A high-speed rotor generates magnetic harmonics in the air gap between stationary pole pieces and the low-speed rotor. While a high-speed rotor rotates, a low-speed rotor rotates in accordance with the gear ratio." + ] + }, + { + "image_filename": "designv11_33_0001627_1.4001668-Figure12-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001627_1.4001668-Figure12-1.png", + "caption": "Fig. 12 Setting errors of the wedge the lead measurement plane of a inclination direction error", + "texts": [ + "1 Mathematical Model of Tooth Lead Measurement Usng Wedge Artifact Considering Setting Error. Equation 5 for he lead deviation does not take the setting error of the wedge rtifact into consideration. The mathematical model with the seting error is constructed and the equation for the lead deviation onsidering the setting error is derived mathematically in this secion. The setting error of the wedge artifact is expressed using two arameters, the actual inclination angle and the maximumnclination direction error , as shown in Fig. 12. The former enotes the actual inclination angle of the lead measurement plane f the wedge artifact, where the inclination error of the plane due o the setting error is included. The latter denotes the angle deviaion between the sensing direction line of the displacement sensor nd the maximum-inclination direction of the lead measurement lane. Figure 13 shows the state where the probe sphere is in ontact with the lead measurement plane S1 of the wedge artifact nder the initial state of the lead measurement, and the state after he wedge artifact has rotated by " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001138_j.mechmachtheory.2011.07.015-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001138_j.mechmachtheory.2011.07.015-Figure3-1.png", + "caption": "Fig. 3. ZN worm-type hob cutter with varying pitch lines represented by parameter dx.", + "texts": [ + " The results of this study can be applied to further investigations on tooth contact analysis, kinematic errors, contact paths and contact patterns of the semi RA, full RA and standard proportional tooth worm gear drives. Fig. 1 shows three types of worm gear drive meshing, semi RA, full RA and standard proportional tooth worm gear drives, with double-depth teeth at the same standard center distance. Fig. 1(a) shows the worm gear drive with standard proportional tooth Nomenclatures bx,bn width of hob cutter at varying pitch line, respectively (Fig. 3) C1 center distance of hob cutter and RA worm gear (Fig. 5) dx distance measured from the middle of hob cutter tooth height to the varying pitch line (Figs. 2 and 3(b)) h straight-lined edge height of hob cutter cutting blade (Fig. 3(b)) ht whole cutting blade height of hob cutter (Fig. 3(b)) Lij coordinate transformation matrix transforming from coordinate system Sj to Si (Eqs. (26), (27), (29), and (31)) l1 surface parameter of hob cutter (Fig. 3(b)) Mij homogeneous coordinate transformation matrix transforming from coordinate system Sj to Si (Eqs. (8), (9), (18), and (19)) mn normal modulus (Fig. 2) mx axial modulus (Figs. 2 and 7), and mx=mn/cos \u03bb1 m21 angular velocity ratio of hob cutter to RA worm gear (Eqs. (28), (33) and (34)) N1,N1 (c) normal vectors of hob cutter tooth surface (Eqs. (12) and (15)) Nx1, Ny1, Nz1 components of the normal vector expressed in coordinate system S1 (Eqs. (34)) p1 lead-per-radian revolution of hob cutter blade surface (Fig. 4) R1, R2 position vectors of hob cutter and RA worm gear, respectively ro, r1, rf outside radius, pitch radius and root radius of hob cutter, respectively (Fig. 3(c)) r2 pitch radius of RA worm gear rc circular tip radius of hob cutter (Fig. 3(b)) rt design parameter of hob cutter (Fig. 3) Si, Sj reference and rotational coordinate systems (i= f, g, p and j=c, 1, 2, 3 ) T1, T2 number of teeth of hob cutter and RA worm gear, respectively tc, tt transverse chordal thicknesses at pitch circle and throat circle of RA worm gear, respectively (Fig. 11) V12 (1) relative velocity vector of hob cutter and RA worm gear expressed in coordinate system S1 (Eqs. (22) and (33)) Vi (1) velocity vectors of hob cutter and RA worm gear (i =1, 2) expressed in coordinate system S1 (Eqs. (22)) \u03b11 pressure angle of hob cutter (Fig. 3(b)) \u03b31 cross angle of hob cutter in generating RA worm gear (Fig. 5) \u03b81 rotation angle of hob cutter in screw surface generation (Fig. 4) \u03bb1 lead angle of hob cutter (Fig. 4) \u03d51, \u03d52 rotational angles of hob cutter and RA worm gear, respectively (Fig. 5) \u03c91, \u03c92 angular velocities of hob cutter and RA worm gear, respectively (Eqs. (24), (25), (28) and (33)) \u03c9i (j) angular velocity vectors, expressed in coordinate system Sj (j=1, p) of hob cutter and RA worm gear (i=1, g) (Eqs. (23)\u2013(26) and (28)) meshing system", + "5\u00b0 to a minimum of 10\u00b0 [2]. This should pay more attention on the checking of the teeth of hob cutter and generated RA worm gear should not be pointed. The pitch lines of the hob cutter in generating semi RA and full RA worm gears become dx= 1.0 and 2.0 of normal modulus, mn (i.e. axial modulus mx=mn/ cos \u03bb1), below the middle of cutting tooth height, respectively, as shown in Fig. 2 where dx is the distance measured from the middle of cutting blade height to the varying pitch line (also refer to Fig. 3(b) and (d)). Thus, tooth thicknesses of the generated semi RA and full RA worm gears at their normal pitch circles become 1.22 and 0.86mn, and their normal throat radii are (T2+2)mn/2mm and T2mn/2mm, individually. Symbol T2 denotes the tooth number of the generated RA worm gear. The above design gives different proportional changes of addendum and dedendum of hob cutter in generating semi RA and full RA worm gears. Especially, for a full RA worm gear, the pitch circle and throat circle are identical", + " Besides, the standard proportional tooth worm gear is a special case of the RA type worm gear when dx equals 0, no doubt, the pitch line of the hob cutter is at the middle of cutting tooth height. This gives that tooth thickness of the generated worm gear at its normal pitch circle is 1.57mn, and normal throat radius is (T2+4)mn/2mm. The ZN worm-type hob cutter with normal profile of straight-lined edge shape is chosen to generate the RA worm gear in this study. A right-handed ZNworm-type hob cutter, as shown in Fig. 3, is chosen to generate RAworm gears in this study. The surfaces of ZN worm-type hob cutter can be cut by a blade. At first, the cutting blade is placed on the groove normal plane of the ZN wormtype hob cutter, as shown in Fig. 3(a). The blade, inclining with a lead angle \u03bb1 , is performed a screw motion with respect to the hob cutter axis. According to Fig. 3(b), the normal section of the cutting blade consists of straight-lined edge and circular tip, which generate the tooth surface and fillet surface of the worm gear, respectively. In Fig. 3(b), l1 denotes a design parameter of the cutting blade straight-lined edge surface, starting from the intersection point Mo of the two straight-lined edges to the end point Mb. And the moving point M1, represents any point on the cutting blade straight-lined edge surface, moving from the initial pointMa to the end pointMb.MoMa denotes the shortest distance of the cutting blade straight-lined edge surface, i.e. l1(min), while MoMb indicates the longest distance of the cutting blade straight-lined edge surface, i.e. l1(max).\u03b11 denotes the pressure angle formed by the straight-lined edge and the Xc-axis, as shown in Fig. 3(b). The cutting blade width bx equals the normal groove width of the hob cutter, varying with the pitch line of the hob cutter in generating RAworm gears, as explained in Section 3. In Fig. 3(c), symbols ro, r1 and rf represent the outside radius, pitch radius and root radius of the ZN worm-type hob cutter, respectively. Therefore, the generating line of cutting blade can be represented in coordinate system Sc(Xc, Yc, Zc) that fixed to the blade normal plane, as shown in Fig. 3(b), as follows: Rc = rt + l1 cos \u03b11 0 l1 sin \u03b11 1 2 664 3 775; \u00f01\u00de where the upper sign of \u201c\u00b1\u201d sign denotes the left-side of cutting blade while the lower sign denotes the right-side of cutting blade. rt is the distancemeasured from the rotation center of ZNworm-type hob cutter O1 or Oc to pointMo, as shown in Fig. 3(b) and (d). Therefore, rt, bx, l1(min), l1(max) and h can be expressed as follows: rt = ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi r21\u2212 bx 2 sin \u03bb1 2 s \u2212 bx 2 tan \u03b11 ; \u00f02\u00de bx = bn\u22122dx tan\u03b11; \u00f03\u00de l1 min\u00f0 \u00de = 1 2cos\u03b11 bn tan \u03b11 \u2212h ; \u00f04\u00de l1 max\u00f0 \u00de = 1 2cos\u03b11 bn tan\u03b11 + h ; \u00f05\u00de and h = ht\u2212rc 1\u2212sin \u03b11\u00f0 \u00de; \u00f06\u00de where bn is the blade width at the middle of cutting blade height, and bn=\u03c0mn/2, h and ht denote straight-lined edge height and whole height of the cutting blade, respectively, and rc represents the cutting blade tip radius, and dx is the distance measured from the middle of cutting blade height to the varying pitch line of the hob cutter" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001529_icphm.2011.6024357-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001529_icphm.2011.6024357-Figure4-1.png", + "caption": "Fig. 4 Diagram of the two-stage planetary gearbox and sensor locations", + "texts": [ + ", k kjr n r n j jT ki ki b j k i k j i j jk i j ki ki k j jk i j ki ki k C s t x b for k j i n x b for k j j r i n \u03b5 \u03b5 \u03b5 \u03b5 \u03c6 \u03b5 \u03b5 \u03c6 \u03b5 \u03b5 \u2212 = = = = = + + \u22c5 \u2212 \u2264 \u2212 + > = = \u22c5 \u2212 \u2264 + + > = + + = \u2211 \u2211\u2211 \u2211 \u2211w w w w w (10) where j runs over 1, 2, \u2026, r-1 and nk is the number of samples in the rank k. By solving the optimization problem in Eq. (10), the optimal w and bj will be found, and thus a ranking model (Eq. (7)) will be built. III. PLANETARY GERABOX TEST RIG A planetary gearbox test rig shown in Fig. 3 was used to perform fully controlled experiments. It includes a 20 HP drive motor, a bevel gearbox, a two-stage planetary gearbox, two speed-up gearboxes and a 40 HP load motor. The transmission ratio of each gearbox is listed in Table 1. 4/8 Fig. 4 shows the schematic diagram of the two-stage planetary gearbox. The 1st stage sun gear is connected to the driven bevel gear by shaft #1. The 1st stage planet gears are mounted on the 1st stage carrier which is connected to the 2nd stage sun gear by shaft #2. The 2nd stage carrier is located on shaft #3. Four accelerometers named LS1, HS1, LS2 and HS2 are located on the housing of the planetary gearbox as also shown in Fig. 4. The pitting damage was artificially created on one of the four planet gears in the 2nd stage planetary gearbox. Four levels of pitting damage were tested, i.e. baseline, slight, moderate, and severe pitting (Fig. 5). Vibration data were collected from four accelerometers with a sampling frequency of 10 KHz at two load conditions (i.e. no load and 10,000 lb load), and four speeds of the driven motor (i.e. 300, 600, 900, and 1200 revolutions per minute, RPM). For each damage level, 10-minute data were recorded at each of the eight combinations of speed and load conditions, and were split equally into 20 samples" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002076_optim.2017.7975006-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002076_optim.2017.7975006-Figure3-1.png", + "caption": "Fig. 3. Elementary coils of the phases U and V, object of the two-phase short-circuit and the circuit associated to the field model", + "texts": [ + " This short-circuit is modeled through the resistor Rshc in the lower image of the associated circuit, resistor which connects the input and output terminals of the short-circuited coil. A very high resistance 107 \u2126 of the Rshc resistor is considered for the healthy (HE) state of the motor operation. The values corresponding to faulty states with increasing fault severity FA1, FA2, FA3 and FA4 are 90 \u2126, 9 \u2126, 5 \u2126, and 0.9 \u2126 respectively. For the first three faulty states, the post processing of the results of the correspondent Flu3D applications show the values of the current in the short- circuited elementary coil of phase U are lower than the motor rated current. Fig. 3 corresponds to a two-phase short-circuit. The output terminal of the first elementary coil of the phase U - red colored, is in electric contact with the input terminal of the first elementary coil of the phase V, yellow colored in Fig. 3. This short-circuit is modeled through the resistor Rshc1 in the lower image of the associated circuit model. The values of the resistance of this resistor for two faulty states FA1 and FA2 of the motor are 500 \u2126 and 50 \u2126. The time dependence of the phase U current Iu(t) for the healthy state of the motor in the interval 0.16 .... 0.20 sec., Fig. 4, that covers two period of the voltage supply frequency, contains besides the fundamental harmonic of 50 Hz, harmonics which are represented in Fig. 4", + " Taking into account the magnetic field in the region of stator winding overhang through the addition Bz[P](t) + Bz[Q](t) of the axial component of the magnetic flux density, Table IX and Table X emphasize the best choice for the fault detection through the magnetic field, the harmonics of 1250 Hz and 1950 Hz. An important property with these harmonics is the consistent efficiency of fault detection in the initial phases of the fault development. The one-phase short-circuit reflects the local weakening of the electrical insulation between turns belonging to a phase. The two-phase short-circuit in Fig. 3 is determined by the local deterioration of insulation between the phases U and V. For the two-phase short-circuit FA2 the supply currents are Iu = 9.529 A, Iv = 15.524 A and Iw = 6.673 A. The current in the last seven elementary coils of phase U is 7.036 A. The current in all eight coils of the phase V is 6.142 A. The currents in all coils being under the rated current, the faulty case FA2 represents a short-circuit that can be supported for the long time by the motor no-load operation. The harmonics 1050 Hz and 1950 Hz, Table XI, are found for the detection of the two-phase short-circuit" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001446_1350650112460799-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001446_1350650112460799-Figure1-1.png", + "caption": "Figure 1. Two cases of the contact of a hemisphere against a rigid flat: (a) deformable base case and (b) rigid base case with additional bottom rigid plane.", + "texts": [ + " The JG model8 is adopted for early stage contact, and empirical expressions for other stages are obtained by fitting finite element results. By comparing with some existing models and experimental results, this study can predict the overall contact behavior of an elastic perfectly plastic hemisphere against a rigid plane well. Since this problem is axisymmetric, the hemisphere could be modeled as a quarter of a circle and the at UNIVERSITE LAVAL on June 13, 2014pij.sagepub.comDownloaded from rigid flat could be modeled as a rigid line, as shown in Figure 1. The radial movement of the symmetry axis is restrained in this model. The axial deformation of hemisphere base is not considered in this study. Since the radial movement of hemisphere base has significant influences on heavily deformed contact behaviors, the radial deformable and radial rigid bases are considered in this study. As shown in Figure 1(a), the hemisphere base could deformable freely in radial direction. The radial rigid base case is illustrated in Figure 1(b). For radial rigid base case, the deformed hemisphere would pass through the base plane of hemisphere, which is rarely happened practically. Therefore, a rigid plane is added at hemisphere bottom base. Note that the axial deformation of hemisphere base is prevented in this study. For the contact behavior of asperities on rough surfaces, the axial deformation of substrate could be considered using the existing models.27,30,31 At initial stage of contact, the hemisphere deforms elastically. Contact area and contact load can be obtained by Hertz solution6 Ae \u00bc R" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000647_0954410011414993-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000647_0954410011414993-Figure1-1.png", + "caption": "Fig. 1 Missile coordinate system", + "texts": [ + " These results show that in the complete closed-loop system, including the input-dependent terms of the model, which are ignored for simplicity in design, the NCEA autopilot accomplishes angle of attack trajectory control, despite large uncertainties in the aerodynamic parameters. These results also show that for properly chosen feedback and adaptation gains, the tracking error control performance of the NCEA law is better than the CEA law. 2 MISSILE LONGITUDINAL DYNAMICS The pitch-plane rigid missile model considered here has two translational degrees of freedom and one rotational degree of freedom. Figure 1 illustrates the coordinate system and variables used for describing the longitudinal dynamics of the missile. U and W are the components of the velocity vector VT along the body-fixed X and Z axes, L the aerodynamic lift, D the drag, and MY the pitching moment. The mathematical model used in this article is based on a tail-controlled missile. This model assumes constant mass, that is, post burnout, zero roll angle and roll rate, no sideslip, and zero yaw rate. Under these assumptions, the motion of the missile is described by two force equations and one kinematic equation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002131_metroaerospace.2017.7999605-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002131_metroaerospace.2017.7999605-Figure7-1.png", + "caption": "Fig. 7. The coordinate system used for the mathematical model of the test bench.", + "texts": [ + " The measured values are shown to the user through the graphical user interface (GUI) depicted in Fig. 5. This GUI consists of three panels, which report the attitude, power consumption, thrust force and raw data measurements, respectively. In the following subsections the measurements provided by \u201cDronesBench\u201d are described. In this subsection, the mathematical model used for evaluating the module of thrust force given by the VTOL-RPAS is reported. In this analysis, the coordinate system has been considered according to Fig. 7. In particular, the chosen origin corresponds to the reference plane center and the points S1, S2, and S3 represent the load cell coordinates. By using the three load cells, the system allows to measure the forces applied by RPAS to the accommodation plane (see Fig. 8). Each load cell measures the force F s i . The corresponding reaction force of the chain Ri can be evaluated as follows: Ri = F s i cos(\u03b1i) (2) where \u03b1i corresponds to the elevation of vector distance \u2212\u2212\u2192 SiTi respect to its projection on the plane defined by the coordinates of the three load cells S1, S2, and S3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002347_iros.2017.8206622-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002347_iros.2017.8206622-Figure3-1.png", + "caption": "Fig. 3. Orientation sensors at B may provide the value of \u03b1V and/or \u03b1H", + "texts": [ + " In this paper we are considering another measurement possibility which consists in getting complete or partial information on the cable direction at the anchor points A and B. The measurement at point A are (figure 2): \u2022 the angle \u03b8V between the x axis and the vertical plane that includes the cable \u2022 the angle \u03b8H between the horizontal direction and the cable tangent at A We introduce a mobile frame attached to the platform whose center is G, the center of mass of the platform, and vectors xr,yr, zr. Let u be the unit vector of the cable tangent at B and up its projection in the xr,yr plane. The measured angles (figure 3) may be \u2022 the angle \u03b1V between up and xr \u2022 the angle \u03b1H between up and u Realizing such measurement has already been considered: for example our CDPR MARIONET-Assist uses a simple rotating guide at A whose rotation is measured by a potentiometer in order to obtain the measurement of \u03b8V while our CDPR MARIONET-VR is instrumented with a more sophisticated cable guiding system which allows for the measurement of both \u03b8V and \u03b8H (figure 4). Our first trials with such a simple system have shown that the accuracy is poor as soon as the cable tension is low" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000546_j.1460-2695.2010.01540.x-Figure11-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000546_j.1460-2695.2010.01540.x-Figure11-1.png", + "caption": "Fig. 11 Subdivision of the interpolation domain.", + "texts": [ + " However, the influence of the variables is generally moderate on the propagation path, while each tested factor weighs significantly on SIF or, in other words, the crack propagation rate. Based on the observations of the previous section, it appears that the highlighted inaccuracies of the model are due to the high level of interaction between \u03b10, hr and np. Apparently, these factors are not perfectly characterized by the initial factorial design. Hence, the proposal is to subdivide \u03b10, hr and np level ranges by introducing two intermediate levels (Fig. 11). This subdivision consists in splitting the interpolation space into eight subdomains ( i). Additional simulation blocks shown in Fig. 11 are composed of 33 new factor combinations. Table 4 lists the factor levels of each subdomain. Therefore, additional BE simulations are required for each sub-factorial design. Despite these extra simulations, the subdivision process does not induce any changes into the interpolation scheme presented in Fig. 6. The only difference is that instead of interpolating the solution over the entire multi-dimensional space (initial interpolation domain) the operation is made within the subdomain that includes the factor values corresponding to the analyzed gear configuration" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000306_mace.2010.5535715-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000306_mace.2010.5535715-Figure1-1.png", + "caption": "Figure 1. Finite element models", + "texts": [ + " It is difficult to control SLS for copper has strong reflectance to the laser. Furthermore, the copper is oxidized easily in the air. Therefore, the studies of copper and copper matrix composites prepared by SLS have been rare reports. In this paper, the distribution of temperature field of copper powder in SLS is studied by ANSYS software, and the factors of sintering parameter on the quality are investigated. . ESTABLISHMENT OF PHYSICAL MODEL The finite element models of sintering are shown in Figure 1. Sponsored by Science and Research Foundation of East China Jiaotong University (01308013), and the Natural Science Foundation of Jiangxi Province(09497, 2009GQC0014, 2008GZC0037). 978-1-4244-7739-5/10/$26.00 \u00a92010 IEEE The models are divided into two parts, namely, substrate and powder. The size of substrate is 4.8mm\u00d72mm\u00d70.5mm, and that of single-layer powder is 3.4mm\u00d71.6mm\u00d70.1mm in figure1(a) and multi-layer powder is 3.4mm\u00d71.6mm\u00d70.3mm in figure1(b). In order to ensure the accuracy of calculation and the increase of the efficiency, the different sizes of mesh are adopted between the substrate and powder. The mesh grids of 0.1mm are applied to sintering layer, while the larger grids are applied to substrate. B. Determination of initial condition and boundary condition [2,3] Assume that the initial temperature T0 of powder bed is uniform distributed before sintering, then the initial condition is shown as follows. 00),,,( TtzyxT t == (1) The main forms of heat transfer are convection and radiation on the upper and other sides of substrate and powder bed in SLS, the boundary condition of heat transfer is shown as follows" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003768_s105261881905008x-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003768_s105261881905008x-Figure3-1.png", + "caption": "Fig. 3. Cantilever constarea beam with the end part with constant cross-section sizes.", + "texts": [ + " This condition is related to the simplicity of obtaining an elastic element, for example, made from steel leaf. (4.2) Constant width \u21d2 ; . This condition is determined, for example, by limitations on the overall dimensions of the spring in a car. (4.3) Constant cross-section area = . This condition is especially important for ensuring the strength of composite beams, as it makes it possible to save the constant number of nonovercut fibers along the whole spring length. Type 4.3 of the constarea beam (Fig. 3) from expression (1) means (6) and from the condition of uniform strength (5): ; . (5) Resistance to shearing force Pmax applied at the beam end (for a distributed load this condition is not necessary, as at the free end the shearing force is zero). Let us accept this condition as the interlaminar shear stresses reaching the critical value , then it defines length a (2) of the end part, the cross-sectional area of which should be kept constant (7) \u2212 = \u03b1 + \u03b2 =2 2 6 ( ) 6 ; from (1) 2 1. ( ) ( ) (0) (0) i i i i i i P l x Pl w x t x w t =1(0) constt \u03b2 =1 0 \u03b1 =1 1 =2(0) constw \u03b1 =2 0 \u03b2 =2 1 2 3 3( ) ( )w x t x 3 3(0) (0)w t \u03b1 + \u03b2 =3 3 0, \u03b1 = \u22123 1 \u03b2 =3 1 \u03c4* \u03b1 +\u03b2\u2212 \u2212 = = = \u03c4 max3*( ) ( ) (0) (0)( ) " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001243_s11044-012-9325-8-Figure11-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001243_s11044-012-9325-8-Figure11-1.png", + "caption": "Fig. 11 The different configurations of the extensor apparatus during extension [16]: (a) the distal section of the extensor digitorum comunis is stretched, (b) the middle section is stretched, (c) the proximal section is stretched. \u201cEDC\u201d stands for extensor digitorum comunis, \u201cL\u201d stands for lumbrical and \u201cIO\u201d stands for interossei", + "texts": [ + " 5, which is a system of tendons that mediates the long extensors (extensor digitorum comunis, the extensor indicis and the extensor digit minimi, for the index and little finger, respectively) and some intrinsic muscles in the finger extension. The major feature of the extensor apparatus is that, unlike the modeling premise, not all segments of the tendons are fully stretched. This is used to move the insertion point of the long extensor throughout the phalanxes depending on the angle of the joints and the intrinsic muscles (lumbrical and the two interossei) [16, 24], Fig. 11. To 4Although the work done in [26] is for the index finger only, the parameters can be adapted to be used in the other fingers since they share common muscles and the specific ones are similar between each other. 5This condition stems from the modeling of the tendons as inelastic and always stretched \u201cwires\u201d. model this behavior, at each iteration, the application point of the external forces from the extensors have to be checked against the joint positions and changed accordingly. The second special case is the lumbrical muscle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002741_012011-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002741_012011-Figure4-1.png", + "caption": "Figure 4. Kit Top View", + "texts": [ + " Therefore, the suitcase that is used as the main casing of the kit has size 41 x 23 x 11.8 cm. Portable concept has given the limit with the size of a rigid suitcase. Thus all the components to be used in the kit should take into account that dimension. Figure 2 and Figure 3 are component layout and kit dimention. 5 1234567890\u2018\u2019\u201c\u201d Another requirement to be compiled with portable concept is the weight of the kit. Therefore components selection is important step. Table 1 shows the selected components taking into consideration the available space. The Figure 4 shows the view of the kit that has been designed. In order to keep the box light, it is made from acrylic 3 mm and the main base is acrylic 5 mm. 6 1234567890\u2018\u2019\u201c\u201d 3.3.1 Output Component and Terminal. The lamp as an input component is selected which has a medium size with a commonly rated 24V, voltage commonly used in the PLC output. But 3 phase electric motor is too big to be placed on kit. Therefore, the motor control simulation can be replaced by setting the direction of DC motor rotation. Clockwise (CW) or Counter clockwise (CCW) adjustments can be done with the help of 2 relays through electrical signals" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002880_s00170-018-2687-1-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002880_s00170-018-2687-1-Figure6-1.png", + "caption": "Fig. 6 Steady thermal-induced deformation fields of the worktable assembly", + "texts": [ + " It is possible because although a part of cutting heat transfers to the worktable assembly by heat conduction, most of it still remained at the workpiece, which caused the highest temperature to appear at the surface of the workpiece. On the other hand, due to the long heat transfer distance and large exposed area to surrounding air, the temperature of the shell at the bottom of the worktable assembly had the lowest temperature. Based on the obtained temperature field of the worktable assembly, analyses of structural deformation in three axisdirections were carried out as well. The thermal-induced deformation fields in the x-, y-, and z-directions are shown in Fig. 6, respectively. Figure 6a revealed that the largest total thermal-induced deformation of the worktable assembly appears at the fixture (49.88 \u03bcm). The explanation for this observation is that the upper end of the fixture directly contacted the workpiece where the cutting heat entries, whereas its opposite end was close to the base shell, which had lowest temperatures in the cutting process. This temperature difference between the two ends of the fixture causes the largest temperature gradient, leading to the largest thermal-induced deformation. Because the thermal-induced deformation in the y-axis direction was the most important effect factor of the tooth profile error in gear hobbing, it was necessary to find out which component is the most vulnerable to the y-axis-direction deformation. As shown in Fig. 6b, the largest y-axis-direction deformation appeared at the workpiece, which was 43.09 \u03bcm. The reason is that the workpiece was fixed by the fixture and the center above it, which limited the displacement in the axial direction. Therefore, the thermal-induced deformation has to happen at other directions. To further comprehend how the temperature and deformation fields vary with time, the transient-state thermal-structure characteristics of the worktable assembly were analyzed as well. Eight points from the surface of the worktable assembly were selected and monitored" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002598_j.ymssp.2018.04.033-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002598_j.ymssp.2018.04.033-Figure4-1.png", + "caption": "Fig. 4. Excitation by (a) rotor mass eccentricity and by (b) dynamic magnetic eccentricity.", + "texts": [ + " The displacements of the bearing housing points B are described by zb and yb. The displacement of the stator mass is described by zs and ys. Additionally also the rotation us of the stator mass has to be considered. The displacements of the motor feet points AL and AR are described by zaL and yaL (left side) and by zaR and yaR (right side) and the displacements of the foundation points FL and FR by zfL and yfL (left side) and by zfR and yfR (right side). The excitation by rotor mass eccentricity e\u0302u and by dynamic magnetic eccentricity e\u0302m is shown in Fig. 4. The rotor mass eccentricity e\u0302u creates an unbalance force F\u0302u, which is rotating with the rotor angular frequency X. With the angle au \u00bc X t \u00feuu the vertical and the horizontal unbalance force, which act at the rotor mass, can be described as follows: f uz \u00bc f\u0302 u cos\u00f0Xt \u00feuu\u00de; f uy \u00bc f\u0302 u sin\u00f0Xt \u00feuu\u00de; with f\u0302 u \u00bc e\u0302u mw X2 \u00f01\u00de The dynamic magnetic eccentricity e\u0302m creates together with the electromagnetic stiffness cm a magnetic force f\u0302m which is also rotating with the rotor angular frequency X, acting at the rotor mass and in opposite direction at the stator mass" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000028_s12206-009-1166-x-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000028_s12206-009-1166-x-Figure4-1.png", + "caption": "Fig. 4. Vector geometry of the serial-link section.", + "texts": [ + " The position of the TCP along the i-j plane is changed by the parallel-link arm section. Also, the position of the TCP along the k-axis is changed by the serial-link arm section. The pose around the yawing direction and pitching direction is changed by the serial-link arm section. The positions and pose of the TCP are controlled individually in each section. Therefore, to calculate the kinematic analysis, we divided this hybrid robot into each section. The geometrical model of each section is defined in Fig. 3 and Fig. 4 individually. Especially, the pose of the constant vector C8 depends on the length of the linear motion vector L3 and L4. The position of the TCP on the i-j plane is changed by the pose of the constant vector C8. So, the constant vector C8 is included in the parallel-link arm section. At the serial-link arm section the origin point is defined as Os. The forward kinematic analysis calculates the position and pose of the TCP for this hybrid robot [3]. At this analysis, each length of the motion vectors L3,L4 and L7 is defined as l3,l4 and l7", + " 51 4k k \u03c9\u03b8\u03b8 \u03b8=E E E E I (1) In this equation, the E is the pose matrix; the vector \u201cI\u201d shows the initial position of the TCP. The position of the TCP on the projected i-j plane, based on Fig. 3, can be calculated as Eq. (2). 2 6 6 8 1 4 6 6 8 1 cos sin sin cos c p cx l p cy \u03b8 \u03b8 \u03b8 \u03b8 + +\u239b \u239e\u239b \u239e = \u239c \u239f\u239c \u239f + +\u239d \u23a0 \u239d \u23a0 (2) In this equation, c2 shows the length of the constant vector C2; p6 shows the length of the polar P6; c8 shows the length of the constant vector C8. And, the position along the k-axis, based on Fig. 4 can be calculated as Eq. (3). 7 9z l g= + (3) In this equation, g9 shows the length of the constant vector G9. Also, the position at the intersection of the polar vector P5 and P6 moves on the center-line as shown in Fig. 3. Therefore, \u03b81, \u03b86 and p6 are shown as Eq. (4). 3 4 1 2 6 1 2 6 1 2 2 cos l l c cp \u03b8 \u03c0\u03b8 \u03b8 \u03b8 \u2212 = = \u2212 = (4) The inverse kinematic analysis calculates the length of each link and rotation angle from the position and pose of the TCP [3]. At first, to calculate the rotation angle \u03b85 and \u03b81+\u03b84, the pose of TCP is defined as Eq. (5). 51 4k k \u03c9\u03b8\u03b8 \u03b8=E E E E I (5) To calculate the length of l7, the position of the TCP which is along the k-axis, as shown in Fig. 4, is defined as Eq. (6). The length of l7, based on Eq. (6), can be calculated as Eq. (7). = +k k 7 9P L G (6) 7 9l z g= \u2212 (7) The polar vector P shows the position of the TCP which is along the k-axis. The position of the TCP on the projected i-j plane, as shown is Fig. 3, is defined as Eq. (8). And the length of l4, based on Eq. (7), can be calculated as Eq. (9). 0 2 4 6 8= + + +P C L P C (8) 2 8 1 0 0 4 0 0 8 1 6 sin sincos cos tan c c pl p c \u03b8 \u03b8\u03b8 \u03b8 \u03b8 + \u2212 = \u2212 + (9) The polar vector P0 shows the position of the TCP on the i-j plane; the p0 shows the length of the polar vector P0; \u03b80 shows the angle between the polar vector P0 and j-axis; \u03b84 shows the angle between the linear motion vector L4 and j-axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000102_s00354-007-0062-0-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000102_s00354-007-0062-0-Figure1-1.png", + "caption": "Fig. 1 System of the Helicopter Coordinates15)", + "texts": [ + " The next section introduces a helicopter system based on coordinates with assumptions for fixed hovering and hovering with disturbances. Section 3 describes the entry fuzzy regulator based on pilot knowledge and on the preceding materials. Section 4 describes the initial model and the modified model. Simulated results with and without the fuzzy regulator both before and after modifications appear in Section 5. In Section 6, an overview is presented of intelligent helicopter controllers. Finally, Section 7 consists of concluding remarks. \u00a72 Description of Helicopter Dynamics 2.1 System of Coordinates In Fig. 1 a right-skewed system of helicopter coordinates is presented. Oxgygzg is a moving coordinate system connected with the direction of gravitation, while Oxyz is a system connected with the fuselage. These systems are interconnected by angles of roll \u03a6, pitch \u0398 and yaw \u03a8. Roll is rotation around the longitudinal axis drawn through the body of the vehicle from tail to nose in the normal direction of flight. Pitch is rotation around the lateral axis running from the pilot\u2019s left to right in piloted aircraft; thus the nose pitches up and the tail down, or vice-versa", + " In practice these values mean that the simulated hover state with the application of the proposed fuzzy regulator is very close to the ideal hover state while, in the same conditions but without the proposed fuzzy regulator, the helicopter would crash. Comparing this result with the result presented in Fig. 3 one can see that without the regulator and without the pilot reacting by steering e.g. the velocity W after 70[s] exceeds the exploitation limit whilst with the proposed fuzzy regulator, the velocity W (like the other parameters) stabilizes. 5.2 Response of the Fuzzy Regulator to Disturbances Research on the influence of disturbances in all three linear directions x, y, z according to Fig. 1 was carried out. Restrained gusts defined in subsection 2.4 were used. Stabilization of the periodic small moves of the fuselage was achieved in all cases. However, a longer parameter stabilization time was observed than in cases where there was no wind. The biggest boggles of parameters took place always in the direction of the gust. Moreover, some of the flight parameters exhibit a small but constant regulation deviation (Fig. 9). After analyzing the initial model there was a possibility to compute a worse adjustment, especially when longitudinal velocity is \u201cmore bigger\u201d than angular velocity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000525_978-3-642-23244-2_54-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000525_978-3-642-23244-2_54-Figure1-1.png", + "caption": "Fig. 1. Kinematic scheme of DELTA robot [5]", + "texts": [], + "surrounding_texts": [ + "This project has been motivated by a requirement from industry. Our department received a proposal for possible cooperation in development of the robot with parallel kinematics \u2013 DELTA type robot. The company Dyger, s.r.o., is an exclusive representative of the company Beckhoff and has at its disposal the recentlyreleased data library for control of robotic appliances with parallel kinematics. Hence, we attempted to make use of these new possibilities and to develop an affordably priced DELTA robot which could be further offered for solutions of specific industrial applications. As the supplied software is not an open source software and it is not possible to simply verify if the mathematical description of the mechanism are valid, we proceeded to develop our own algorithms for calculations of position, speed, acceleration and, first of all, for calculation of the mechanism\u2019s working space and equally important singular positions. Our objective was also to verify the mathematical apparatus used in the supplied software [1]." + ] + }, + { + "image_filename": "designv11_33_0003741_oceanse.2019.8867249-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003741_oceanse.2019.8867249-Figure1-1.png", + "caption": "Fig. 1. Mechanical Structure of UWG", + "texts": [ + " The overall system was designed through an iterative testing and validating process considering the requirements established according to the environmental characteristics of the Peruvian coastline and the UWG state-of-art. Those requirements are presented as follows. \u2022 The main structure must house the electronics in a watertight compartment. \u2022 The shape and configuration of the hull must ensure the buoyancy and stability of the whole system. \u2022 The UWG must resist depths of 150m maximum accord- ing to the Peruvian coastal shelf. \u2022 the UWG must have a minimum time of autonomy of seven (07) days. The UWG is composed of a total of five parts as shown in Figure 1: a) An acrylic enclosure with static seals to protect electronic from water damage. b) Buoyancy engine to regulate the mass of the UWG. c) Battery block which acts a sliding and rolling mass. d) Internal actuators to regulate the position of the battery block and change the vehicle pitch and roll. 1) Battery Holder: The battery holder was designed as the main moving mass inside the UWG which enables the shifting of the center of mass in order to regulate the vehicle\u2019s pitch and roll. It is designed to hold 16850 lithium batteries, providing a total vehicle autonomy of seven (07) days" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002485_aab928-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002485_aab928-Figure1-1.png", + "caption": "Figure 1. Schematic cut view of the outlet of the printhead, the nozzle chip, and the rinse gas flow (blue) (left); and SEM picture of the inner part of a StarJet nozzle chip fabricated from silicon (right).", + "texts": [ + " The StarJet technology enables the direct printing of molten metals in the shape of microdroplets or jets by a pneumati cally driven printhead. The following sections\u00a0 introduce the general technological aspects of the StarJet principle, and the design of the StarJet printhead used for printing molten Al alloys. Printheads based on the StarJet technology feature a heatable reservoir with a cap to close the reservoir, which contains the molten metal. In the lower part of the printheads, there is an interchangeable nozzle chip with a starshaped orifice, from which the droplets are ejected (see figure\u00a01). The nozzle orifice diameter of the nozzle chip mainly determines the droplets\u2019 size. For printing lower melting point and less cor rosive metals, such as solder, these nozzles can be fabricated from silicon by deep reactive ion etching (DRIE). Figure\u00a0 1 also depicts a scanning electron microscopy (SEM) image of an exemplary nozzle chip made from silicon, with the main orifice and the bypass channels clearly visible. Due to non wetting contact angles of the metal melt on the nozzle chip material, and the dedicated star shaped geometry, the droplets are formed inside the nozzle chip orifice by capillary forces J. Micromech. Microeng. 28 (2018) 074003 before they are ejected [21]. StarJet printheads are operated with two independent gas pressures: the socalled actuation pressure is used to apply short pressure pulses by a solenoid valve onto the reservoir of molten metal; and the socalled rinse gas flow constantly supplies gas (typically argon or nitrogen) through the bypass channels of the nozzle chip. This gas flow surrounds the droplet after ejection with the rinse gas, as shown in figure\u00a01. In dropondemand mode, a pres sure pulse of a few milliseconds, applied by a solenoid valve, leads to the ejection of a single droplet from the nozzle orifice. Typically applied pressures are in the range of 50\u2013500 mbar, depending on the size of the nozzle chip orifice. StarJet printhead for molten Al alloys The cap closing the reservoir features a gas inlet and is sealed via a custommade temperaturestable phlogopite mica seal (novamica THERMEX, Frenzelit, Germany). Sealing is vital for the pneumatically actuated printhead because short pressure pulses of a few milliseconds are used for actuation, and even a small leakage can impede operation", + " Therefore, nozzle chips from silicon carbide (SiC) are used for printing Al alloys because of their inert nature toward the Al alloy melts. Such SiC chips cannot be fabricated by DRIE but by laser structuring, which com promises the structural quality of the design. The nozzle chips used in this work were fabricated at FraunhoferInstitut f\u00fcr Lasertechnik (Fraunhofer ILT, Aachen, Germany) with a pico second laser (pulse duration 7 ps at a wavelength of 515 nm). The fabrication of the complex starshaped geometry in SiC with high aspect ratios remains especially challenging for small orifice diameters, and the intended shape (see figure\u00a01 for reference) cannot be perfectly generated. Still, nozzle chips with minimum diameters of 130 \u00b5m in the center were successfully fabricated from SiC, and the starshaped struc ture was sufficiently established, as shown in figure\u00a04. The droplet generation process was experimentally investigated with a stroboscopic setup that comprises a com plementary metaloxidesemiconductor camera (\u00b5eye, IDS Imaging Development Systems GmbH, Germany) with a Schneider Kreuznach Coponon lens system (Jos. Schneider Optische Werke GmbH, Germany) and a lightemitting diode illumination for a transmission light setup, where the printhead is placed between the light source and camera" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002896_tia.2018.2877186-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002896_tia.2018.2877186-Figure1-1.png", + "caption": "Fig. 1. Structure of a 12-/8-pole DSDSEM.", + "texts": [ + " The variable reluctance structure with the distributed arrangement of the field windings is given in [5] and [6] to obtain the sinusoidal back electromotive force (EMF) for the torque ripple suppression, and the traditional skewed rotor of DSEM is proposed in [7] with the corresponding six-state commutation principle to reduce the torque ripple. Nevertheless, the power density is decreased due to the lower salient ratio of the variable reluctance structure. A novel 12-/8-pole dual-stator DSEM (DSDSEM) has been proposed in [8]. As shown in Fig. 1, the dual-stator structure shares the common field winding. The rotor is divided into two sections and one section is shifted 7.5 mechanical degrees to another, which is equal to 60 electrical degrees of the 12/8 poles. Although the power density of DSDSEM is lower than the conventional DSEM due to the large volume of the two stators, with the corresponding two-section twisted-rotor structure, the output torque ripple of the DSDSEM is much lower than the conventional DSEM in the driving mode, and in the generating mode, the output dc voltage ripple could also be reduced by the compensation of the dual stators; thus, it could be the candidate as the starter/generator for the hybrid electric vehicle or aircraft 0093-9994 \u00a9 2018 IEEE", + " Then, the commutation principle of the driving system is proposed in Section III. The interactions of the coupled outputs are analyzed and the optimized control strategy is proposed in Section IV. Section V shows the simulation results of the proposed system operating in the normal control strategy and the optimized control strategy, and experiments are presented in Section VI to validate the feasibility and effectiveness of the proposed driving system. Finally, the conclusion is drawn in Section VII. Due to the 60 electrical degrees shift of DSDSEM\u2019s rotor, as shown in Fig. 1, the inductances of phase windings A, B, and C in stator 1 are 60 electrical degrees ahead of the inductances of phase windings U, V, and W in stator 2, respectively. As the inherent feature of the DSDSEM, the linearized self-inductances and mutual inductances between the phase windings and field windings are shown in Fig. 2, where LA , LB , LC , LU , LV , and LW are self-inductances, LAF , LBF , LC F , LU F , LV F , and LW F are mutual inductances, \u03b8e is the electrical degrees of the rotor position, and six modes related to the rotor position are defined from I to VI" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001517_ical.2011.6024750-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001517_ical.2011.6024750-Figure1-1.png", + "caption": "Fig. 1 The inertial earth-fixed frame and the body-fixed frame for a ship.", + "texts": [ + " The concept of group biasing, adaptive group biasing strategy and thrust allocation algorithm based on the strategy are proposed in Section III. The simulation results of optimizing adaptive thrust allocation algorithm of a dynamic positioning ship are presented in Section IV. Finally, the conclusion is proposed in Section V. The horizontal motion of a DP ship is usually described by the motion components in surge, sway, and yaw. The different reference frames used in dynamic positioning are illustrated in Fig. 1. The position [x, y] and yaw angle \u03c8 of the 978-1-4577-0302-07/11/$26.00 \u00a92011 IEEE 394 vessel are expressed in the earth-fixed reference frame while the surge, sway and yaw velocities [u, v, r] are expressed in the body-fixed frame. A low-frequency ship model in 3 DOF is employed in [11] since dynamics at higher frequencies are negligible in station-keeping. ( )J M D w \u03b7 \u03c8 \u03c5 \u03c5 \u03c5 \u03c4 = + = + (1) where [ , , ]x y\u03b7 \u03c8 \u03a4= , [ , , ]u v r\u03c5 \u03a4= and: cos sin 0 ( ) sin cos 0 0 0 1 J \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u2212 = \u23a1 \u23a4 \u23a2 \u23a5 \u23a3 \u23a6 (2) \u03c4 is a force vector provided by the thruster system, w is environmental disturbances, M and D is mass matrix and damping matrix respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000837_jfm.2012.45-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000837_jfm.2012.45-Figure2-1.png", + "caption": "FIGURE 2. Schematic of the moving Fo\u0308ppl problem and the relevant variables.", + "texts": [ + "4) using Nv = 1, \u03ba1 = \u03ba and L0 = 0 as dZc dt = i\u0393 \u03c0a2(\u03c1s/\u03c1 + 1) [ Zv1 \u2212 a2 l2 1 (Zv1 \u2212 Zc) ] , (3.5) where \u0393 = 2\u03c0\u03ba is the circulation of the free vortex. Hence, the evolution trajectories for the point vortex and the cylinder presented in the figure 1 of their work can be exactly reproduced. 3.3. With a vortex pair: the moving F\u00f6ppl problem Next, we consider the moving Fo\u0308ppl problem, where a pair of vortices of opposite strengths, \u03ba1 = \u2212\u03ba2 = \u03ba , move behind a non-rotating cylinder of radius a and density \u03c1s along OX1 at velocity U, as depicted in figure 2. This vortex pair is in a symmetric configuration with respect to OX1 with a half-span of H and a horizontal distance D from Zc. The complementary set-up where a uniform flow passes a fixed cylinder with a downstream vortex pair is the classic Fo\u0308ppl problem, well examined in the literature (Fo\u0308ppl 1913; Milne-Thompson 1968; Marshall 2001). The moving Fo\u0308ppl problem appears in a few recent works that treat it as a perturbation problem, as introduced below. Firstly, Shashikanth et al. (2002) develop a model for the interaction between a neutrally buoyant cylinder and N free vortices in a potential flow using the Hamiltonian mechanics" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002630_1.5034584-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002630_1.5034584-Figure1-1.png", + "caption": "FIGURE 1. Correspondence between angles in double pendulum with ellipse constraint.", + "texts": [ + " The singularity lies only in the denominator of the second term. 030004-2 We are going to study a mechanical system with singular configuration space using the mathematical double pendulum as an example. We use the following notation: point A is fixed, the rod AB of length l1 is attached to the point A, and the rod BC of length l2 is attached to the point B. We consider a two-dimensional torus Q = T2 with coordinates (\u03d5, \u03c8). Angles \u03d5 and \u03c8 measure the deviation of the rod AB and BC from the vertical axis AX (Fig.1). The system is located in a vertical plane and only gravity affects the system. At the point B there is a point mass m1 and at the point C there is mass m2. We suppose that a geometric constraint F(\u03d5, \u03c8) = 0 is imposed on the coordinates. The Lagrange equations of the second kind with multipliers for the given mechanical system: l21(m1 + m2)\u03d5\u0308 + m2l1l2[cos(\u03d5 \u2212 \u03c8)\u03c8\u0308 + sin(\u03d5 \u2212 \u03c8)\u03c8\u03072] + l1(m1 + m2)g sin\u03d5 = n F\u03d5 |\u2207F| ; (9) l22m2\u03c8\u0308 + m2l1l2[cos(\u03d5 \u2212 \u03c8)\u03d5\u0308 \u2212 sin(\u03d5 \u2212 \u03c8)\u03d5\u03072] + l2m2g sin\u03c8 = n F\u03c8 |\u2207F| . (10) Using the following notation: A(q) = ( l21(m1 + m2) m2l1l2 cos(\u03d5 \u2212 \u03c8) m2l1l2 cos(\u03d5 \u2212 \u03c8) l22m2 ) ; (11) B(q, q\u0307) = ( m2l1l2 sin(\u03d5 \u2212 \u03c8)\u03c8\u03072 + l1(m1 + m2)g sin\u03d5 \u2212m2l1l2 sin(\u03d5 \u2212 \u03c8)\u03d5\u03072 + l2m2g sin\u03c8 ) , (12) it is possible to represent system (9)-(10) as A(q)q\u0308 + B(q, q\u0307) = N", + " But on the torus we can consider only one constraint instead of three. Denote coordinate B(x2, y2), C(x1, y1) and the center of ellipse by S = (a, 0); in singular case a = (l1 \u2212 l2) \u2212 r1. Then: x1 = l1 cos\u03d5, y1 = l1 sin\u03d5; (13) x2 = l1 cos\u03d5 + l2 cos\u03c8, y2 = l1 sin\u03d5 + l2 sin\u03c8; (14) (x2 \u2212 a)2 r2 1 + y2 2 r2 2 = 1. (15) We get the equation F(\u03d5, \u03c8) = 0 if we substitute expressions for x2 and y2 from (14) in (15). Let us study how angles \u03d5 and \u03c8 depend on the u which is the angle between straight line AC and a vertical axis (Fig.1). We introduce the angles \u03b81 := \u2220CAB, \u03b82 := \u2220CBA. The \u03b81 angle is measured from the AC axis counterclockwise, the angle \u03b82 is counted from the AB axis counterclockwise. Angles u, \u03b81, \u03b82, \u03d5, \u03c8 correspondence in the case of \u03b81 > 0, \u03b82 > 0, \u03d5 < 0 is shown in Fig.1. The values |\u03b81(u)| and |\u03b82(u)| could be found from the cosine theorem for 4ABC, and the signs of angles follow from geometric considerations. Denote d(u) = |AC(u)|. We will study a motion of pendulum near the point (0, \u03c0) (Fig.1). Calculation formulas for the parametrization of the configuration space on the torus are the following (we get expression for d(u) as intersection of ellipse (15) and straight line y2 = tan(u)x2): d(u) = \u221a 1 + tan2(u) \u00b7 a r2 1 + \u221a1 \u2212 a2 r2 1 ( tan(u) r2 )2 + 1 r2 1 \u00b7 1 r2 1 + ( tan(u) r2 )2\u22121 ; (16) \u03b81(u) = \u00b1 arccos l21 + d(u)2 \u2212 l22 2l1d(u) ; \u03b82(u) = \u00b1 arccos l21 + l22 \u2212 d(u)2 2l1l2 ; (17) \u03d5(u) = u + \u03b81(u) ; \u03c8(u) = (u + \u03b81(u)) + \u03c0 + \u03b82(u) . (18) 030004-3 If we choose the sign \u201c+\u201d in both equations in (15), we get the following geometric picture: rod BC is always located right of the axis AB" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002564_s1995078017060039-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002564_s1995078017060039-Figure1-1.png", + "caption": "Fig. 1. Scheme of functioning of created screen-printed electrodes modified with SWCNTs.", + "texts": [ + " NANOTECH The measurements were carried out using a Kapel\u2019104T capillary electrophoresis system (Lumex, Russia) with a negative polarity of high voltage and equipped with a quartz capillary (internal diameter is 50 \u03bcm, effective length is 65 cm, and total length is 75 cm) and a photometric detector, allowing measurements at a wavelength of 254 nm. Before analysis, the samples were filtered on a membrane (Millipore, United States). The area of the peak was the analytical signal. The processing and interpretation of the data was carried out on a personal computer using specialized software. A schematic representation of the electrode structure, when the enzyme is encapsulated in the membrane of the cross-linked BSA, is shown in Fig. 1. Figure 2 shows microphotographs of the surface of the developed electrodes at various stages of modification obtained by the SEM method. The graphite working electrode has a highly developed surface (Fig. 2a), which provides the large contact area of the enzyme and nanomaterials with conductive material and allows a high sensitivity of biosensors to be achieved based on the developed electrodes. The matrix (enzyme, nanomaterial, and mediator) evenly covers the surface of the working electrode (Fig", + "89 \u00d7 10\u20132 nmol/cm2) [35]. However, these data correspond to the total concentration of glucose oxidase deposited on the working electrode (~6 nmol/cm2). Thus, prac- tically all of the glucose oxidase deposited on the elec- trode takes part in the electrochemical process. This can be explained by the fact that, due to the presence of carbon nanotubes in the volume of the immobiliz- ing hydrogel, the transfer of electrons from the enzyme molecules located at a distance from the electrode sur- face is facilitated (Fig. 1). This ultimately leads to an improvement in the positive characteristics of the bio- sensor. Figure 6 presents the calibration dependences of the response on glucose concentration for biosensors based on electrodes modified with GOD and SWCNTs with the addition of ferrocene and without a mediator. Both calibration curves are hyperbolic and were approximated using the Michaelis\u2013Menten equation. It can be seen from the comparison that the maximum level of responses of the biosensor based on the modified electrode without ferrocene is significantly higher than that of the analog containing ferrocene" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001997_iccre.2017.7935043-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001997_iccre.2017.7935043-Figure2-1.png", + "caption": "Fig. 2: Rotor", + "texts": [ + " Here, subscripts \u201cl\u201d, \u201cr\u201d, \u201cv\u201d and \u201ch\u201d represent the left-hand of the rotor, right-hand side of the rotor, the vertical direction and horizontal direction, respectively. For example, the vertical direction at the left-hand side of the rotor is represented as \u201clv\u201d. Both the perturbation gj and the rotor angular velocity \u03c9 are the measurable parameters in operation. Each physical parameter of the AMB system is shown in TABLE I. To obtain the motion equations of the AMB system, coordinate axes X, Y, and Z are introduced to the rotor. Fig. 2 shows a schematic diagram of the rotor with the introduced X, Y, and Z axis. Here, the origin of three axes is the center of gravity of the rotor at the equilibrium state. Let x, y, z, \u03c6, \u03b8 and \u03c8 be the position of the center of gravity on X axis, Y axis and Z axis, the rotation angle about X, about Y and about Z, respectively. In this study, the followings are assumed for the AMB system. \u2022 The rotor is rigid body. \u2022 The rotor is symmetrical with respect to both Y and Z axis of rotation. \u2022 The rotor position does not change significantly" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000945_ssd.2010.5585533-Figure15-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000945_ssd.2010.5585533-Figure15-1.png", + "caption": "Figure 15. 3-D flux path in the stator lamination.", + "texts": [ + " Main Flux Paths The flux moves through the magnetic circuit of the CPAES within three dimensional (3D) paths. These are described as follows: \u2022 axially in the stator yoke (see figure 12), \u2022 radially down through the collector and crossing the air gap down to the magnetic ring (see figure 12), \u2022 radially then axially in the magnetic ring (see figure 12), \u2022 axially-radially in the claws (see figure 13), \u2022 radially up in air gap facing the stator laminations (see figure 14), \u2022 radially up in the stator teeth and circumferentially in the stator core back (see figure 15), \u2022 radially down in stator teeth facing the two adjacent claws (see figure 16), \u2022 radially-axially in the adjacent claws (see figure 16), \u2022 axially then radially in the magnetic ring (see figure 17), \u2022 radially up through the air gap and the mag netic collector on the other side of the machine (see figure 17), \u2022 axially in the other side of the stator yoke (see figure 18). 3.2.2. Leakage Flux Paths Not all the flux produced by the excitation winding and the armature contributes to the EMF generation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002297_ldia.2017.8097244-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002297_ldia.2017.8097244-Figure4-1.png", + "caption": "Fig. 4. Sectional view of vernier model", + "texts": [ + " In this paper, the following combination is used. 221 ZZ (5) The current applied to the coil needs to be controlled according to the position of the mover. The thrust is maximum when the phase angle difference between the magnetomotive force due to the coil and the harmonics in the air gap is \u03c0 / 2 in the electrical angle. Therefore, the current of the U, V and W phases are as follows 3 22 sin2 3 22sin2 2sin2 2 L zZ II L zII L zII W V U (6) where the initial position of the mover are set as shown in Fig. 4. L is the pitch width of the mover, z is mover position, and I is the effective value of the current. In this paper, the vernier motor in Fig. 2 is called an 8- 6 vernier motor, and that in Fig. 3 is called a 14-12 vernier motor. The 14-12 vernier motor has twice as many magnets as the 8-6 vernier motor. By increasing Z1 and Z2, the 14-12 vernier motor is expected to generate a larger thrust. The static thrust characteristics of each model were investigated by using 3-D FEM. The amplitude of the current is 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003864_s11012-019-01081-5-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003864_s11012-019-01081-5-Figure6-1.png", + "caption": "Fig. 6 a 3-CPPU parallel manipulator; b particularization of the generic kinematic chain unit vectors", + "texts": [ + " All the figures and simulations of the possible motions described in this section have been generated using a free software package developed by the authors, called GIM [22], intended for the simulation and analysis of mechanisms and robots. Based on the 3-CPCR manipulator, other possible manipulators with the same multioperational characteristics studied in Sect. 5 can be obtained. For this purpose, the orientations of the unit vectors of the generic kinematic chain that define the motion must be maintained. This is the case with the 3-CPPU parallel manipulator represented in Fig. 6(a), where the italicized PU refers to the parallelism between the uc translational vector and the ub rotational vector (the first revolute pair of the U joint), as shown in Fig. 6(b). Moreover, the revolute pairs that connect chains 2 and 3 with the moving platform are located in such a way that uc2 and uc3 are parallel to each other and perpendicular to uc1 [Eq. (27)]. Figure 6 shows a 3-CPPUmanipulator in which the particularization based on the generic chain is also that of Fig. 3b. Its system of velocity problem equations is given by Eq. (31), but substituting d \u00bc 0 in Eq. (5) because points D and E are coincident. All the conclusions derived from the multioperationality analysis of the 3-CPCR described in Sect. 6 are applicable to the 3-CPPU manipulator. With the aim of validating all these theoretical results and verifying that the sets of group displacements established in Sect" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000832_1.3657241-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000832_1.3657241-Figure1-1.png", + "caption": "Fig. 1 Schematic relationship between a roller and two raceways in roller bearing", + "texts": [ + " 1 Numbers in brackets designate References at end of paper. Contributed by the Lubrication Division of THE AMERICAN SOCIETY OP MECHANICAL ENGINEERS and presented at the A S L E - ASME Lubrication Conference, Chicago, 111., October 17-19, 1961. Manuscript received at ASME Headquarters, May 4, 1960. Paper No. 61\u2014Lub-14. Fundamental Equation Denoting that Pb is the total pressure of lubricant film entrapped between the raceway and the roller at the position angle 6 which is measured from the bottom point in unloaded side, as shown in Fig. 1, the load capacity L of roller bearing is given by On (*2t L = - V Pb cos 6 = - Pb cos ddd (1) fee, 2 7 r J o where n is the total number of rollers in the bearing. Then the frictional moment of roller bearing is J = FnUi + FbeUe + 2Fbr Ur, . where J is the frictional work per unit time 011 contact surfaces between a roller and raceways. Pb and Fb are the function of the minimum amounts of film thickness hoi and ho, which are obtainable from hoi + hoe = Cr(l + 6 cos 0) (3) where subscripts i and e mean the inner and outer race, respectively", + " In general, it is concluded that the load capacity is only affected by the effect of squeeze film, the variation of friction is very small compared with load variation, and that the average value of friction for variable load is nearly equal to that of friction under the constant basic load which has the average value of variable load. T h e Theory With Consideration of Sliding of Roller Denoting M as the moment operating on a roller, M, as frictional moment caused by the friction between roller and raceway, and Me as the moment caused by the friction between roller and cage, as shown in Fig. 1, the following relation exists fif = UiW I{ VI + T cos 8 3 K, dd + C } + Kc )d0 (28) V 1 + e cos 6 where C is an integral constant. This can be rewritten by C f u r (29) (21) In the steady state, Fc = Ft = F and M, = 2FR. From eq. (64) in Part I This is determined with the boundary condition that Ur = Uro at 8 = 8o. Then, (30) where qo is a value of q at 8 = do. Considering the range of (6 \u2014 do) < 2t, since e' is the increasing function, there is the relation of e-i I e\" dO < 2T. The value of R2ICJIk,W is practi-J 0 o cally very small except the case in extremely low speed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003787_j.ymssp.2019.106415-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003787_j.ymssp.2019.106415-Figure3-1.png", + "caption": "Fig. 3. 1-D contact model with normal load variation.", + "texts": [ + "106415 The strategy used to build the reduced basis of Uiw is later discussed in Section 4, since it strictly depends on the physics behind the contact phenomena occurring at the hooks. The one-dimensional Jenkins contact element with normal load variation is used to compute the periodic contact forces for a given periodic relative displacement, by taking into account possible separation of the contact interfaces [6]. The Jenkins contact element models three different contact states: stick, slip and separation. A schematic view of this contact model can be found in Fig. 3, where the two dimensional relative displacement is decomposed into two perpendicular directions: two in-plane tangential component denoted by the u and w components, and one out-of-plane normal component v. The contact model\u2019s parameters are represented by the tangential and normal contact stiffnesses, kt and kn respectively, the coefficient of friction l and the normal preload f 0 (see Fig. 3). At every time instant the normal contact force f n\u00f0t\u00de is defined as: Please turbin f n \u00bc max\u00f0f 0 \u00fe kn v ;0\u00de \u00f027\u00de If f 0 is positive, the bodies are in contact before vibration starts, while if f 0 is negative an initial gap g0 \u00bc f 0 kn exists between the two bodies. Along the tangential direction, the contact force is defined as: f t \u00bc kt \u00f0u w\u00de sticking mode sgn\u00f0 _w\u00de lf n slipping mode 0 lift-off mode 8>< >: \u00f028\u00de Since the EQM are expressed in the frequency domain, the Fourier coefficients of the non-linear contact forces Fvhnl have to be computed from the Fourier coefficients of the corresponding relative displacement Xrel" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002816_1.4041364-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002816_1.4041364-Figure1-1.png", + "caption": "Fig. 1 A schematic sketch of the direct shear test conducted on the gel block and fixed glass plate for at a constant velocity (v0)", + "texts": [ + " Quantity of the gelatin powder and distilled water is varied in order to prepare the gelatin hydrogel \u00f0c \u00bc 10% wt:=vol:\u00de of different specimen thickness \u00f0h \u00bc 5; 6; 7; 9; and 12 mm\u00de. Shear modulus \u00f0Gs\u00de of hydrogel is estimated from the initial slope of shear stress versus shear strain plots which were obtained in shear sliding experiments [16\u201320]. Moreover, elastic modulus \u00f0E\u00de for the incompressible solids having Poisson\u2019s ratio 0:5 is evaluated using the expression \u00f0E \u00bc 3Gs\u00de. In the present experiments, Gs and E of the gelatin hydrogel \u00f0c \u00bc 10%\u00de are found to be 9:72 and 29:16 kPa, respectively. Figure 1 presents a schematic diagram of the direct shear test which is conducted on the hydrogel. A hydrogel specimen having contact area \u00f021 25 mm2\u00de and a fixed thickness is placed gently on the glass substrate. A thin aluminum plate is attached to the top surface of specimen using a commercially available adhesive. Top plate is pulled using a linear direct current servomotor at a constant velocity \u00f0v0\u00de which varies between 0:5 and 2 mm=s. The normal force (P) is applied by placing different weights \u00f00:5 0:8 N\u00de on the top of the plate" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002871_s10118-019-2180-9-Figure10-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002871_s10118-019-2180-9-Figure10-1.png", + "caption": "Fig. 10\u00a0\u00a0\u00a0\u00a0The\u00a0flat\u00a0bar\u00a0with\u00a0a\u00a0hole\u00a0in\u00a0tension\u00a0FE\u00a0simulation:\u00a0(a)\u00a0The\u00a0coordinate\u00a0system and\u00a0characteristic\u00a0points;\u00a0(b)\u00a0The\u00a0FE\u00a0mesh\u00a0(C3D8\u00a0elements)", + "texts": [ + " \u00a0It \u00a0was found\u00a0that\u00a0in\u00a0the\u00a0considered\u00a0case\u00a0of\u00a0uniaxial\u00a0stress\u00a0states,\u00a0the differences \u00a0between \u00a0results \u00a0obtained \u00a0for \u00a0the \u00a0two \u00a0values \u00a0of volumetric \u00a0modulus \u00a0were \u00a0negligible. \u00a0This \u00a0fact \u00a0supports \u00a0the proposed\u00a0method\u00a0of\u00a0material\u00a0parameter\u00a0identification,\u00a0which initially\u00a0makes\u00a0use\u00a0of\u00a0the\u00a0material\u00a0incompressibility\u00a0assumption\u00a0(with\u00a0the\u00a0true\u00a0value\u00a0of\u00a0the\u00a0volumetric\u00a0constant\u00a0being\u00a0calculated\u00a0in\u00a0the\u00a0next\u00a0step). Simulation of Flat Bar Deformation A\u00a0simulation\u00a0of\u00a0a\u00a0flat\u00a0bar\u00a0in\u00a0tension\u00a0is\u00a0described\u00a0below\u00a0as\u00a0an example \u00a0of \u00a0nonhomogeneous \u00a0deformation. \u00a0Due\u00a0to \u00a0the \u00a0symmetries \u00a0in \u00a0this \u00a0case, \u00a0only \u00a0one-eighth \u00a0of \u00a0the \u00a0bar\u2019s \u00a0geometry was\u00a0considered\u00a0(Fig.\u00a010a).\u00a0The\u00a0bar\u2019s\u00a0dimensions\u00a0are\u00a070\u00a0mm \u00d7\u00a070\u00a0mm\u00a0\u00d7\u00a05\u00a0mm.\u00a0There\u00a0is\u00a0a\u00a0centrally\u00a0placed\u00a0hole\u00a0with\u00a0the radius \u00a0of \u00a010 \u00a0mm. \u00a0The \u00a0following \u00a0boundary \u00a0conditions \u00a0were defined:\u00a0u1\u00a0=\u00a00\u00a0on\u00a0the\u00a0face\u00a0AA'EE',\u00a0u2\u00a0=\u00a00\u00a0on\u00a0the\u00a0face\u00a0BB'CC', u3\u00a0=\u00a00\u00a0on\u00a0the\u00a0face\u00a0A'B'C'D'E'. \u00a0The\u00a0displacement\u00a0component u2\u00a0 defined \u00a0on \u00a0the \u00a0face\u00a0DD'EE'\u00a0 was \u00a0used \u00a0in \u00a0the \u00a0form \u00a0of \u00a0a kinematic \u00a0excitation. \u00a0The \u00a0displacement \u00a0increases \u00a0linearly until\u00a0the\u00a0maximum\u00a0value\u00a0of\u00a03.5\u00a0mm\u00a0is\u00a0reached.\u00a0Subsequently,\u00a0it\u00a0changes\u00a0with\u00a0time\u00a0according\u00a0to\u00a0a\u00a0triangular\u00a0wave\u00a0function. \u00a0The\u00a0bar\u2019s\u00a0geometry\u00a0was\u00a0meshed\u00a0using\u00a0C3D8\u00a0elements (Fig.\u00a010b).\u00a0In\u00a0Fig.\u00a011(a),\u00a0the\u00a0contour\u00a0plot\u00a0of\u00a0Huber-von\u00a0MisesHencky\u00a0(HMH)\u00a0equivalent\u00a0stress\u00a0at\u00a0the\u00a0time\u00a0instant\u00a0t\u00a0=\u00a060\u00a0s is\u00a0depicted.\u00a0The\u00a0distribution\u00a0of\u00a0displacement\u2019s\u00a0magnitude\u00a0at the\u00a0same\u00a0time\u00a0instant\u00a0can\u00a0be\u00a0observed\u00a0in\u00a0Fig.\u00a011(b).\u00a0The\u00a0time histories\u00a0of\u00a0the\u00a0excitation\u00a0and\u00a0the\u00a0maximum\u00a0HMH\u00a0stress\u00a0are shown \u00a0in\u00a0 Fig. \u00a012. \u00a0The \u00a0simulation \u00a0was \u00a0performed \u00a0using \u00a0the constant\u00a0values\u00a0gathered\u00a0in\u00a0set\u00a02\u00a0(Table\u00a02)\u00a0and\u00a0the\u00a0volumetric modulus\u00a0D1\u00a0=\u00a06.74\u00a0\u00d7\u00a010\u22124\u00a0MPa\u22121. \u00a0 https://doi.org/10.1007/s10118-019-2180-9 \u00a0 CONCLUSIONS In\u00a0this\u00a0work\u00a0the\u00a0inelastic\u00a0mechanical\u00a0properties\u00a0of\u00a0polyamide 66 \u00a0were \u00a0investigated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002379_1350650117748096-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002379_1350650117748096-Figure1-1.png", + "caption": "Figure 1. Schematic configuration of externally pressurized double-layered porous journal bearing.", + "texts": [ + " The present analysis bears significance in terms of dealing with a modified form of Darcy\u2019s law19 valid for the flow of coupled-stress fluid through the porous medium with the effect of percolation of additives\u2019 microstructure into the pores. Using the steady-state film pressure distribution, the nondimensional load capacity, attitude angle, end flow, and frictional parameter have been obtained at various parametric conditions and the influence of these parameters on the bearing performance has been investigated and exhibited in the form of graph. A schematic configuration of externally pressurized double-layered porous journal bearing is shown in Figure 1. It is assumed that the flow of fluid through the porous medium in presence of percolation of additives into pores is purely viscous and obeys the Darcy\u2019s law.17,20 The governing equations of steady-state pressures in the porous mediums incorporating the additives effects into the pores of an anisotropic porous journal bearing are expressed as follows. For the coarse layer kxc kyc 1 yc 1 xc @2p0c @x2 \u00fe @2p0c @y2 \u00fe kzc kyc 1 yc 1 zc @2p0c @z2 \u00bc 0 \u00f01\u00de For the fine layer kxf kyf 1 yf 1 xf @2p0f @x2 \u00fe @2p0f @y2 \u00fe kzf kyf 1 yf 1 zf @2p0f @z2 \u00bc 0 \u00f02\u00de Modified Reynold\u2019s equation appropriate to the film region of an anisotropic porous journal bearing lubricated with the couple-stressed fluid with the effect of velocity slip at the fine porous layer\u2013film interface is represented by the following equation @ @x f h, kxf, l, xf @p @x \u00fe @ @z f h, kzf, l, zf @p @z \u00bc 6 U @ @x h 1\u00fe 0x\u00f0 \u00de \u00fe 12kyf 1 yf @p0f @y y\u00bc0 \u00f03\u00de Here f h, knf, l, nf \u00bc h3 1\u00fe n 1 nf ( ) 6h2l 0n tanh h 2l 12l2 h 2l tanh h 2l 0n \u00bc ffiffiffiffiffiffi knf p h\u00fe ffiffiffiffi knf p n \u00bc 3 ffiffiffiffiffiffi knf p 2 ffiffiffiffiffiffi knf p \u00fe h 1 nf h h\u00fe ffiffiffiffi knf p n \u00bc x, z With the help of following substitutions \u00bc x R , z \u00bc 2z L , y \u00bc y H , h \u00bc h C l \u00bc l C , Kxf \u00bc kxf kyf , Kxc \u00bc kxc kyc , Kzf \u00bc kzf kyf Kzc \u00bc kzc kyc , xf \u00bc Cffiffiffiffiffiffi kxf p , zf \u00bc Cffiffiffiffiffiffi kzf p xf \u00bc yf Kxf , zf \u00bc yf Kzf , p0f \u00bc p0f ps , p \u00bc p ps The nondimensional forms of equations (1), (2), and (3) are as follows xcK 2 xc @2 p0c @ 2 \u00fe R H 2@2 p0c @ y2 \u00fe D L 2 zcK 2 zc @2 p0c @ z2 \u00bc 0 \u00f04\u00de xfK 2 xf @2 p0f @ 2 \u00fe R H 2@2 p0f @ y2 \u00fe D L 2 zfK 2 zf @2 p0f @ z2 \u00bc 0 \u00f05\u00de And @ @ f h, xf, l, xf @ p @ \u00fe D L 2 @ @ z f h, zf, l, zf @ p @ z \u00bc S @ @ h 1\u00fe 0x\u00f0 \u00de \u00fe 1 yf @p0f @ y y\u00bc0 \u00f06\u00de Here f h, nf, l, nf \u00bc h3 1\u00fe n 1 nf ( ) 6 h2 l 0n tanh h 2 l " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000945_ssd.2010.5585533-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000945_ssd.2010.5585533-Figure8-1.png", + "caption": "Figure 8. CPAES rotor study domain considered in FEA. Legend: (4) magnetic ring, (5) claw, (6) unmagnetic core.", + "texts": [], + "surrounding_texts": [ + "its DC-excitation winding is located in the stator rather than in the rotor in conventional machines. The proposed model takes into account for the saturation of the mag netic circuit. A special attention is paid to the distribution of the air gap flux density and the back-EMF. A validation of the results yielded by FEA is achieved considering both measurements carried out on a prototype of the studied machine and analytical results obtained by a reluctance model developed in a previous work.\nIndex Terms- Claw pole alternator, DC-excitation in the stator, 3D finite element analysis, flux path, air gap flux density, back-EMF, validation.\n1. INTRODUCTION\nDuring the last four decades, talking about automotive power generation, it is quite commonly dealing with claw pole alternators [1, 2, 3]. The popularity of such electric machine is mainly due to its heteropolar topology offering the possibility of integration of a high pole pair number (6 to 10) in a low volume, leading to high torque densities and interesting generation capabilities.\nHowever, the machine is penalized by a crucial main tenance problem due to brush-ring system through which the rotor field winding is fed by DC current. An approach to discard this drawback and hence increase the availabil ity of the machine has been reported in the literature [4, 5]. Basically, this approach consists in transferring the DC excitation winding from rotor to stator, yielding the so called \"Claw Pole Alternator with DC Excitation in the Stator\" (CPAES). Previous investigations of the CPAES have highlighted interesting features [5, 6]. These have been carried out considering reluctance models. However, accounting for the complicated magnetic phenomena involved in the CPAES, it is expected that the results yielded by reluctance models are more or less accurate. Therefore, finite ele ment analysis (FEA) tum to be a necessity for the sake of a deeper investigation of the CPAES features. The present paper develops this idea.\n978-1-4244-7534-6/10/$26.00 \u00a920 1 0 IEEE\nThe CPEAS concept presents a claw pole topology where the DC-excitation winding is located in the stator rather than in the rotor as in conventional claw pole machines. As a result, the brush-ring system and the associated mainte nance problem have been discarded, which represents cru cial cost and availability benefits.\nThe machine is equipped by a three phase armature winding. In the manner of conventional claw pole alterna tors, the stator is made up of a laminated cylindrical mag netic circuit as shown in figure 1.\nThe field winding is simply wound in a ring shape. Figure 2 shows the photo of one half of the stator field winding.", + "The two halves of the field winding are inserted in both sides of the machine between the armature end-windings and the housing as illustrated in figure 3. They are con nected in series in such a way to produce additive fluxes.\nFollowing the transfer of the DC-excitation winding from rotor to stator, appropriate changes of the magnetic circuit have been introduced. These concerned mainly the rotor where the two iron plates with overlapped claws facing the air gap tum to be magnetically decoupled. Figure 4 shows a photo of the rotor of the CPAES.\nFollowing the removal of the field winding from rotor to stator and for the sake of an efficient flow of the flux, two magnetic collectors have been included in the stator mag netic circuit. These guarantee the flux linkage between rotor and stator. They are embedded on the two flasks of the machine. Figure 5 shows a photo of one magnetic col lector.\nThe CPAES static and rotating components are illustrated in figure 6. One can notice that each claw plate includes six poles. Moreover it is to be noted that the two claw plates are magnetically decoupled.\nIn order to reduce the computation time, the FEA study domain is limited to a one pair of poles of the CPAES. Figures 7 and 8 show the stator and the rotor study domains, respectively.", + "The meshing of both stator and rotor study domains has been automatically generated in the Ansys-Workbench environment.\nFigure 9 shows a mesh of the stator yoke and of the two associated magnetic collectors.\nA mesh of the stator lamination is illustrated in\nfigure 10.\nFigure 11 shows a mesh of the rotor claws and the as sociated magnetic rings.\n3.2.1. Main Flux Paths\nThe flux moves through the magnetic circuit of the CPAES within three dimensional (3D) paths. These are described as follows:\n\u2022 axially in the stator yoke (see figure 12),\n\u2022 radially down through the collector and crossing the\nair gap down to the magnetic ring (see figure 12),\n\u2022 radially then axially in the magnetic ring (see figure 12),\n\u2022 axially-radially in the claws (see figure 13),\n\u2022 radially up in air gap facing the stator laminations\n(see figure 14),\n\u2022 radially up in the stator teeth and circumferentially\nin the stator core back (see figure 15),\n\u2022 radially down in stator teeth facing the two adjacent\nclaws (see figure 16),\n\u2022 radially-axially in the adjacent claws (see figure 16),\n\u2022 axially then radially in the magnetic ring\n(see figure 17),\n\u2022 radially up through the air gap and the mag\nnetic collector on the other side of the machine (see figure 17),\n\u2022 axially in the other side of the stator yoke\n(see figure 18).\n3.2.2. Leakage Flux Paths\nNot all the flux produced by the excitation winding and the armature contributes to the EMF generation. Different leakage fluxes have been distinguished in the CPAES. The two main ones are\n\u2022 the leakage flux linking adjacent claws,\n\u2022 tow dimensional flux paths which flow through the\nmagnetic circuit as follows:\n- axially in the stator yoke,\n- radially down through the magnetic collector and air gap el,\n- radially then axially in the magnetic rings holding the claws,\n- axially in the claw,\n- radially in air gap e2 and in the stator teeth.\nFigure 19 illustrates the flux vectors flowing within 20 paths." + ] + }, + { + "image_filename": "designv11_33_0002077_ffe.12654-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002077_ffe.12654-Figure1-1.png", + "caption": "FIGURE 1 Cross section of the nacelle (1, blade; 2, spin; 3, hub; 4, bearings; 5, axle pin; 6, generator; 7, main column; 8, axial bearing; 9, column; 10, nacelle). [Colour figure can be viewed at wileyonlinelibrary. com]", + "texts": [ + " Mechanically, the axle pin is the most critical part of the structure and is subjected to bending rather than torsion. The axle pin in the middle is connected to the hub with 2 bearings. The radial bearing and the larger axial\u2010radial bearing are mounted on the axle pin, at the tip and within the radial transition region of the axle pin, respectively. Three blades are mounted to the hub. Torsion is induced between the hub and the generator's rotor. The generator's stator is fixed onto the axle pin and the main carrier, as shown in the cross section of the wind\u2010turbine housing in Figure 1. The direct current is induced between the permanent magnets on the generator's rotor and the stator fixed on the axle pin. Therefore, the axle pin can be mechanically treated as a bending loaded part of the wind turbine. The axle is usually made of cast steel (grade G24CrMn6), cast by the gravity casting process. Figure 2A shows an axle pin made by the casting process and painted after thermomechanical treatment. The design of the axle pin shows changes in wall thicknesses and reduction of diameters", + " The idealized unique surface crack in the most critical region, due to conservatism, can be considered by the FITNET2 procedure, as shown in Figure 9B. Per the FITNET2 procedure, it is allowable to assume and perform the assessment for the worst\u2010case crack geometry. This conservative approach leads to conservative results. An idealization of the crack geometry is shown in Figure 9B, where a = 20 mm is the crack depth and 150\u00b0 is the radial surface crack length. The wind\u2010power turbine's axle pin is loaded by bending, as 2 bearings are mounted onto the pin, shown schematically in Figure 1. Therefore, the bearings e of the crack. B, The idealized unique surface crack in the most critical on the pin ensure rotation of the hub with a direct drive to the direct current generator, while on the axle pin, the stator generator was fixed, as shown in Figure 1. The stress intensity factor is calculated via Equation 2 below: KI \u00bc \u03c3\u22c5 ffiffiffiffiffiffiffi \u03c0\u22c5a p \u22c5Y a T ; \u03b2 ; (2) where T is the thickness of the axle pin, Y(a/T,\u03b2) is the stress intensity function, a/T is the crack depth ratio, \u03b2 is the surface crack length in radians, and \u03c3 is the principal opening load, obtained by FE modelling. The stress intensity function for a hollow thick tube is available in the compendia of the stress intensity factors for FITNET7 procedure. Y a T ; \u03b2 \u00bc 1:1\u00fe a T \u22120:09967\u00fe 5:0057\u22c5 a T \u22c5 \u03b2 \u03c0 0:565 \" \u22122:8329 \u22c5 a T \u22c5 \u03b2 \u03c0 # (3) The distribution of the crack opening stress along the axle is shown in Figure 8" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002902_00423114.2018.1531135-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002902_00423114.2018.1531135-Figure9-1.png", + "caption": "Figure 9. Two-line vectors of instantaneous screws.", + "texts": [ + " For the spring with the stiffness ks, the relation fs = \u2212ks\u03b4ls can be substituted into Equation (8) to obtain w\u0302 = f r\u0302 = \u2212ks\u03b4ls r\u0302Ts S\u0302H r\u0302TS\u0302H r\u0302. (39) Figure 8. Kinestatic relations of suspensionmechanism: (a) double-wishbone suspension and (b)multilink suspension. If the relative vertical displacement of the wheel with respect to the vehicle body is defined as z\u2032 w (see Figure 8), the equivalent suspension rate Keq can be determined from Keq = \u2212 f z\u2032w = ks\u03b4ls z\u2032 w r\u0302Ts S\u0302H r\u0302TS\u0302H . (40) In Figure 9, the instantaneous twist T\u0302H of the wheel with respect to the vehicle body can be expressed in Pl\u00fccker\u2019s axis coordinates as T\u0302H = q\u0307H S\u0302H = q\u0307H [ rH \u00d7 sH + hsH sH ] , where sH is the unit direction vector of S\u0302H , rH is the position vector from the origin O to the axis of S\u0302H , and h is a pitch. The velocity vC of the wheel centre C can be determined by vC = q\u0307H[sH \u00d7 (rC \u2212 rH)+ hsH], (41) where rC is the position vector of the wheel centre C. The vertical velocity z\u0307\u2032w of the wheel centre C can be found by projecting vC onto the vertical axis, s = [ 0 0 1 ]T, z\u0307\u2032w = sTvC" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002076_optim.2017.7975006-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002076_optim.2017.7975006-Figure2-1.png", + "caption": "Fig. 2. Coils of the phase U, the elementary coil object of the one-phase short-circuit and the circuit associated to the field model", + "texts": [ + " The finite element analysis in time domain of the electromagnetic field gives the time dependence of the three stator currents and of different components of the magnetic flux density in points outside the motor. The value 0.05 ms of the time step considered by the step by step in time domain computation ensures a good enough precision of the numerical results for the evaluations of the amplitude of harmonics of the currents and of the magnetic field until 2000 Hz. An electric circuit is associated to the field model of the motor in order to take into account the three phase 380 V rms voltage supply of the motor and to model the short-circuits through low resistance resistors. The upper image in Fig. 2 highlights in yellow the first elementary coil of the phase U, red colored in Fig. 1, which is the object of the one-phase short-circuit. This short-circuit is modeled through the resistor Rshc in the lower image of the associated circuit, resistor which connects the input and output terminals of the short-circuited coil. A very high resistance 107 \u2126 of the Rshc resistor is considered for the healthy (HE) state of the motor operation. The values corresponding to faulty states with increasing fault severity FA1, FA2, FA3 and FA4 are 90 \u2126, 9 \u2126, 5 \u2126, and 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003534_iemdc.2019.8785093-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003534_iemdc.2019.8785093-Figure4-1.png", + "caption": "Fig. 4. Parasitic load components that can affect torque transducer output. The axial forces affect in the shaft direction, lateral forces in the radial direction, and bending moments are the force components trying to twist the shaft from the correct axial direction causing stress to the transducer.", + "texts": [ + " As the sensitivity tolerance is given as the permissible deviation of the actual sensitivity at rated load from the nominal sensitivity, the uncertainty caused by sensitivity tolerance is relative to the actual torque being measured [15]. For the example case, the sensitivity tolerance is ( ) = \u2219 \u2219 1 \u221a3 = 2 \u2219 239.7 Nm \u2219 0.0005 \u2219 \u221a = 0.14 Nm. (23) The last source of uncertainty defined in Table IV is the group of different parasitic loads (PARA) that cover the mechanical stresses other than the actual load torque subjected to the torque transducer. Fig. 4 illustrates the different parasitic load components that are axial force, lateral force, and bending moment. The axial force is the component affecting in the shaft direction that can be caused by inaccurate positioning of the transducer in the direction of the shaft, but also by the thermal expansion of the shaft. The lateral force, in turn, is the component in the radial direction that can naturally originate from inaccurate positioning in the perpendicular direction. The bending moment, as the name suggests, is the parasitic load component twisting the transducer shaft and is caused by inaccuracies in the angle of the transducer mounting" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003852_j.mechmachtheory.2019.103628-Figure19-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003852_j.mechmachtheory.2019.103628-Figure19-1.png", + "caption": "Fig. 19. The ground tooth surfaces.", + "texts": [ + "538 mm and 65 mm, respectively. Then the grinding wheel location of tooth grinding can be calculated according to the method described in Section 5 , which are shown in Figs. 17 and 18 . At the same time, the grinding wheel model (the green part) is built according to Table 4 , and the imported gear model (the orange part) is shown in Fig. 10. There is no feeding during the first grinding, and the grinding wheel should be tangent to the tooth surface theoretically. The actual ground tooth surfaces are shown in Fig. 19(a) , which shows that the grinding wheel and the tooth surface are in intermittent contact state, and the ground area extends from the top of the tooth to the root of the tooth evenly at the contact positions. This is because the material removal simulation of machining simulation software is based on the discrete grid model, but not on Boolean operation of accurate 3D model. The intermittent grinding state can be considered as tangent contact between the grinding wheel and the tooth surface along the b curve, and there is no curvature interference. Then tooth grinding is carried out again after adjusting the gear rotational position (corresponding to 0.02 mm grinding feed), and the ground tooth surfaces are shown in Fig. 19(b) . It can be seen that a layer of material is removed evenly for both sides, which proves the correctness of the grinding path. This paper presents a new method for face-hobbed hypoid gear tooth grinding using large diameter conical grinding wheel. The analysis shows that the proposed approach can efficiently and accurately grind the FFHHG without theoretical tooth deviation. The main conclusions are summarized below: (1) In order to avoid curvature interference, the parameters of the grinding wheel working side for concave grinding should be determined by the principal curvature radius and its change rate at the root position of the concave toe" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000638_232002-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000638_232002-Figure2-1.png", + "caption": "Figure 2. Illustration of the profile at the mid-section of the cylinder (radius R, length L) in a 2-fingered experiment. Fingers indent the cylinder radially (dotted line) by a distance z (equal for both). The left (l+, l\u2212) and right (r\u2212, r+) pairs of vertices are shown. (+) and (\u2212), respectively indicate the far and near points from the plane of symmetry (chain line). We measured forces Fl and Fr with two load cells (Cooper LPM 512). The profile was determined by laser sheet tomography and is representative of the geometry of the deformation of the cylinder due to symmetry. A tilted mirror located inside the cylinder reflected the image of the profile towards a camera that was placed outside the transparent cylinder. The cylinder was constructed from a flat transparent 2\u03c0R \u00d7 L PVC sheet. The cylinder was clamped at the ends in such a way that air circulation was assured and the cylinder was free of tension except for its own weight. The thickness, h = (100 \u00b1 5) \u00b5m, and aspect ratio, L/R = 4.58, were constant, while R = 50, 75, 100 and 125 mm. The bending stiffness of the sheets was in the range (2\u20134) \u00d7 10\u22125 N m.", + "texts": [ + " The procedure can be illustrated by squeezing an aluminum soda can with two fingers as shown in figures 1(a)\u2013(d). Eventually the shallow depressions that form merge through a snap-buckling instability and leave a swirled scar, which is simply the track left by two facing, interacting vertex singularities. In the experiment, vertices are nucleated pairwise by indenting a cylindrical shell (cylinder, from now on) radially at its midplane [7, 10]. An illustration of this procedure is given in figures 1(e)\u2013(f ). Figure 2 shows the configuration relevant for this experiment in which two pointed rigid fingers are separated by an angle 2\u03b1. They move slowly inwards, radially indenting the cylinder and nucleating two pairs of vertices (l+, l\u2212) and (r\u2212, r+). By increasing the pushing distance z (the same for both fingers) the vertices of each pair move away from one other. At all stages the vertices and the tips of the fingers lie in the same azimuthal plane. For small values of \u03b1 we observe that the inner vertices l\u2212 and r\u2212 mutually annihilate at a critical indentation distance zc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001575_robot.2010.5509764-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001575_robot.2010.5509764-Figure5-1.png", + "caption": "Fig. 5 QuadHelix CAD prototype", + "texts": [ + "1) So two rope ends of the first DoHelix-coiling (5) are guided over two pulleys (7) and are pulling the turning wheel (8) into one direction, while the other two rope ends of the second DoHelix-coiling (6) are working in the opposite way. With this new layout, named QuadHelix-Drive because of its usage of two DoHelix-coilings, an improved actuator system is possible. IV. MECHANICAL REALIZATION The next step towards a new system is the mechanical realization. To show the potential of the actuation concept a CAD prototype is drafted, as shown in Fig. 5. This prototype consists of a 200 W motor unit with a small gear box (a), with a reduction rate between 4.3:1 for the testing facility and 61:1 for a 2-DoF-module in a robotic arm. A long shaft with 6 mm diameter (b) with a worm gear (c), two 1.5 mm diameter DoHelix ropes (d) with a breaking load of 2200 N and two aligned 100 mm turning wheels as representations of a 1-DoF-axis (e) are other key components. In addition, the motor unit is placed on a linear guiding rod (f), the worm gear has a fixed linear gear rod as a counterpart (g) and the two DoHelix ropes are guided to the aligned turning wheels by eight pulleys (h)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000041_cae.20391-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000041_cae.20391-Figure6-1.png", + "caption": "Figure 6 A 3D model of the processing station.", + "texts": [ + " Transfer cylinder carries parts by means of vacuum handling and a pressure sensor also controls the vacuum. Part handling by means of vacuum between stations modules was realized via Cartesian axes movements and could be called as a Cartesian Robot. X- and Y-axis are obtained by using two DC gear motors (D.AS-DC-G-24-2) and two slide units, spindle drive (D.ERKSP-250) and Z axes movement is given by the pneumatic transfer cylinder [11]. Processing Station consists of four modules, turntable module, marking module, drilling module, and cleaning module as shown in Figure 6. On the turntable, parts were marked, drilled, and cleaned by sequence at marking module, drilling module, and cleaning module, respectively. The turntable was driven by a 24 V DC motor and positioned by a relay circuit at processing steps, with the position of the being detected by an inductive sensor signal connected to the turntable body. The transfer cylinder retrieves the differentiated parts from testing module and locates to the loading/retrieving position of the turntable. After loading the material, turntable turns according to the material type. For plastic parts, marking, drilling, and cleaning operations were required. For metal parts only marking and cleaning operations were required. Transfer cylinder retrieves the processed parts from loading/retrieving position and locates to the sorting magazine according to the material type as metal and plastic. Turntable and processing modules picture is shown in Figure 6. Marking operation was modeled by a single acting pneumatic cylinder. Drilling operation was simulated with a cylinder and an electrical motor and cleaning process was simulated by a nozzle expelling compressed air to the materials. Movements and position names are given in Figure 7. Transfer times of the materials in the MTMPS unit with respect to positions are given in Table 1. Since the plastics part is required drilling operation, the average total time for plastic part is higher than metal average total time" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002938_icelmach.2018.8507253-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002938_icelmach.2018.8507253-Figure1-1.png", + "caption": "Fig. 1. Description of the studied PM machines: (a) without pole gap (M1), (b) with pole gap (M2), (c) with pole filler (M3), and (d) Inset PM (M4).", + "texts": [ + " These machines have the same 36-slots stator with distributed winding. Three machines have surfacemounted PM (SPM) rotor structure with a carbon-fiber retaining sleeve. The other one have an inset PM rotor structure with a carbon-fiber sleeve. The rotor diameter and the PM thickness are the same. The studied machines have different PM opening angle to pole pitch ratios (1 and 0.8) and different inter-pole filler materials (glass fiber, air and iron). The different rotor structures are depicted in Fig. 1 and the main geometrical parameters are reported in Table II. The stator laminations are made of NO20 steel sheets, the PMs are SmCo Recoma32 [10] and solid rotor has been adopted. Electromagnetic, Structural and Thermal Analyses of High-Speed PM Machines for Aircraft Application R. Benlamine, T. Hamiti, F. Vangraefsch\u00e8pe, and D. Lhotellier H 978-1-5386-2477-7/18/$31.00 \u00a92018 IEEE 212 The machines have been analyzed for the transient and the nominal operating points: 60 Nm-10,000 rpm and 100 kW-60,000 rpm, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003861_peami.2019.8915294-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003861_peami.2019.8915294-Figure4-1.png", + "caption": "Fig. 4. Partitioning into finite elements", + "texts": [ + " The emission of the electronic circuit is built using the Maxwell Circuit program, which is an application of the program. The fig.3 shows a three-phase rectifier and voltage regulator that controls the excitation winding to stabilize the output voltage of the generator. To solve the problem, the program splits the model into a large number of finite elements. For each element, a system of equations for the local matrix is formed. A global matrix is formed from the local matrices ifor solution of the field problem . The results of splitting the model into finite elements are shown in fig. 4. The results of the calculation of the magnetic field are shown in fig.5 The fig.6\u201310 show the results of the calculation of the main parameters and characteristics The analysis of curves shows that the created model is close to the real generator in terms of calculation accuracy. VI. DISCUSSION The proposed generator has a complex magnetic system, which contains two sources of magnetic field. The ways of closing the magnetic flux from the permanent magnets and the excitation winding have a complex shape" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003301_10402004.2019.1607641-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003301_10402004.2019.1607641-Figure6-1.png", + "caption": "Fig 6 Schematic diagram of experimental setup (a) and test rig (b)", + "texts": [ + " 5 illustrates the schematic diagram of the measurement method and pictures of the test device. The test device has a short shaft, split in half lengthwise. One half is rigidly fixed and the other is mounted on the shaft, which is movable and can induce pressure on the force transducer. The test plant is capable of operating in a wide temperature range (from room temperature to300\u00b0C). The bench test was used to measure the reverse pumping rate and the friction torque in order to validate the numerical simulation results. Fig. 6 shows a schematic diagram of the experimental setup and pictures of the test rig, respectively. To obtain the reverse pumping rate, the seal was installed reversely. The air side and the oil side exchange in place. To avoid confusion, the \u201cair side\u201d was termed the exterior side in this paper, which will be referred later. Consequently, the leakage rate measured by this configuration is equivalent to the pumping rate under the normal operating conditions. The leaking lubricant oil was collected by a beaker and weighed using a precision Acc ep te d M an us cr ipt electronic scale with a resolution of 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003536_bs.afnr.2019.07.002-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003536_bs.afnr.2019.07.002-Figure4-1.png", + "caption": "Fig. 4 The experimental set-up used for the formation of stabilized lipid films on glass fiber filters; the micromachined chambers are separated by a thin (12.5\u03bcm thick) polyvinylidene chloride wrap and enclose the microfiber disk. For more details, see text. From ref. Nikoleli, G.-P.; Nikolelis, D.; Siontorou, C. G.; & Karapetis, S. (2018). Lipid membrane nanosensors for environmental monitoring: The art, the opportunities, and the challenges. Sensors, 18(1), 284.", + "texts": [ + " An Ag/AgCl reference electrode was positioned at the center of the cylindrical cell. An external voltage of 25 or 50mV d.c. was applied between the two reference electrodes. A Keithley digital electrometer was used as a current-to-voltage converter. A peristaltic pump was used for the flow of the carrier electrolyte. Sample injections were made with a Hamilton repeating dispenser. The electrochemical cell and electronic equipment were isolated in a grounded Faraday cage. A simple scheme of the apparatus used is presented in Fig. 4. Details for the procedure followed for the formation of the stabilized BLMs can be found in (Andreou & Nikolelis, 1998; Nikolelis et al., 1995). The methods of polymer stabilized in air lipid film devices were given in previous reports (Nikolelis, Raftopoulou, Chatzigeorgiou, Nikoleli, & Viras, 2008; Nikolelis, Raftopoulou, Nikoleli, & Simantiraki, 2006). These techniques include polymerization using UV irradiation and 8 Georgia-Paraskevi Nikoleli not heating at 60 \u00b0C because the latter deactivates the enzyme or antibody molecules" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003541_iemdc.2019.8785161-Figure10-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003541_iemdc.2019.8785161-Figure10-1.png", + "caption": "Fig. 10 Configuration of the improved magnetic-geared motor.", + "texts": [ + " However, the eddy current loss is large since the magnets facing to the air gap are affected by the slot harmonics. Thus, the IPM type that the magnets are embedded in the inner rotor core is employed for the improved machine. Fig. 9 shows the comparison of eddy current loss when a current density is 5 A/mm2. It is clear that the eddy current loss of the IPM type is reduced to 1/4 in comparison with the SPM type. The reason is that harmonic flux flows into the rotor core instead of the rotor magnets by embedding the magnets. Fig. 10 shows a configuration of the improved magneticgeared motor which is designed based on the investigation in the previous sections and Table II indicates specifications. The open-slot structure is employed in the improved machine as shown in the figure. As a result, the winding space factor is 46.1% which is about 9% higher than the previous one. The 6.5% Si steel is also employed in the stator and rotor core. In addition, the IPM type rotor is employed to reduce the eddy current loss of the magnets" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002248_s00170-017-1085-4-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002248_s00170-017-1085-4-Figure3-1.png", + "caption": "Fig. 3 Decomposed model of multiple robot arm system", + "texts": [ + " The relation between the joint torque and the joint angular deformation can be expressed as \u03c4 i \u00bc Kqidqi; i \u00bc 1; 2\u2026; n \u00f02\u00de where Kqi is the stiffness of the ith joint. In order to develop a static stiffness model of the entire multiple robot system, the mapping of the joint stiffness of each robot to the system stiffness needs to be established. For this reason, the coordinated robot system is decomposed into several single robot arms with the constraints in force transmission and geometry construction as shown in Fig. 3. For each single arm of this multiple robot system, there is a relation between the deflection associated with the center point on the workpiece and that associated with the end effector of each robot. Without loss of generality, the deformation model is established for the ith robot arm as DCi \u00bc I \u2212S O EiR Ei\u2022EiPCi 0 I \u2022DEi \u00f03\u00de where O EiR Ei is the rotation matrix from the frame {O} attached to the base to the frame {Ei} attached to the end effector of the ith robot;DCi andDEi are the deflection of pointCi (selected at the center of the workpiece) and point Ei (i", + " (23) can be rewritten as DCi \u00bc D\u03b8 Ci \u00fe DL Ci \u00bc K\u03b8 i \u22121 \u2022PCi \u00fe KL i \u22121 \u2022PCi \u00f027\u00de where K\u03b8 i K L i denotes the stiffness matrix of the joint stiffness and link stiffness respectively, Pciis the force exerted on point C from the end effector of ith robot. The overall stiffness of the ith robot can be obtained by the addition of compliance matrix [23] K\u22121 i \u00bc K\u03b8 i \u22121 \u00fe KL i \u22121 \u00f028\u00de To obtain the stiffness of the entire multiple robot system, it is necessary to establish the deflection equation for the object held by the robots. Here, the system is decomposed into n single robots as shown in Fig. 3. In the decomposed model, the observation point C is shared by all robots expressed as Ci (i = 1, 2,\u22ef, n). Since the same observation point is used for all robots, a parallel spring system as shown in Fig. 7 can be used under which the deflection that each robot undergoes would be identical, that is DC \u00bc DC1 \u00bc DC2\u22ef \u00bc DC n\u22121\u00f0 \u00de \u00bc DCn \u00f029\u00de Then the total force at point C should be the summation of all the robot\u2019s end effector forces contributing to point C, i.e., Kall\u2022Dc \u00bc K1\u2022DC1 \u00fe K2\u2022DC2\u22ef\u00fe Kn\u2022DCn \u00f030\u00de In light of Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002595_j.triboint.2018.05.003-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002595_j.triboint.2018.05.003-Figure5-1.png", + "caption": "Fig. 5. Film thickness in dynamic conditions, at the end of an oscillation cycle ( =t 3.33 ms in Ref. [1]). Contour plots 5(a) to 5(c) correspond to numerical examples 1 to 3 in Table 2. The line plots in 5(d) compares these with the experimental results from Ref. [1].", + "texts": [ + " Results for film thickness fluctuations in dynamic conditions are shown in Fig. 3 using case 3 in Table 2. These illustrate four different solution times during one ball sliding stroke covering half of the last simulated oscillation cycle as indicated by the red dots in Fig. 4. The periodic nature of the generated film fluctuations becomes apparent in these results when noting that the film shapes at the start and end of the ball sliding stroke mirror each other. The effect of the lubricant rheology is investigated in Fig. 5 based on the film fluctuations at the simulation time =t 9.99 ms. As illustrated by the last red dot in Fig. 4, this time corresponds to the end of the third speed oscillation cycle where the instantaneous transverse velocity of the ball equals zero. According to cases 1 to 3 in Table 2, the lubricant rheology is considered in Fig. 5(a) as Newtonian, in 5(b) as Eyring and in 5(c) as Carreau-Yasuda. These results correspond to the measurements made at =t 3.3 ms presented by Fryza et al. in their Fig. 6(a). Therefore, a direct comparison with the relevant film profile given in the reference is attempted here in Fig. 5(d). The location where these profiles are taken in the current results is indicated by the dashdotted lines in the contour maps. Qualitatively, there is a good agreement between the experiments and the numerical results for all three lubricant rheology types considered. All major local and global features of the lubricant film thickness distribution seen in the interferogram of Fig. 6(a) in Ref. [1] are well reproduced by the simulations. Quantitatively, however, some minor shortcomings are evidenced when comparing film profiles in Fig. 5(d). The simulated results are unable to accurately predict both local maxima and minima of the measured film thickness. This is in line with the observations made by Fryza et al., who noted that the peak film thickness values near the contact inlet were larger than the steady state, pure rolling values at the maximum effective entrainment speed. They pointed that this could be attributed to rig dynamics and squeeze film effects introduced by the rapid lateral oscillations of the ball. The results here obtained are not sufficient to provide an explanation to the above described deviations with regard to the measurements", + " And since the lubricant SN650 + PIP shows no appreciable shear thinning under these conditions, then when combined with the squeeze film effect, which delays film thickness decay when the effective entrainment speed decreases in an oscillation cycle, it leads to the observed net global film thickness increase. Conversely, the fact that the central film thickness decreases can be attributed to being a local evaluation in combination with the particular set of operating conditions. As seen in Fig. 3 here and also in Fig. 5 of the reference, the contact side lobes are stretched either towards or away from the contact center depending on the direction of the transverse speed, hence accounting for the film reduction observed there. Once this cyclic process has been identified it is then apparent that by using different combinations of frequency and stroke length for the transverse speed oscillation the side lobes may be stretched in ways that they might or might not fully reach the contact center point. Or, if the evaluation is fixed to a point other than the contact center then it is possible that the film thickness evolution shows an entirely different trend", + " For example, coming back to the results for SQL + PIP in Fig. 6(b) the normalized central film thickness oscillates around a value of \u2248 0.88. In their Fig. 10 the corresponding result with \u2248v u/ 1 is \u2248 0.7. The differences are thus substantial and can perhaps be explained noting that in the reference the film thickness is measured in the high pressure region of the contact using a single film thickness profile that is extracted from an interferogram taken at a time corresponding to the end of a full speed oscillation cycle, such as shown for example in Fig. 5 here. As also noted by Fryza et al., with such approach there is thus a risk of excluding relevant film fluctuations no longer present at the end of the speed oscillation cycle. The same can perhaps be found when analyzing combinations of frequency and stroke length other than those used here for the transverse oscillation speed. The film fluctuations created might have a wavelengths that is either much shorter or longer than the contact length, such that the missing or additional film thickness information leads to inaccurate evaluations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001673_sav-2010-0501-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001673_sav-2010-0501-Figure1-1.png", + "caption": "Fig. 1. Quarter vehicle McPherson suspension system.", + "texts": [ + " The excitation is analytically calculated from the power spectral density (PSD) acting on the lower suspension arm of a vehicle through a tire of stiffness constant Kp, where the return chain is made up of a linear shock absorber and a linear spring. To evaluate the dynamic behaviour of the suspension system and the roadway profile model, Rahnejat [13] studied the dynamics of the Macpherson suspension system in a quarter vehicle. The spring and the shock absorber are represented as a single element with stiffness constant k s (Fig. 1). In fact, Rahnejat proposes a simple model where the mass and inertia are represented by parameters m and I respectively. The model developed in this study is based on vertical motion and road excitation. In the model shown in Fig. 2, the partial stability of the vehicle is ensured by a suspension control system. The action chain of the system has a tire of stiffness constant kp. The return chain in negative feedback has a spring of stiffness ks assembled with a shock absorber Cs. The excitation force F2 of the road irregularity is balanced through the tire by the F1 negative feedback" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002839_j.mechmachtheory.2018.06.021-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002839_j.mechmachtheory.2018.06.021-Figure9-1.png", + "caption": "Fig. 9. Structural schematic of the outer cone CVT.", + "texts": [ + " (13) into the plane coordinate system XOY ( Fig. 8 ). As shown in Fig. 8 , the circle is suitable for contacting to an outer cone roller with a straight-lined generatrix, and a relative disc can be obtained by using the derived circle as its generatrix. During the speed ratio variation process, the circle and the cone apex of the roller need to be moved along the X -axis and. The roller is also tilted, as illustrated in Fig. 8 . By using the derived curves of Eq. (35) , a type of outer cone CVT can be designed, as shown in Fig. 9 . Input disc and output disc are placed on the left and right sides of the outer cone roller, respectively. The outer cone roller is assembled precisely such that its cone apex is on the disc rotation axis. The cone apex is required to move along the X -axis ( Fig. 8 ), but it can be fixed at a single point because the translational motions of the discs and roller are relative to one another. In other words, the rollers and the discs are required to have a tilting and translation motion ( Fig. 9 ) in the process of speed ratio variation, respectively. The formula of speed ratio that represents the ratio of input speed to output speed also needs to be derived. As shown in Fig. 9 , A is the contact point between the roller and the input disc, while B is the contact point between the roller and the output disc. Segments AA\u2019 and BB\u2019 are perpendicular to the disc rotation axis at A\u2019 and B\u2019 , respectively. Segments AA\u201d and BB\u201d are perpendicular to the roller rotation axis at A\u201d and B\u201d, respectively. The length of segment BB\u2019 can be derived from Eq. (34) . \u2223\u2223BB \u2032 \u2223\u2223 = R sin ( \u03b8 \u2212 \u03c9t ) , (36) where \u03b8 is the half-cone angle of roller, \u03c9 is the tilting angular speed of the roller rotation axis, R is the radius of the generatrix circle of disc and t is an individual parameter that represents time", + " (44) , we can obtain \u03c3spinL = cos ( \u03b8 \u2212 \u03b3 ) \u2212 cos ( \u03b8 \u2212 \u03b3 ) cos \u03b8 cos \u03b8 = 0 , (50) where \u03c3 spinL is the spin ratio of logarithmic CVT between the input disc and the roller. The logarithmic CVT can also transmit power without the spin motion. In accordance with Fig. 6 , the spin ratio of the inner cone CVT between the input disc and the roller can be written as \u03c3spinI = cos ( \u03b8+ \u03c9 \u00b7 t ) \u2212 \u03c9 r \u03c9 i sin \u03b8, (51) where \u03c3 spinI is the spin ratio of the inner cone CVT between the input disc and the roller. By substituting Eq. (28) into Eq. (51) , we obtain \u03c3spinI = cos ( \u03b8 + \u03c9 \u00b7 t ) \u2212 cos ( \u03b8 + \u03c9 \u00b7 t ) sin \u03b8 sin \u03b8 = 0 . (52) In accordance with Fig. 9 , the spin ratio of the outer cone CVT between the input disc and the roller can be written as \u03c3spinO = cos ( \u03b8+ \u03c9 \u00b7 t ) \u2212 \u03c9 r \u03c9 i sin \u03b8, (53) where \u03c3 spinO is the spin ratio of the outer cone CVT between the input disc and the roller. By substituting Eq. (41) into Eq. (53) , we derive \u03c3spinO = cos ( \u03b8 + \u03c9 \u00b7 t ) \u2212 cos ( \u03b8 + \u03c9 \u00b7 t ) sin \u03b8 sin \u03b8 = 0 . (54) The slip condition (i.e. the torque is transmitted) is only slightly different from the cases of Eqs. (47) , (50) , (52) and (54) because the slip is always extremely small" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000862_s1350-4789(10)70534-1-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000862_s1350-4789(10)70534-1-Figure6-1.png", + "caption": "Figure 6. A composite PTFE and elastomer seal designed for ease of fitting and flexibility.", + "texts": [ + " An immediate problem created is the ability to maintain a static leak, and a number of methods have been tried to achieve this.[24] Sometimes an elastomer is used in series to provide the leakage control. PTFE seals are now widely used for powertrain duties as they address the progressive increases in temperature combined with increasingly aggressive additives used in longlife lubricants. There are many variations aimed at providing convenient assembly and increased resilience by means of a composite assembly with elastomer flexible element (Figure 6). An alternative approach has been to use what appears to be a conventional elastomer seal, but with a PTFE coated lip. These have been shown to provide a reduction in friction,[23] but applications appear to be confined to rela- tively exotic duties, suggesting that comparative cost may be a deciding factor. A reduction in seal friction provides a double benefit as not only does it become more attractive because of reduced power consumption, but the reduction of under-lip temperature automatically offers a potential for extended service life under the same operating conditions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000761_j.proeng.2012.04.030-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000761_j.proeng.2012.04.030-Figure2-1.png", + "caption": "Fig. 2. Body-fixed coordinate system Fig. 3. Aerodynamic coefficients as functions of the angle of attack at 7 revolutions per second and 25 ms-1", + "texts": [ + " A self-organizing map (SOM) is a type of artificial neural network that trains a set of high-dimensional input data through an unsupervised learning process, and projects them onto a low-dimensional output map while preserving their own features. A SOM is useful for enabling low-dimensional views of high-dimensional data [2]. The inertial coordinate system is shown in Figure 1. The origin is defined as being at the center of the throwing circle, while the XE-axis is in the horizontal forward direction, the YE-axis is the horizontal lateral direction and the ZE-axis is vertically downward. The body-fixed coordinate system for the discus is shown in Figure 2. The origin is defined as being at the center of gravity of the discus. It is assumed that the geometric center of the discus coincides with its center of gravity. The xb and yb axes are each in the plane of the discus, and the zb-axis is orthogonal to these. Fig.1.Inertial coordinate system In terms of coordinate transformations [3] we then have W V U m Z Y X ij E E E (1) 2 3 2 2 2 1 2 010322031 1032 2 3 2 2 2 1 2 03021 20313021 2 3 2 2 2 1 2 0 22 22 22 qqqqqqqqqqqq qqqqqqqqqqqq qqqqqqqqqqqq mij (2) Here, (U, V, W) are the (xb, yb, zb) components of the velocity vector as shown in Figure 2. The Eulerangle transformation matrix [mij] is defined by the quaternion parameters (q0, q1, q2, q3) in Equation (2) [4]. The equations of motion and the moment equations are RVQWqqqqgmX m U da d 203121 (3) PWRUqqqqgmY m V da d 103221 (4) QUPVqqqqgmZ m W da d 2 3 2 2 2 1 2 0 1 (5) 1 L T L a I IQR I LP (6) L T L a I IRP I MQ 1 (7) T a I NR (8) Here, (Xa, Ya, Za) are the (xb, yb, zb) components of the aerodynamic force, (P, Q, R) are the (xb, yb, zb) components of the angular velocity vector, md is the mass of the discus, g is the gravitational acceleration, (La, Ma, Na) are the (xb, yb, zb) components of the aerodynamic moment, and IL and IT are the moments of inertia of the discus about its longitudinal (i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000680_125003-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000680_125003-Figure2-1.png", + "caption": "Figure 2. Schematic layout of a LIGO interferometer. The Fabry\u2013Perot arm cavities have lengths Lx Ly 4 km.", + "texts": [ + " Potential astrophysical sources of gravitational waves include supernovae and compact binary inspirals. The expected strain from such sources requires a detector sensitive to displacements on the order of 10\u221219 m/ \u221a Hz over a few km baseline, at frequencies ranging from \u223c50 Hz to 7 kHz. Figure 1 shows a typical displacement sensitivity for the LIGO interferometers. 0264-9381/12/215008+13$33.00 \u00a9 2012 IOP Publishing Ltd Printed in the UK & the USA 1 Using Fabry\u2013Perot cavities for the interferometer arms as well as power recycling [2] helps improve the detector\u2019s sensitivity. Figure 2 shows a simplified schematic of a LIGO interferometer. The motion of the ground is many orders of magnitude above the required sensitivity, so the mirrors are seismically isolated, using both passive and active techniques. The cavities are held on optical resonance by suppressing external disturbances with a family of length and alignment feedback control loops. Suspending the mirrors as pendulums provides passive vibration isolation above the pendular resonance, which is arranged to be \u223c1 Hz. The support point of the pendulum is attached to a passive in-vacuum seismic isolation stack, consisting of four mass and spring layers", + " Small magnets are glued to the backs of the mirrors to allow for actuation via magnetic fields generated by currents flowing in nearby wire coils. Above \u223c50 Hz, where ground vibrations have been sufficiently suppressed, the detector is fundamentally limited by thermal noise in the mirrors and seismic isolation systems and by photon shot noise. There are four length degrees of freedom that need to be controlled. They are defined in terms of differential and common length changes, L+ = Lx + Ly 2 (1a) L\u2212 = Lx \u2212 Ly (1b) l+ = lx + ly 2 (1c) l\u2212 = lx \u2212 ly (1d) where Lx, Ly, lx, and ly are as shown in figure 2. An RF modulation scheme [3] is used to generate control signals by relating position and angle fluctuations of the test mass mirrors to power fluctuations measured by photodetectors at various locations around the interferometer [4]. The detector is brought to and held at its operating point by sequentially bringing cavities into resonance using dynamically calculated error signals [5], a process called lock acquisition. Precision measurement of the differential arm length degree of freedom, called L\u2212 or DARM, enables the detector to find potential gravitational waves" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002779_1464419318789185-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002779_1464419318789185-Figure4-1.png", + "caption": "Figure 4. Three-dimensional FE model of the first gear pair for meshing impact force calculation: (a) constraints, boundary conditions, and loads and (b) contact relationship between the pair.", + "texts": [ + " Finally, all the nodes of the tooth mesh model are generated through axially extruding and simultaneously rotating all the nodes on the (front) end face. In this way, the structured mesh model of the tooth pair is obtained as shown in Figure 3(b). To obtain the impact forces under different rotational speed, a parameterized program is developed in Python language for automatic generation of the mesh model, which improves the calculation efficiency of finite element analysis (FEA). Meshing impact force model With the main parameters of the example EV gearbox listed in Tables 1 to 3, a three-dimensional model of the first helical pair (Figure 4) is constructed by the FEM in Abaqus for the meshing impact force calculation. First, the structured mesh models of the pinion teeth are accurately generated through the circular array of one tooth mesh model (Figure 3(a)) according to the number of teeth. Second, the pinion shaft (i.e. input shaft) is created in order to include the effect of the elastic deformation because it also influences the position of the initial impact point. Third, the teeth, the hub, and the shaft of the pinion are merged as a single component to simplify the FE model. In the same way, a simplified component of the gear is also obtained accordingly. Based on the gear meshing theory,24 both components are precisely assembled under a global coordinate system, as shown in Figure 4(a). Next, the deep groove ball bearings supporting the pinion shaft and the gear shaft are simplified as equivalent springs, one end of which is fixed on the ground and the other of which is coupled to the relevant surfaces of the shafts. Afterward, the surface-to-surface contact algorithm is applied to simulate the interaction between the pair. Then, an input rotational speed is given to the pinion shaft, and a load torque (equivalently transformed from an input torque acting on the input shaft) is applied to the gear shaft. Finally, with good grid quality, the FE mesh model of the entire pair is established to investigate the meshing impact force, as shown in Figure 4(b). Considering both the calculation efficiency and accuracy, there are 20 sampling points adopted in the impact time and 10 sampling points determined in a complete meshing period. In this way, there are totally 30 points selected to record relevant simulation results, such as the impact time and the amplitude of the impact force. The FE model of the pair is solved using implicit integration algorithm, and three meshing periods under steady-state conditions are investigated for the impact force calculation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001339_icar.2011.6088557-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001339_icar.2011.6088557-Figure2-1.png", + "caption": "Fig. 2. Rescue scenario in the TB: the green rescue vehicle must assist the red parked vehicle (left); path solution for the same path for both wheels (center); independent paths for each wheel (right).", + "texts": [ + "00 \u00a92011 IEEE 106 The motion planning problem for the CPRHS rhombic vehicle was first addressed in [1] by a planning methodology, which granted the generation of feasible and optimized trajectories in agreement with the above specified criteria. The inbuilt path optimization module followed closely the elastic bands approach proposed in [15]. However, this methodology was compliant with a line guidance requirement entailing that both vehicle wheels should follow the same physical path and therefore, the inherent rhombic flexibility was only partially explored. Fig.2 illustrates part of the scenario in TB of ITER where a CPRHS, acting as a rescue vehicle, has to dock in a Vacuum Vessel Port Cell (VVPC) where another CPRHS is parked. For the particular case where both wheels are constrained to follow the same path (center), no solution is found. However, the use of independent references for the wheels would simplify the motion problem (right). The achievement of this solution requires the use of dedicated motion planning techniques, in particular, the employment of a efficient path optimization method capable of handling the high maneuvering ability of the rhombic vehicle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003861_peami.2019.8915294-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003861_peami.2019.8915294-Figure1-1.png", + "caption": "Fig. 1. Sketch of the combined excitation generator.", + "texts": [], + "surrounding_texts": [ + "Located in the Krasnodar region. Length \u2013 89 km, basin area \u2013 885 km2. The longest river in Russia from flowing into the Black sea. Originates on the southern slope of the Main Caucasian ridge at an altitude of 2980 m, in the upper reaches of the mountain lakes flows from the Small Kardyvach and Kardyvach, below the river \u2013 Emerald waterfalls. In the middle reaches and breaks through the ridge Aibga, the Achishkho, forming a Greek gorge below runs through the gorge Ahtsu, Akhtyrskoe gorge. The power of the river is mixed; spring and summer high water and rain floods are characteristic. The average water consumption is about 50 m3/s (the highest \u2013 764 m3/s)." + ] + }, + { + "image_filename": "designv11_33_0002183_aim.2017.8014173-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002183_aim.2017.8014173-Figure1-1.png", + "caption": "Fig. 1: PPNNPPNN octorotor", + "texts": [ + " (7) Equations (1) to (7) represent the rigid body dynamics of any MAV and the model is parametrized by the system\u2019s mass mo, the inertia tensor J and the propulsion subsystem that produces the generalized forces T and \u03c4 . In the sequel, we briefly describe the propulsion system. In our previous work [16] we have considered a PNPNPNPN octorotor configuration. However, due to its apparent higher degree of fault tolerability (see [7]), we now consider an alternative PPNNPPNN configuration. The octorotor system, as shown in Figure 1, consists of eight arms of length l with a 45 degrees angle between adjacent arms, each arm carrying a DC motor that drives a fixed pitch rotor, and a support plate in the center of the system on which all additional hardware is mounted. Each rotor, as shown in Figure 1, produces a force in the positive direction of the local Z axis. For control modeling purposes, these forces can be approximated statically (see [20]) as Fi = b\u21262 i (i = 1..8), where b [Ns2/rad2] is the rotor thrust constant, and Fi [N ] and \u2126i [rad/sec] are the force and the angular velocity ot the i-th rotor, respectively. In addition, as a consequence of Newton\u2019s third law, the rotors produce the counter torques Mi (i = 1..8). We will also approximate them statically as Mi = d\u21262 i , where d [Nms2/rad2] represents the rotor drag coefficient. Following the octorotor model derivation given in [16], we can express the system\u2019s actuation as u = A\u2126s, where \u2126s represents the squared rotor velocity vector given as \u2126s = [ \u21262 1 \u21262 2 \u21262 3 \u21262 4 \u21262 5 \u21262 6 \u21262 7 \u21262 8 ]T and A is the actuation matrix defined as A = b b b b b b b b bl \u221a 2 2 bl 0 \u2212 \u221a 2 2 bl \u2212bl \u2212 \u221a 2 2 bl 0 \u221a 2 2 bl 0 \u2212 \u221a 2 2 bl \u2212bl \u2212 \u221a 2 2 bl 0 \u221a 2 2 bl bl \u221a 2 2 bl \u2212d \u2212d d d \u2212d \u2212d d d . (8) The axes of the local coordinate system, as shown in Figure 1, are principle axes of inertia, and consequently the inertia tensor has the diagonal form J = diag ([ Ixx Iyy Izz ]) , where the components Ixx, Iyy and Izz can be determined via the Huygens-Steiner theorem as shown in [16]. For the sake of more accurate actuation modeling, we also include both the gyroscopic effect and the motor dynamics (see [16] for details). Based on the octorotor model from the previous section, we will design a PD tracking controller and extend it into a fault-tolerant controller by applying a novel RLS-based propulsion FDI technique" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001855_s11249-017-0818-8-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001855_s11249-017-0818-8-Figure1-1.png", + "caption": "Fig. 1 Experimental and numerical setup. a Experimental setup previously used [1]; lower disk textured with ten periodic regions, each with an identical circular texture. b Experimental periodic texture cell. Simulations involved one periodic textured cell. c Simulated periodic texture with top flat plate moving in the direction of positive angular velocity (note z axis stretched compared to r)", + "texts": [ + " To minimize sliding friction with angled cylindrical textures, an optimal angle of asymmetry b exists. The optimal angle depends on the film thickness but not the sliding velocity within the applicable range of the model. Outside the scope of this paper, the model is being used to optimize generalized surface texture topography (Lee et al. in J Mech Design, to appear). Keywords Surface textures Reynolds equation Pseudospectral method Optimization We have previously shown experimentally (setup shown in Fig. 1) that asymmetry is required to reduce friction in fullfilm lubrication by reducing the apparent shear stress and by producing a separating normal force between the two surfaces [1, 3]. There, we minimized experimental effects (inertia, gap accuracy, non-parallelism, and surface tension) in order to obtain accurate results with gap-controlled bidirectional sliding conditions, and the normal forces were attributed to viscous effects up to gap-based Reynolds number Reh qXRh g \u00bc 1:21. We choose this data set for validating a design-driven model because it involves asymmetric depth profiles (slanted-bottom circular textures) with the most complete data set including bidirectional sliding, shear and normal load measurement, controlled gap conditions, and precisely known texturing profiles" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003541_iemdc.2019.8785161-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003541_iemdc.2019.8785161-Figure1-1.png", + "caption": "Fig. 1 Basic configuration of flux-modulated type magnetic gear.", + "texts": [ + " In a previous paper, a magnetic-geared motor was prototyped, and the feasibility and usefulness were demonstrated. However, the efficiency of the prototype motor is not enough high [8]. This paper discusses the efficiency improvement of the magnetic-geared motor from the view point of increasing torque and reducing losses. The improved geared motor is prototyped and evaluated experimentally. Furthermore, the feasibility for applying to walking support machines is studied. II. INVESTIGATION OF EFFICIENCY IMPROVEMENT Fig. 1 shows a basic configuration of a flux-modulated type magnetic gear. It consists of coaxial inner and outer rotors with surface-mounted permanent magnet, and ferromagnetic stationary parts placed between both rotors, which are called pole-pieces (PP). The PP have an important role since the flux-modulated type magnetic gear can be worked as a gear when the PP modulates the magnet flux. Fig. 2 illustrates a configuration of the magnetic-geared motor which was prototyped in a previous study, and Table I shows specifications" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000856_2011-01-2117-Figure14-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000856_2011-01-2117-Figure14-1.png", + "caption": "Fig. 14 Bearing insulation from engine block", + "texts": [ + " The TC method is based on the above principle and determines the minimum oil film thickness hmin from the following equation (6) [8]: where CM represents the capacitance between the crankshaft and main bearing, CR represents the clearance (radial), R represents the radius of the main bearing, W represents the width of the main bearing, and (= o r) represents the dielectric constant of oil. In the oil film measurement by the TC method, the rear surface of the bearing was covered with a 50 m thick polyamide-imide film to electrically insulate the bearings from the engine block (Fig. 14). Figure 15 compares the oil film thicknesses measured by the TC method and thin-film method. Under 2000 rpm and half-load conditions, the TC method measured the minimum oil film thickness of 0.9 m at an angle near 370\u00b0CA. This thickness was smaller than that measured by the thin-film method. 2 MR Rmin CC RW 211Ch -(6) TDC of Cyl. No.2 TDC of Cyl. No.3 TDC of Cyl. No.2 TDC of Cyl. No.1 SAE Int. J. Fuels Lubr. | Volume 5 | Issue 1 (January 2012) 431 The difference in both measured values may be caused by the fact that the thin-film method did not estimate the minimum oil film thickness, because the thickness of the position where the electrode exists is measured in the thin-film method" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000758_s11465-012-0317-4-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000758_s11465-012-0317-4-Figure2-1.png", + "caption": "Fig. 2 The 3-UPU parallel mechanism and its constraint system", + "texts": [ + " When there is linear dependence in the constraint wrench system, the moving platform will possess uncontrollable DOF. When there is linear dependence in the actuated wrench system, namely the actuated wrenches cannot fully actuate the DOF motions, the moving platform occurs under actuated phenomenon. At these configurations, the parallel mechanism cannot afford external forces. The 3-UPU mechanism proposed by Tsai [1] is a 3-DOF parallel mechanism with three pure translational motions in theory, as shown in Fig. 2. Considering a universal joint is equivalent to two intersecting revolute joints, we may take the ith limb as a serial chain connecting the moving platform to the fixed base by five number 1-DOF joints. And the instantaneous motion of the moving platform can be expressed as a linear combination of 5 instantaneous twists, $p \u00bc X5 j\u00bc1 _ i,j$\u0302i,j, i \u00bc 1,2,3, (11) where _ i,j is the intensity and $\u0302i,j denotes a unit screw associated with the jth joint of the ith limb. The twist of the moving platform $p \u00bc \u00bd\u03c9T vTp T is a six-dimensional velocity vector, which includes the possible angular velocity \u03c9, of the moving platform and the linear velocity vp, of a point in the moving platform. Since the DOF of each limb is equal to five, there exists a reciprocal screw $ri , which is reciprocal to the basis screws of each limb. From the relationship between twists and wrenches, we could further know that the reciprocal screw is a constraint couple, with its direction being perpendicular to the plane determined by the two orthogonal axes of universal joint Bi, as shown in Fig. 2. The constraint couple applied on the moving platform can be obtained from the geometric characteristics of the parallel mechanism and the absolute coordinate system shown in Fig. 2, $ri \u00bc \u00bd0 0 0 ; cos\u03b2icos\u03b1i cos\u03b2isin\u03b1i sin\u03b2i , i \u00bc 1,2,3, (12) where \u03b2i denotes the angle from the direction of $ri to fixed base plane and \u03b1i represents the angle from projection line of the twist $ri to x-axis. Due to the fixed base and moving platform are all equilateral triangles, the values of angles \u03b1i are as follows: \u03b11 \u00bc \u03c0 2 , \u03b12 \u00bc 7\u03c0 6 , \u03b13 \u00bc 11\u03c0 6 : Now taking the orthogonal product [28] of both sides of Eq. (11) with the obtained constraint couples for each limb produces three equations, which can be written in matrix form as $r$p \u00bc 0, (13) where $r \u00bc \u00bd$r1 $r2 $r3 ", + " One is obtaining three relative stable parameters \u03b2i by placing the universal joint at the proper position and orientation. These three parameters can make the value of norm\u00f0\u03c9\u00de be lower. Another is decreasing the possible clearance to the greatest extent. From the above analysis and the unexpected motion exhibited in practice [17], the SNU 3-UPU parallel mechanism surely possess parasitic rotations because of some unexpected errors. The exactly description [19] is \u201cAll looked perfect until Prof. Frank Park pushed the mobile platform and the mechanism seemed to collapse under its weight\u201d , see the Fig. 2 in Ref. [17]. Bonev and Zlatanov [19] named this special type of singularity as the constraint singularity. And Wolf [22] pointed that the self motion is a pure rotation about any screw axis, which belongs to the flat pencil, defined by the two axes of the linear complexes. In this section, we will analyze how to eliminate the possible parasitic motions by adding redundantly actuated limbs. The reason for parasitic rotations usually is analyzed as there is linear dependence among constraint wrenches provided by each limb" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002920_mmar.2018.8486052-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002920_mmar.2018.8486052-Figure1-1.png", + "caption": "Fig. 1. Principle configuration of a hydrostatic transmission system.", + "texts": [], + "surrounding_texts": [ + "I. INTRODUCTION\nA hydrostatic transmission system typically consists of two main components \u2013 a hydraulic pump and a hydraulic motor, each with variable volumetric displacement \u2013 that are connected by hydraulic hoses in a closed hydraulic circuit. The pump is driven by an electric motor or a combustion engine, which supplies mechanical power to the system. This power is converted to hydraulic power \u2013 volume flow times pressure \u2013 and transmitted to the hydraulic motor, where it is converted back to mechanical power at the output shaft.\nHydrostatic transmissions are used in classical industrial applications like heavy working machines, construction and agriculture vehicles but also in recent applications such as power-split gearboxes and off-road vehicles, cf. [1], as well as wind turbines, cf. [2]. Hydrostatic transmissions offer many advantages in comparison to purely mechanical solutions: they provide a continuously variable transmission ratio, a high power density, a directional reversion without changing gears, and it is able to serve as a wearless braking system, cf. [1]. The most popular form of a hydrostatic pump and motor involves a design with axial pistons. With a changing swash-plate angle of the hydraulic pump and a changing bent angle of the hydraulic motor, the transmission ratio can be varied by means of displacement units. As a result, both the motor torque and angular velocity of motor output shaft can be adjusted. Beside many advantages like the flexible geometrical arrangement and the operation principle that let hydrostatic transmissions become a new trend in industry, energy efficiency and control issues still need to be addressed. From a control point of view, sophisticated control approaches are required because hydrostatic transmissions are characterized by nonlinear differential equations and,\nDang Ngoc Danh and Harald Aschemann are with the Chair of Mechatronics, University of Rostock, Justus-vonLiebig-Weg 6, Rostock D-18059, Germany, {Ngoc.Dang, Harald.Aschemann}@uni-rostock.de.\nmoreover, are affected by load disturbances and unavoidable parameter uncertainty. Model uncertainty is physically related, for example, to variations in the fluid temperature, to kinematic viscosity, to the elasticity of the hydraulic hoses, and the leakage oil flow, cf. [3].\nIn current industrial practice, gain-scheduled PID controllers are still the typical choice to control hydrostatic transmissions, cf. [4]. Nevertheless, if higher performance demands are to be fulfilled, many nonlinear control approaches have proven advantageous and have been successfully validated in the last decade, see [5] for a general overview and a comprehensive list of references. Regarding the control structure, both centralized and decentralized topologies are applicable and enable an accurate tracking control. The decentralized control of the motor bent angle and the motor angular velocity, see [5], outperformed the centralized topology. The achievable tracking performance is higher, and the implementation is simpler in direct comparison to the centralized approach, see [5] for further details.\nIn this study, the control-oriented model of a hydrostatic transmission is presented in Sec. II. A decentralized scheme discussed in Sec. III serves as the basis for the investigation of alternative estimators \u2013 a state and disturbance observer, an adaptive parameter estimator and a neural network \u2013 employed for the feedback linearization of the nonlinear system. The estimator-based approaches considered in Sec. IV are meant to cancel uncertain nonlinear terms as well as to compensate for disturbances. A comparison is performed in Sec. V by means of both simulation and corresponding experiments. Successful experiments at the test rig validate the concept and show that all variants are applicable and, moreover, allow for an accurate trajectory tracking. Conclusions and a short outlook finish the paper in Sec. VI.\n692978-1-5386-4325-9/18/$31.00 \u00a92018 IEEE", + "The overall system comprises two physical domains: a hydraulic subsystem and a mechanical subsystem, cf. [5]. The hydraulic subsystem characterizes the pressure dynamics of the hydraulic oil, whereas the mechanical subsystem describes the dynamics of motor output shaft.\nAssuming small values of the pump swash-plate angle (|\u03b1P| \u2264 18o ) and of the motor bent axis angle (\u03b1M \u2264 20o ), the corresponding volume flow rates of pump and motor can be approximately expressed by\nqP = V\u0303P\u03b1\u0303P\u03c9P, qM = V\u0303M\u03b1\u0303M\u03c9M.\n(1)\nHere, \u03c9P/M are the angular velocities of pump and motor. The variable \u03b1\u0303P, which is limited to the interval \u03b1\u0303P \u2208 [\u22121,1], represents the normalized swash-plate angle. The variable \u03b1\u0303M \u2208 [\u03b5M,1], with \u03b5M > 0, denotes the normalized bent axis angle. V\u0303P/M stand for the volumetric displacements of the pump and the motor, respectively. For practical reasons, the pressure dynamics is reduced to the dynamics of the difference pressure between the high pressure and the low pressure side. Assuming symmetric conditions and neglecting the pressure losses in the hydraulic hoses, the differential equation for the difference pressure becomes\n\u0394p\u0307 = 2\nCH\n( V\u0303P\u03b1\u0303P\u03c9P\u2212V\u0303M\u03b1\u0303M\u03c9M\u2212 qu\n2\n) . (2)\nHere, CH denotes the hydraulic capacitance, where qu is a disturbance resulting from unknown leakage flows in the physical system.\nThe dynamics of the motor angular velocity is governed\nby the following equation of motion\nJV \u03c9\u0307M +dV \u03c9M = V\u0303M\u0394p\u03b1\u0303M\u2212 \u03c4u . (3)\nHere, JV is mass moment of inertia, and dV denotes the velocity proportional damping coefficient at the rotational output shaft. The variable \u03c4u stands for an unknown load disturbance torque.\nThe dynamics of the displacement units, which represent the actuators of the hydrostatic transmission and which are responsible for changing the swash-plate angle and the bent axis angle, cover the relationships between normalized tilt angles and the input signals. They can be described by firstorder lag systems as follows\nTuP \u02d9\u0303\u03b1P + \u03b1\u0303P = kPuP, TuM \u02d9\u0303\u03b1M + \u03b1\u0303M = kMuM.\n(4)\nHere, TuP/uM are the time constants. kP,M are proportional gains and uP/M are normalized input signals of the pump and motor actuators. Due to given limits of the mechanical design, the actuator dynamics is subject to saturation. In this study, for the sake of simplicity, the saturation is not considered explicitly in system model but is addressed properly by the planning of feasible desired trajectories.\nCombining all subsystems, the overall system can be described by a fourth-order nonlinear state-space representation\n\u23a1 \u23a2\u23a2\u23a3\n\u02d9\u0303\u03b1M \u02d9\u0303\u03b1P \u0394 p\u0307 \u03c9\u0307M\n\u23a4 \u23a5\u23a5\u23a6= \u23a1 \u23a2\u23a2\u23a2\u23a3\n\u2212 1 TuM \u03b1\u0303M + kM TuM uM\n\u2212 1 TuP \u03b1\u0303P + kP TuP\nuP 2\nCH V\u0303P\u03b1\u0303P\u03c9P\u2212 2 CH V\u0303M\u03b1\u0303M\u03c9M\u2212 qu CH\n\u2212 dV JV \u03c9M + V\u0303M JV \u0394p\u03b1\u0303M\u2212 \u03c4u JV\n\u23a4 \u23a5\u23a5\u23a5\u23a6 . (5)\nIn the decentralized control scheme, the overall system dynamics is partitioned into two subsystems: the first subsystem corresponds to the differential equation for the motor bent axis angle, whereas the second one comprises the remaining last differential equations and serves for the control of the motor angular velocity.\nA flatness-based controller, see [6] for an overview of this design method, is developed for the normalized motor bent axis angle, cf. [7]. Solving the first differential equation (5) for the input signal uM and introducing the first time derivative as stabilizing control input \u03c5M = \u02d9\u0303\u03b1M results in the inverse dynamics\nuM = \u03b1\u0303M +\u03c5MTuM\nkM . (6)" + ] + }, + { + "image_filename": "designv11_33_0000274_icara.2000.4803910-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000274_icara.2000.4803910-Figure7-1.png", + "caption": "Fig. 7. Search Space with Time Dimension", + "texts": [ + " We describe the way we extend the steering sets to include planning in dynamic environments and a human tracking system. 1) Path Planning with Moving Obstacles: The search space in a changing environment, includes moving obstacles can be expressed as a three-dimensional space with a time axis in addition to the spatial xy axes. In this search space, static obstacles are expressed as rectangular solids vertically elongated along the time axis, and moving obstacles are expressed as columns obliquely elongated against the time axis. The planning for that three-dimensional space becomes a geometric spatial search (Fig. 7). We assume that the planner gets the position and speed data of the moving obstacles using external devices and uses these data to predict the future positions of moving obstacles. Using the predicted position of moving obstacles, this three-dimension search space will be recognized as the two-dimensional search space which is sliced by time period [18] (Fig. 8). We assume that the search space including all static obstacles is known. The planner receives the obstacles prediction path of the time t+n at a time t at each planning cycle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002035_s12206-017-0515-4-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002035_s12206-017-0515-4-Figure5-1.png", + "caption": "Fig. 5. 3D-FEM models of the LSR: (a) SPM_CMG; (b) FC_CMG.", + "texts": [ + " Thus, the main excitation frequency of the force that acts on each magnet is calculated by the following equation: ( )_ 1,2,3,...ex p r lowf Q fm m= = , (4) where Qp is the number of MPs, and f r_ low is the rotational frequency of the LSR. The number of MPs in the analytical models is 26, as shown in Table 1. Therefore, the electromagnetic force frequencies are multiples of 26fr_low, as shown in Fig. 4(b). Structural analysis is required to investigate the influence on vibration according to the electromagnetic force. The commercial software Ansys 17.2 is used to perform the structural analysis. Fig. 5 shows the 3D-FEM models of the comparison model for structural analysis. The material properties for vibration analysis are shown in Table 3. First, modal analysis is performed using the 3D-FEM, in which the mode shapes and their resonant frequencies can be obtained using the following equation [7, 8]: ( ){ } { }2[ ] [ ] 0i i K Mw- F = , (5) where [K], [M], \u03c9i and { }i F are the stiffness matrix, the mass matrix, the i-th natural frequency, and the eigenvector of the i-th resonance frequency, respectively", + " The vibrations due to the electromagnetic force are calculated by the mode superposition method as follows [9]: { } { } { } { } 1 1 1 [ ] [ ] [ ] N N N i i i i i i i i i M y C y K y F = = = F + F + F =\u00e5 \u00e5 \u00e5&& & , (6) where [C], N, { },F yi and {f} are the damping matrix, the number of modes considered, the i-th mode shape, the displacement in modal coordinate, and the applied radial force, respectively. In this paper, 200 modes are provided to calculate the displacement. Then, the harmonic forces in Fig. 4(b) are applied to the center of each pole of the LSR for the mode superposition analysis. Fig. 7 shows the radial deformation spectra of the analytical models on points A and B, which are in the same location on the LSR surface shown in Fig. 5 when the speed of the LSR is 600 rpm. The rotation frequency is 10 Hz because the rotational speed of the LSR is 600 rpm. Therefore, the electromagnetic force frequencies are multiples of 260 Hz using Eq. (4). Accordingly, radial deformation occurs in the component of a multiple of 260 Hz, as shown in A and C of Fig. 7. Furthermore, resonance occurs in B and D of Fig. 7, with a slightly small excitation force of 12480 Hz, which is close to the resonance frequency of the 5th mode in Fig. 6(e). Consequently, the deformations of the SPM_CMG and the FC_CMG are generated at 12480 Hz" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001677_iros.2011.6095023-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001677_iros.2011.6095023-Figure2-1.png", + "caption": "Fig. 2. An overview of information processing in the system. The input signal I(t) multiplied with input weights win is applied to several springs as external forces u(t). The output of the system y(t) is computed as a sum of all length of springs multiplied with output weights wout. While win is randomly set and then fixed, wout is adapted in order to emulate a desired signal.", + "texts": [ + " To allow diversity of the system, the parameters of each spring (k1, d1, k3, and d3) are randomly drawn from defined huge ranges: for each spring, k1 and d1 are drawn from [1, 0\u00d7104, 1.0\u00d7106] (log-uniform distribution) and k3 and d3 are drawn from [5.0\u00d71012, 1.0\u00d71013] (uniform distribution). We limit the movement of each spring in a direction where the spring lifts up or down the lower link to avoid twist of the lower link. The simulation of physical dynamics is carried out using ODE (Open Dynamic Engine) (R. Smith, http://www.ode.org/ode.html). The information processing based on the musculoskeletal model is shown in Fig. 2. A single input signal I(t) at time step t is transformed into external forces u applied to Nin springs (Nin \u2264 N), which are chosen randomly, by multiplying input weights win, which are randomly drawn from [\u22121.0, 1.0]. The lower link moves, depending on the actuation of springs caused by their own dynamics and the external forces. The state of the model is measured as the lengths of the springs. A linear and static readout unit computes an output of the model y(t) based on the lengths of all springs with output weights wout = (wout,1, wout,2, \u00b7 \u00b7 \u00b7 , wout,N )T : y(t) = \u2211N j=1 wout,j lj(t), where, wout,j and lj(t) indicate the output weight for j-th spring and the length of the spring at time t, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001163_andescon.2010.5633569-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001163_andescon.2010.5633569-Figure4-1.png", + "caption": "Fig. 4. Array of two accelerometers instrumenting the 4WDSS rover.", + "texts": [], + "surrounding_texts": [ + "The robot angular acceleration is measured by deploying a geometric arrangement of accelerometers to determine the total acceleration of a geometric point of the robot chassis (see fig.5). The angular acceleration causes a tangential acceleration which is a component perpendicular to the position vector of such point, where the acceleration is being calculated with the rotation center of the body w.r.t. the origin. The orientation of the tangential acceleration is bearing toward the direction of the angular acceleration and perpendicular to the position vector, ~r, from the rotation center G of the robot, defined by, ~aAG = ~rA \u00d7 ~\u03b1+ \u03c92~rA; ~aBG = ~rB \u00d7 ~\u03b1+ \u03c92~rB (20) Where ~aAG is the relative acceleration of point A (the front side) with respect to G (geometric center), ~aBG is the relative acceleration of point B (the rear side) with respect to G, ~rA is the position vector of point A with respect to G. ~rB is the position vector of point B with respect to G, ~\u03b1 is the angular acceleration vector of the robot, which is normal to the reference plane. The total acceleration of a point is equal to the acceleration of the rotation center G, plus the relative acceleration of the point with respect to G. Thus, for points A and B the following models are stated, ~aA = ~aG + ~aAG (21) ~aB = ~aG + ~aBG (22) Where ~aG is the acceleration of the rotation center of the object. If total accelerations of A and B are subtracted, the acceleration of G is taken out, so this acceleration difference is independent of the chassis acceleration. ~aB \u2212 ~aA = ~aBG \u2212 ~aAG (23) If (21) and (22) are substituted in (23) the result is as follows, ~aB \u2212 ~aA = ~rB \u00d7 ~\u03b1+ \u03c92~rB \u2212 (~rA \u00d7 ~\u03b1+ \u03c92~rA) (24) ~aB \u2212 ~aA = (~rB \u2212 ~rA)\u00d7 ~\u03b1+ \u03c92(~rB \u2212 ~rA) (25) After simplification it leads to: ~aAB = ~rAB \u00d7 ~\u03b1+ \u03c92~rAB (26) Being ~aAB the difference of accelerations of A and B. By definition, ~rAB\u00d7~\u03b1 and \u03c92~rAB are orthogonal, so the following expression is valid. \u2016~aAB\u2016 = \u221a (\u2016~rAB\u2016\u2016~\u03b1\u2016)2 + (\u2016~rAB\u2016\u03c92)2 (27) If we solve for \u2016~\u03b1\u2016, the following equation turns out, \u03b1 = \u2016~\u03b1\u2016 = \u221a( \u2016~aAB\u2016 \u2016~rAB\u2016 )2 \u2212 \u03c94 (28) With this resulting equation is possible to calculate the angular acceleration of the moving object if the acceleration of two different geometric points are known. This formula includes the angular speed, which depends on the angular acceleration itself. As for implementation, such angular acceleration needs to be calculated iteratively using sensor readings. The next formulation is a recursive algorithm to calculate the angular acceleration, the angular speed, and the orientation angle of the robot, \u03b1t = \u221a\u221a\u221a\u221a(\u221a(axB \u2212 axA)2 + (ayB \u2212 ayA)2 l )2 \u2212 \u03c94 t\u22121 (29) \u03c9t = \u03c9t\u22121 + \u222b t \u03b1tdt (30) \u03b8t = \u03b8t\u22121 + \u222b t \u03c9tdt (31) Where \u03b1t is the current angular acceleration of the robot. axA, axB , ayA and ayB are the acceleration values obtained from both two-axis accelerometers. l is the distance between the position of the accelerometers on the robot. \u03c9t is the current angular speed of the robot. \u2206t is the time interval between each sensor reading, and \u03b8t is the current angle of the robot." + ] + }, + { + "image_filename": "designv11_33_0003222_ffe.12997-Figure13-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003222_ffe.12997-Figure13-1.png", + "caption": "FIGURE 13 The geometry of the pinion and wheel with points T1 and T2 for which the evolution of stresses were determined5,33", + "texts": [ + " In numerical calculations, material assigned to both gears was steel with following material parameters: modulus of elasticity E = 206 800 MPa; Poisson ratio \u03bd = 0.3. Value of the total normal force used amounted to F bt = 12 600 N, and the Hertzian load distribution was used. Friction was neglected in this case. Evolutions of stress components during a meshing cycle were determined for point T1 situated 2.5 mm under the tooth flank's surface and point T2 located 1.0 mm under the surface as shown in Figure 13. Values of stresses components determined using developed model (\u03c3x, \u03c3y, and \u03c4xy) and numerically (\u03c3x,FE, \u03c3y, FE, and \u03c4xy,FE) are presented in diagrams in Figures 14 and 15. Values of stress components \u03c3x, \u03c3y, and \u03c4xy were first calculated in local coordinate system corresponding to Figure 4 and were then recalculated to the global coordinate system of the numerical model to enable direct comparison with values determined numerically, \u03c3x,FE, \u03c3y,FE, and \u03c4xy,FE. Results of numerical (FE) and proposed model, and for both points T1 and T2, coincide very closely for the most part of the loading cycle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000675_icelmach.2010.5608123-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000675_icelmach.2010.5608123-Figure1-1.png", + "caption": "Fig. 1. Main subparts of a rolling-element bearing.", + "texts": [ + " This approach is often redundant and does not guarantee the avoidance of abnormal functioning. Moreover, it is inefficient from an economical pointof-view because still sound components could be replaced. Furthermore, disregarding the cost of the replaced component, the downtime incurred for an unnecessary maintenance event should be included in the performance balance. Consequently, the ability to determine the state-of-health of a rolling-element bearing assumes considerable practical and economical importance. As illustrated in Fig. 1 a rolling-element bearing is composed of three main components: the inner race, the outer race, and the rolling-elements. Components can either be damaged independently, or as part of a multi-component failure mode. Consequently, several fault diagnosis problems can be defined for a rolling-element bearing. Of course, the main objective is to determine if a rolling-element bearing is healthy or damaged. In the case of bearing diagnostics through analysis of vibration, this problem is straightforward to solve, due to the relevant difference in the spectral frequency content between a healthy and damaged bearing", + " DE and FE accelerometers data were digitally recorded at 12 KHz sampling frequency. Data were acquired in three different conditions: motor operating with two normal bearings, motor operating with faults seeded in the DE bearing, and motor operating with faults seeded in the FE bearing. Single point faults with three different diameters i.e., 7 mils, 14 mils and 21 mils (1 mils = 10\u22123 in) were seeded by electrodischarge machining on deep grove ball bearings, at three different locations i.e. inner race, outer race, and rollingelement (see Fig. 1), into a 6205-2RS JEM SKF bearing and into a 6203-2RS JEM SKF bearing mounted on the DE and the FE, respectively. The data downloaded from the web repository consists of 95 time series from tests with at least one faulty bearing, and four time series from tests with two healthy bearings, for a total of 99 time series. Each time series is contained in a Matlab file named nnn.mat where nnn is a three digit integer number. As previously stated, each time series corresponds to a working condition with at least one faulty bearing described by three main characteristics (position, location, and size)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001113_cca.2010.5611189-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001113_cca.2010.5611189-Figure3-1.png", + "caption": "Fig. 3. The three control objectives: (i) length, (ii) height and (iii) angle.", + "texts": [ + " Also the acceleration needs to be limited for each actuator (max acc) to ensure that the desired motion is actual possible for both cranes. This summarizes the pre-control and is enough to move both cranes theoretically synchronously. Since no crane is perfect in its dimensions, weight and actuation, positional errors will occur. Therefore, three objectives of the virtual traverse need to be feedback controlled to overcome these errors. (i) The horizontal distance, (ii) the vertical distance, and (iii) the angle with respect to a vertical axis between the hooks need to be controlled, see Fig. 3. The virtual traverse is initiated to these objectives at the beginning of the tandem-lift. Continuing with Fig. 2, we transform the speeds to the end-effector velocities via Jacobian transformations, where the length error w.r.t. the traverse is added via a compensator. Finally, we transform them back to the actuator speeds and limit each single one to its maximum speed and acceleration. The feedback control of the angle is added via compensator at the very beginning, when the rotational speeds are added to the end-effector/hook speeds" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002380_0954410017716477-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002380_0954410017716477-Figure1-1.png", + "caption": "Figure 1. Geometric of the leader\u2013follower formation.", + "texts": [ + " In the Formation control analysis section, analysis of formation keeping in various conditions is presented. The structure of the proposed guidance law is also treated step by step in this section. Simulation results for the formation-keeping of multi unmanned vehicles via two examples is provided in the Simulation results section. The Conclusions section summarizes our work. The leader\u2013follower formation problem is formulated in this section. Let us consider the geometric of the leader\u2013follower formation as shown in Figure 1. In this figure, three coordinate frames I, L and V are defined as the inertial reference frame, the line of sight (LOS) frame and the follower velocity frame, respectively. The leader\u2013follower relative distance is denoted by rL and, VL is the relative angle. Moreover, VF and L denote the follower velocity vector and the LOS angle, respectively. Now, consider the following definition for the leader\u2013follower formation. Definition 1. A formation is said to be achieved if the relative distance rL and the relative angle VL converge to constant values" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003920_s00500-019-04522-1-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003920_s00500-019-04522-1-Figure3-1.png", + "caption": "Fig. 3 High yo\u2013yo maneuver trajectory characteristic analysis", + "texts": [ + " Take the high yo\u2013yo maneuver as a case to demonstrate the process of extracting the characteristic parameters. A high yo\u2013yo maneuver is performed when a UCAV is approaching an enemy aircraft at a high speed and has lost shot opportunity, converting the speed advantage to a height advantage and thus avoiding flying over the enemy or being in a passive situation. A schematic sketch of the high yo\u2013yo maneuver is shown in Fig. 2. The high yo\u2013yo maneuver consists of three segments: climb, observation and dive. A detailed analysis is shown in Fig. 3. The climb starting point S1 and dive ending point S5 are the starting point and ending point of the whole maneuver, respectively; the observation starting point S2 and dive starting point S4 are the turning points; and S3 is the extreme point, which is also the highest point in the maneuver. The five points S1, S2, S3, S4 and S5 can completely and accurately describe the characteristics of a high yo\u2013yo maneuver; thus, these five points are extracted as the characteristic parameters of the high yo\u2013yo maneuver" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002165_j.jmapro.2017.08.008-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002165_j.jmapro.2017.08.008-Figure7-1.png", + "caption": "Fig. 7. (a) Support-model interface full contact and (b) Support-model interface partial contact.", + "texts": [ + " A partial contact based support structure removal technique is pplied in this paper to assist the removal of support material. The usion between model and support material is the key force to overome during the removal of support material. The surface energy etween them is usually reduced through differentiating the mateial composition especially in extrusion based system. Here, the elationship between sagging and raster-width is used along the odel support interface for the ease of separation as shown in Fig. 7. The raster width in AM is defined as the distance between onsecutive filaments. In layer by layer process, materials are suported by the deposited layer underneath and most often are upported with 00\u2013900 pattern. Increasing the raster width will ecrease the material uses which is a common technique for conomical 3D printing. However, depending upon the material haracteristics and process parameter it can also introduce saging of the filament due to the gravitational load. If, MRWt and the support needed points accordingly, (d) Model-support interface surface, (e) SRWc is the maximum raster distance required to print the model and support material without sagging respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000723_s11661-010-0561-3-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000723_s11661-010-0561-3-Figure1-1.png", + "caption": "Fig. 1\u2014Finite-element model of the substrate and deposition region for one-layer cladding; the first activated set of elements is shown in a higher magnification.", + "texts": [ + " The temperature history of the nodes at the surface are then transferred to a subroutine developed in the programming language MATLAB (Natick, MA) to analyze the phase transformation and final hardness. The finite-element (FE) model is described briefly here, but more details can be found in Reference 14. The FE analysis of the temperature of the LPD process with respect to time and location requires the geometry of the part that is defined by a mesh of finite elements updated over time to represent the additive nature of the process. Figure 1 shows the geometry of the finite-element model used in this study, and in a higher magnification, the first set of activated elements on the first bead. The deposition area is 15 9 5 mm, and the spacing between deposition tracks is 0.5 mm. The continuous movement of the laser beam over the substrate is divided into small divisions of the stationary heat transfer analysis, called time steps. In every time step, the energy balance Eq. [1] is solved using the finiteelement method, and the results are used as the initial conditions for the next time step", + " The thermal behavior of the material is defined by specifying the temperature-dependent specific heat, thermal conductivity, and density of the material. The latent heat of fusion as suggested by Reference 22 is considered 1908\u2014VOLUME 42A, JULY 2011 METALLURGICAL AND MATERIALS TRANSACTIONS A in the definition of specific heat. Table I shows the temperature-dependent material properties of AISI H13. In the current study, the laser scanning speed is 10 mm/s and the powder feeding rate is 3 g/min with a powder efficiency of 30 pct. The deposition area, as shown in Figure 1, is 15 9 5 mm, and the spacing between deposition tracks is 0.5 mm. With a time step of 25 milliseconds, a set of elements with volume of 0.25 9 0.5 9 0.4 mm3 is activated (Figure 1). The area on which the heat flux is activated is 0.25 9 0.50 mm2. The thermal analysis of the process results in determining the temperature with respect to time and location. These results are used to predict the phase transformations during the process. The microstructure of the deposited material, just after the solidus temperature, consists of austenite. Depending on the cooling rate of the process, different phases can be developed from the austenite.[25] In the LPD process, the high cooling rate from a large low temperature mass of substrate results in a martensite transformation only when the temperature drops below Ms", + "314 1910\u2014VOLUME 42A, JULY 2011 METALLURGICAL AND MATERIALS TRANSACTIONS A Based on the developed thermokinetic model, the effect of path planning on surface modification is studied. The material AISI H13 is used to cover the surface of a substrate of AISI 1018 with three different deposition patterns as shown in Figure 4. The hardness of the surface of the deposited material is calculated for each deposition pattern. To minimize the computational costs of calculations in MATLAB, the nodal solution of the nodes located at the sides of each bead (like along the dashed line shown for the first bead in Figure 1) is taken into account. The results show that the deposition pattern can change the hardness distribution of the deposited material significantly, as a result of changing the heat input. The surface hardness of the one-section pattern, as shown in Figure 5, has a minimum value at the starting bead and a maximum value at the last two beads. In between the starting and ending edges, the hardness has a fairly uniform distribution with a lower value than the maximum hardness. The surface hardness of the deposition on the twosection pattern is shown in Figure 6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002109_1350650117723484-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002109_1350650117723484-Figure1-1.png", + "caption": "Figure 1. Two kinds of preloads: (a) constant preload using spring; (b) fixed-position preload using nut lock.", + "texts": [ + " Bearing characteristics, constant preload, bearing stiffness, thermal effects, lubricant film thickness Date received: 5 April 2017; accepted: 26 June 2017 An angular contact ball bearing is widely used in a high-speed spindle because of its low friction, simple structure, easy lubrication, and low cost. When bearings are mounted in the spindle, a preload is applied to create a high stiffness. There are two ways to apply the preload to the ball bearings: fixed-position preload and constant preload. Figure 1(a) and (b) illustrates a spindle-applied constant preload using the spring system and fixed-position preload using nut locks, respectively. The constant preload method is often employed for a high-speed spindle to avoid exceeding the preload due to a rise in temperature. Numerous researchers have established models for analyzing angular contact ball bearing characteristics. Palmgren1 presented two degrees-of-freedom (2- DOF) for a spindle bearing based on the displacement of the bearing rings. Based on the Hertz contact theory,2 Jones3 modeled the bearing with 5-DOF under arbitrary load and speed conditions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000988_s00542-010-1189-3-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000988_s00542-010-1189-3-Figure4-1.png", + "caption": "Fig. 4 Finite element model and pressure distribution of the coupled journal and thrust bearings", + "texts": [ + " WX, WY, WZ and Lf are the external force of each axis and the z distance of the external force from the mass center G, respectively. The nonlinear equations of motion of the HDD spindle system are solved by using the fourth-order Runge\u2013Kutta method to investigate the motion of the HDD spindle system, including the whirling, tilting, and axial motions. Table 1 shows the major design specifications of the FDBs in this research that support the HDD spindle system with a 3.5-inch disk. They consist of two grooved journal bearings, four plain journal bearings, two grooved thrust bearings, and one plain thrust bearing. Figure 4 shows the finite element model and the pressure distribution of the coupled journal and thrust bearings. The fluid film was discretized by 10,740, four-node, isoparametric bilinear elements, and the Reynolds boundary condition was applied in order to guarantee the continuity of the pressure and pressure gradient. The accuracy of the developed program was verified by comparing the calculated flying height of the coupled journal and thrust bearings in equilibrium (where the axial load generated by the FDBs was equal to the weight of a rotor) with the measured flying height at various rotation speeds (Jang et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000856_2011-01-2117-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000856_2011-01-2117-Figure8-1.png", + "caption": "Fig. 8 Bearing tester", + "texts": [ + " We determined hi by substituting these values into equation (3) and using the capacitance CM that was measured between the lead wires. Thus, we could determine the oil film thickness that was independent of the effects of the protective layer, stray capacitance, and other factors. PROOIL PROOIL STRINSM CC CCCCC -(2) OILSTRINSM CCCC )Clearanceh( i -(4) nF02.0CCC )Clearanceh( iMSTRINS -(5) PROSTRINSM r0i C 1 CCC 1S h -(3) SAE Int. J. Fuels Lubr. | Volume 5 | Issue 1 (January 2012) 427 To verify the accuracy of the thin-film method, we conducted an oil film thickness measurement test on a bearing tester (Fig. 8), holding the test bearing rigidly to simplify its behavior. We then compared the measurement results with oil film thickness calculation results based on the elasto-hydrodynamic lubrication (EHL) theory that assumed the Reynolds equation: Where, R is the bearing radius, h0 and p are the oil-film thickness and pressure respectively at time t, \u03bc is the lubricant dynamic viscosity, is the angular bearing abscissa, z is the axial coordinate and is the analytical crankshaft rotating speed [14] [15]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000882_srin.201000032-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000882_srin.201000032-Figure6-1.png", + "caption": "Figure 6. Representativecast part (top: steel die insert, bottom:TZM die insert).", + "texts": [ + " A visual inspection revealed no spallation or coating defects. Phase analysis using Xray diffraction (XRD) revealed significant amounts of a-Al2O3 as well as g-Al2O3 as constituents. When thixocasting steel, it is particularly necessary to strip the scaled shell off the heated billet so that no oxides and www.steelresearch-journal.com 2010 W insulation fibers are carried into the cast part during the mould filling. For this purpose, the billet is pressed through an aperture, consequently only the interior of the billet enters the cast part (Figure 4 and Figure 6). Since the aperture is subjected to maximum stress caused by the heat energy and the shrinkage of the solidifying steel, it is recommended that the apertures are designed to be disposable. A practical solution is the use of apertures made from steel sheet (3 mm unalloyed steel was used here). Hence, no complex core slider technology is necessary. Moreover, the disposableaperture controls the heat balance. On the one hand, it withdraws energy from the semi-solid steel, by which the die iley-VCH Verlag GmbH & Co", + " It is assumed that a higher die temperature is advantageous for the tool life since the thermal shock load is reduced. Commercial Al2O3-coated ejectors were used for both the steel and the TZM die inserts. Notwithstanding a thermal expansion value twice that of steel (aTZM 7 10-6 K 1, asteel 13 10-6 K 1 between 20 and 700 8C), these ejectors led to no problems. In an initial series of tests, 10 cast parts were manufactured prior to plunger failure. The reasons for the failure are explained below. Figure 6 shows a representative cast part. All components exhibit very good cast surfaces. Contrary to the strongly oxidized aperture, the combination spanners show almost no tarnishing. This results from the high cooling rates and despite this, the cavity was filled completely. The combination spanners produced from the steel side tarnish slightly more than those from the TZM side. This was expected from the simulation. Owing to the high cooling rates, the cast steel partially forms martensite which leads to a high hardness value. For www.steelresearch-journal.com 2010 W this reason, a hardening processing step, which is necessary after forging, can be omitted thus reducing costs. In order to determine the mechanical properties, both notched-bar impact as well as tensile test specimens were taken from the partly tempered runners. Metallographic samples were taken from a small number of combination spanners (Figure 6). For the evaluation of the dies, one die insert each of TZM and of steel was sectioned and examined by means of SEM. The etched samples exhibit a very fine, globular microstructure. The shell is clearly recognizable, which is formed by the segregated melt in the wake of the Magnus-effect and is typical for thixocastings at regions of high flow velocity (Figure 7a). Advantages of the shell, which is mainly formed by the liquid phase, are the outstanding surface quality and reproduction of accuracy as well as its good burnishing properties", + " In this condition, it is therefore impossible to employ the material for hand tools. Owing to the very high strengths, which are comparable to those of hardened X39CrMo17, it must be assumed that the tempering was not successful. As previously mentioned, part-ejection after thixocasting proved to be effortless, indicating negligible adhesion of the 586 2010 Wiley-VCH Verlag GmbH & Co. KGaA, Wein as-cast parts to the mould surface. Excellent surface quality of the cast combination spanners was observed (Figure 6) whilst the die tool coatings remained macroscopically undamaged following 10 shots. As shown in Figure 9a, Stage 2, the cross-sectional microscopic images at different positions of the tool surface indicate that the coating is intact. Tool examination after use revealed a thinning of the coating at sharp edges within the cavity that were exposed to both high thermal stresses as well as high flow stresses during mould filling (Figure 9b). After four castings, TZM already cracked in the region of the inlet" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001256_s1068366611010053-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001256_s1068366611010053-Figure1-1.png", + "caption": "Fig. 1. Shaft bush contact diagram.", + "texts": [ + " The present paper considers a flat contact between an elastic shaft and an elastic bush with the account of wear of both bodies. No limits are imposed on the angle of contact between the shaft and the bush, so the obtained solution is suitable to calculate heavily loaded sliding bearings. The equation of local wear in the point is assumed as an exponential function of pressure and velocity. Consider the contact of elastic cylinder S0 with radius R0 having a round hole with radius R1 in the infinite elastic agent S1 (Fig. 1). Assume that the radii R0 and R1 differ little. The shaft is loaded with the lin ear force Q that is distributed over the upper semicircle of the shaft and acts in the direction opposite to the axis y. Assume also that the interface cross section is exposed to planar strain. The elastic cylinder S0 rotates in a cylindrical notch in the elastic medium S1, with both bodies undergoing wear. The wear rates of the shaft and the bush depend on the contact pressure p and the linear sliding velocity v is assigned by the relations (1) where Ki, \u03b1i, \u03b2i, i = 0, 1 are the set parameters" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002101_med.2017.7984240-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002101_med.2017.7984240-Figure2-1.png", + "caption": "Fig. 2. 2DOF Testbed reference frames and angles.", + "texts": [ + " Dealing with the firmware code can be conducted both in the original Arduino software environment, and with more powerful and convenient software such as Microsoft Visual Studio, e.g. It should be noted that the microcomputer might not have enough of processing power to execute the complex algorithms. For this case it is possible to communicate quadrotor with programs written in Python by means of Xbee modems. The feature of this approach lies in the fact that the large processing power may be needed for control signal computation. To describe the Testbed dynamics let us use the following nomenclature (see Fig. 2): OXgYgZg denotes the inertial normal orthogonal reference frame associated with the basement. OXY Z stands for the movable body-fixed orthogonal reference frame associated with the quadrotor. \u03d1 and \u03b3 denote the Euler angles (pitch and roll, respectively). Denote the projections of the quadrotor angular velocity vector \u03c9 to the corresponding axis of body-fixed reference frame OXY Z as \u03c9x, \u03c9z (rotation velocity \u03c9y around OY axis is zero). It should be noted that due to the 2DOF Testbed construction, its center of gravity lies in the close vicinity to the origin point O of the reference frames OXgYgZg and OXY Z" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001290_j.mseb.2010.12.001-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001290_j.mseb.2010.12.001-Figure2-1.png", + "caption": "Fig. 2. Two-layered wires used for producing of new types of fuse links by laser alloying or melting.", + "texts": [ + " Achieving proper functional properties of fuse links requires creation of narrowings on a conducting element. A steep decrease of cross section in narrowings is necessary for shortening of switch-off-time and pre-arcing integral [2]. The search is still on for the best way of F , alloy w o s m a d c s p m a i a m n n c 2 2 t r c r s T e A m C f c t a l N T i s f o of very small diameter, consisting of the core with high electrical resistivity and the surface layer with high electrical conductivity were considered (Fig. 2). Laser micro-melting-alloying has been suggested for the purpose of producing HRMA. To this end the Cu- ig. 1. (a) \u2013 A view of HRMA made by cw and pulse laser on Cu foil (thick. 0.18 mm ith current of Cu wire, diameter 0.18 mm, with HRMA. btaining good switching properties of fuses. Laser cutting with mall heat affected zone, laser alloying of a very small area and theral or ablative removal of conducting layers by the laser radiation re the microtechnologies which could be used for production of escribed fuse links", + " The Gaussian distribution of power density of radiation of laser beam was assumed: q = q0 \u00b7 exp ( \u2212x2 + y2 r2 ) (5) density 2.2 \u00d7 108 W/m2:(a) \u2013 after 2.16 ms; (b) \u2013 after 3.3 ms from the begining of where q0 is maximum of power density, r \u2013 radius of laser spot equal 150 m. The module responsible for convection heat transfer has been omitted in calculations. The error caused by this simplification was not greater than 10% [12,13] (which is acceptable for good engineering estimation of laser beam parameters). Simultaneously the time of calculation was reduced. Due to low thickness (Fig. 2) and big difference between melting temperatures of the metals forming two-layered system Ag/kanthal (Fig. 2b, TmAg = 1234 K, Tmkanthal = 1773 K) the kanthal core did not reach melting temperature under influence of the laser impulse, which otherwise was sufficient for removal of the outer layer. Therefore a thermocapillary convection could not considerably change a shape and a depth of the laser pool [14\u201316] in this case. Laser treatment on two-layered system Cu/FeNi caused melting of core and outer layer. The thermocapillary convection omission in the analysis is justifiable by the sufficient agreement of results from experiments and simulations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002870_joe.2018.8345-Figure11-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002870_joe.2018.8345-Figure11-1.png", + "caption": "Fig. 11 Vibration acceleration of the 4 pole 36 slot motor at 900 Hz (a) Motor A, (b) Motor B", + "texts": [ + " The electromagnetic force calculated by Maxwell was transferred to the harmonic response calculation module of the structure field, so as to perform the coupling analysis of the electromagnetic field and the structure field. The harmonic response analysis module was applied to solve the vibration displacement. By selecting the modal superposition method for calculation, the frequency response, vibration displacement, velocity, acceleration of the entire vibration of the stator, and vibration displacement, velocity, and acceleration at a certain node can be obtained. Fig. 11 shows the overall vibration acceleration of the stator at the slot frequency of 900 Hz for model motor A and model motor B. Fig. 12 shows the overall vibration acceleration of the stator at the slot frequency of 750 Hz for model motor C and model motor D. According to the simulation results, the motor C was a typical second-order vibration mode with a maximum amplitude of 2.9 \u00d7 10\u22128 m, located between the two magnetic poles. Motor D was also a typical second-order mode with a maximum amplitude of 9" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002058_aer.2017.57-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002058_aer.2017.57-Figure1-1.png", + "caption": "Figure 1. (Colour online) Twin rotor MIMO system.", + "texts": [ + " 2.0 TRMS DYNAMIC MODEL TRMS is a laboratory setup multi-input\u2013multi-output MIMO system developed by Feedback Instruments Ltd. for identification, analysis and real-time control experiments(35). The TRMS simulates the main and tail rotor subsystems of a rotorcraft system with their interactions resulting in highly nonlinear cross-coupled dynamics. It is widely used as scaled model of highly nonlinear vehicles with strong cross-coupling effects. The phenomenological modelling of the TRMS shown in Fig. 1 is based on the use of the first principal approach, which leads to the following highly nonlinear cross-coupled MIMO dynamic model:(36) ht tp s: // do i.o rg /1 0. 10 17 /a er .2 01 7. 57 D ow nl oa de d fr om h tt ps :/w w w .c am br id ge .o rg /c or e. C ol um bi a U ni ve rs ity L ib ra ri es , o n 04 Ju l 2 01 7 at 0 3: 44 :4 4, s ub je ct to th e Ca m br id ge C or e te rm s of u se , a va ila bl e at h tt ps :/w w w .c am br id ge .o rg /c or e/ te rm s. \u23a7\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a8\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a9 \u03b1\u0307h \u0307h \u03c9\u0307h \u03b1\u0307v \u0307v \u03c9\u0307v \u23ab\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23ac\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23ad = \u23a7\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a8\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a9 h Lt f2(\u03c9h )\u2212K h h\u2212 f3(\u03b1h ) Jh \u2212 ( (Kah\u03d5h )2 JtrRah + Btr Jtr ) \u03c9h \u2212 f1(\u03c9h ) Jtr + Kah\u03d5h JtrRah hh (uh) v( Lm f5(\u03c9v )\u2212K v v+g[(A\u2212B) cos(\u03b1v )\u2212Csin(\u03b1v )] Jv ) \u2212 ( (Kav\u03d5v )2 JmrRav + Bmr Jmr ) \u03c9v \u2212 f4(\u03c9v ) Jmr + Kav\u03d5v JmrRav hv (uv ) \u23ab\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23ac\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23ad \u2026 (1) where \u03b1h, \u03b1v , h, v , \u03c9h, \u03c9v , uh, uv , refer to the horizontal and vertical angular positions of the beam, yaw and pitch rates, rotational speeds of the main and tail rotors, and input voltage signals of main and tail rotors, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002655_jifs-169558-Figure13-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002655_jifs-169558-Figure13-1.png", + "caption": "Fig. 13. Drivetrain dynamics simulator (DDS): 1. a two stage planetary gearbox; 2. an accelerometer; 3. a two stage parallel shaft gearbox with rolling or sleeve bearings; 4. a programmable magnetic brake.", + "texts": [ + " The CIs of energy, RMS, STD, AM and kurtosis show better performance on the indication of PG health state. Using the energy as an example, the change for a small crack, such as the crack of 2.5 mm or 5 mm, is difficult to be noticed, and the change for the latter three cracks presents a visible rising. Based on above results, the use of CIs presents a useful way to monitor and evaluate the vibration of healthy as well as faulty planetary gearboxes. Experiments were conducted on a drivetrain dynamics simulator (DDS) shown in Fig. 13. The drivetrain consists of a two stage planetary gearbox, a two stage parallel shaft gearbox with rolling or sleeve bearings, a bearing loader, and a programmable magnetic brake. An accelerometer was vertically mounted on the planetary gearbox housing to collect vibration signals. The tested PG has three planet gears in the first stage and four planet gears in the second stage. It is noted that, limited by manufacturing error and small size of the sun gear, the artificially made crack is relatively thick, which is not as slim as a real crack" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002525_28465-ms-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002525_28465-ms-Figure5-1.png", + "caption": "Figure 5\u2014Schematic showing the key stages in the build process from design to finished product.", + "texts": [ + " Build parameters and design optimisation For the build parameters themselves, there are many variables and factors. Additive manufacturing machines are able to store parameter sets, and it is common practice to correlate these to the resultant properties of the component during the trial stages of a job. Although in theory this should ensure consistency, there are still effects such as powder quality that can affect the final build quality. Regular maintenance, verification and calibration of the machine set up are necessary. The key steps in the build of an additively manufactured item are illustrated in Figure 5. To take the designers CAD drawings and convert them to a printed item is currently not a trivial task. The design needs to be translated to a file format written in a language that the additive manufacturing machine understands such as .STL files or .AMF files. This process carries a risk of creating voids in the build as during the conversion, some of the elements may not map precisely. Software is available that assists with finding potential defects, and independent validation of this type of software would be beneficial to give the users confidence in the capabilities" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003893_0954406219888240-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003893_0954406219888240-Figure2-1.png", + "caption": "Figure 2. 3D model of the combined spherical joint.", + "texts": [], + "surrounding_texts": [ + "The three-dimensional (3D) model of the ideal spherical joint and the combined spherical joint is shown in Figures 1 and 2, respectively. Usually, the workspace shape of the ideal spherical joint is a cone that can be simply determined by a constant angle, but it is not accurate for the combined spherical joint. If the spherical joint is assembled by a universal joint and a rotating unit, the workspace of the combined spherical joint is related to two Euler angles (yaw angle and pitch angle) of the universal joint and a roll angle of the rotating unit. The roll angle of the combined spherical joint is free, but the yaw angle and pitch angle are constrained by the non-interference domain of universal joint. Thus, the workspace of the universal joint is important for the combined spherical joint, which will be discussed first. Although the workspace analysis of this class of universal joints has been discussed,11 they are too complex to use. Inspired by the cross-sectional area method, a concise projection method is presented to study the joint workspace in this paper, so that only two parameters can determine the non-interference domain of this class universal joint." + ] + }, + { + "image_filename": "designv11_33_0002697_codit.2018.8394945-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002697_codit.2018.8394945-Figure1-1.png", + "caption": "Fig. 1. Quadrotor scheme", + "texts": [ + " Stability analysis of the closed loop system is illustrated in section 4. In section 5 simulation results are presented. Conclusions and remarks are drawn in section 6. In this section, we will first introduce the quadrotor dynamic model; then, we will present a stochastic model of the wind gust disturbances. To study the quadrotor motion dynamics, two frames are used: an inertial frame attached to the earth defined by Ea(ea1, ea2, ea3) and a body-fixed frame Eb(eb1, eb2, eb3) fixed to the centre of mass of the quadrotor as depicted in Fig. 1. The absolute position of the quadrotor is described by p = [x, y, z]T and its attitude by the Euler angles \u03b7 = [\u03c6, \u03b8, \u03c8]T . The attitude angles are respectively Yaw angle (\u03c8 978-1-5386-5065-3/18/$31.00 \u00a92018 IEEE -1023- rotation around z-axis), Pitch angle ((\u03b8 rotation around y-axis), and roll angle (\u03c6 rotation around x-axis) [11]. The dynamic model for quadrotor vehicle can be derived as mx\u0308 = (cos\u03c6 sin \u03b8 cos\u03c8 + sin\u03c6 sin\u03c8)u1 my\u0308 = (cos\u03c6 sin \u03b8 sin\u03c8 \u2212 sin\u03c6 cos\u03c8)u1 mz\u0308 = u1 \u2212mg Ix\u03c6\u0308 = (Iy \u2212 Iz)\u03b8\u0307\u03c8\u0307 + J\u03b8\u0307$ + lu2 Iy \u03b8\u0308 = (Iz \u2212 Ix)\u03c6\u0307\u03c8\u0307 \u2212 J\u03c6\u0307$ + lu3 Iz\u03c8\u0308 = (Ix \u2212 Iy)\u03c6\u0307\u03b8\u0307 + u4 (1) where $ = \u03c91\u2212\u03c92+\u03c93\u2212\u03c94" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002741_012011-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002741_012011-Figure5-1.png", + "caption": "Figure 5. DC Motor Control using Relay", + "texts": [], + "surrounding_texts": [ + "A series of manufacturing and testing processes have been undertaken to ensure the condition of the kit is safe to use. The main frame is made using acrylic 3 mm instead of metal. In addition to facilitate the 7 1234567890\u2018\u2019\u201c\u201d process of making and lightweight, acrylic is also useful to avoid the occurrence of short circuit on the main panel. The overall weight of the kit is 6.76 kg, Figure 6. Thus still within the maximum weight standard. The Figure 7 shows internal component and Figure 8 is the kit along with the completeness of the power cable, USB cable and connection cable." + ] + }, + { + "image_filename": "designv11_33_0000547_asjc.454-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000547_asjc.454-Figure1-1.png", + "caption": "Fig. 1. Regular triangular array of optical mice (N =3).", + "texts": [ + " Section II obtains the mobile robot velocity estimation as the simple average of the optical mouse velocity readings. Section III shows the robustness of the proposed velocity estimation against measurement noises and partial malfunction. Section IV suggests simple but effective method of position calibration of imprecisely installed optical mice. Section IV gives some experimental results to validate the proposed velocity estimation method. Finally, the conclusions are provided in Section VI. II. MOBILE ROBOT VELOCITY ESTIMATION As shown in Fig. 1, assume that N optical mice are installed at the vertices, Pi , i =1, . . .,N , of a regular polygon that is centered at the center, Ob , of an omnidirectionalmobile robot traveling on the xy plane. Figure 1 shows an example of a regular triangular array of optical mice with N =3. The position of the ith optical mouse, i =1, . . .,N , from the center Ob to the vertex Pi , can be expressed as pi = [ pix piy ] = \u23a1 \u23a2\u23a2\u23a3 r cos { +(i\u22121)\u00d7 2 N } r sin { +(i\u22121)\u00d7 2 N } \u23a4 \u23a5\u23a5\u23a6 , i =1, . . .,N (1) where represents the heading angle of a mobile robot moving forward in the direction of p1, and r represents the distal distance of each optical mouse. It should be noted that N\u2211 i=1 pix = N\u2211 i=1 piy =0 (2) regardless of the heading angle ", + " Kim: Systematic Robustness Ana ysis of Least Squares Mobile Robot Velocity Estimation \u00a9 2011 John Wiley and Sons Asia Pte Ltd and Chinese Automatic Contr l Society S. Kim and H. Kim: Systematic Robustness Analysis of Least Squares Mobile Robot Velocity Estimation random measurement errors are almost completely identified and isolated from the new velocity estimation. As discussed previously, simple averaging or median filtering may work well for sporadic measurement errors, but cannot be as effective for prolonged measurement errors as our consensusbased remedy. For a regular triangular array of optical mice (N=3), shown in Fig. 1, with the heading angle =\u2212 6 [rad], we set the nominal (wrong) optical mouse position, p\u0304, and the actual (true) optical mouse position, p\u0302, as follows: p\u0304= 0.3\u00d7[0.866 \u22120500 0.000 \u22120.866 \u22120.500]t (40) p\u0302= 1.05p\u0304=0.3\u00d7[0.909 \u22120.525 0.000 1.050 \u22120.909 \u22120.525]t . (41) It should be noticed that we set the actual optical mouse array to be slightly bigger than the nominal one on purpose, as will be shown later. The mobile robot is commanded to make a selfrotation at a constant speed b= 1 3 [rad/s] with vbx = vby =0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001434_046009-Figure10-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001434_046009-Figure10-1.png", + "caption": "Figure 10. (a) Tertiary structure of the flagellin protein constructed with atomic positions of the R-type flagellin from [35] and our model protein superimposed. The elongated shape of the D0 (red) and D1 (blue) domains of the flagellin molecule coincides with the axial distance of the inner and outer binding sites in the model protein. (b) Arrangement of model proteins in filaments without twist (left) and with twist (right). Note the large deflection of the inner binding sites from the central axis when twist is present. Dark gray: three model proteins along a 1-start helix; light gray: two model proteins from a neighboring 1-start helix.", + "texts": [ + " Several authors have stressed that release of strain between the inner (D0) and outer (D1) domains of flagellin determines the helical ground state of the filament [2, 19, 35, 39, 40]. The most important result of our work is that the axial distance between the inner and outer binding sites reveals the mechanism via which the associated stress is produced and released. We will discuss this mechanism below. It only depends on the overall shape of the flagellin protein and does not rely on molecular details. In figure 10(a), we present a representation of the real flagellin protein produced from data of its atomic positions [35]. We also superimpose our model protein with an axial displacement of 9 nm between inner and outer binding sites which was necessary to stabilize the normal form of the filament. There is a remarkable agreement between the elongated shape of flagellin and our model protein. We note that with their original binding scheme Namba and Vondervistz could successfully calculate curvature and torsion of the different polymorphic configurations within a coarsegrained model that does not rely on molecular details", + " Extending this successful coarse-grained model, we could identify that the elongated shape of flagellin favors a helical ground state. Now we ask the question how this elongated shape favors the helical ground state? We know from the results in figure 6 that a filament built from our model protein with a more compact shape, which means axial distance D = 0 nm, has the largest elastic energy when curvature is the largest. Thus, straight filaments are favored. On the other hand, straight filaments with internal twist built from elongated model proteins contain internal stress. This is illustrated in figure 10(b). On the left-hand side the model proteins are arranged without twist and the inner binding sites stack onto one another along the central axis with little cost of elastic energy. If the filament possesses an internal twist, the inner binding sites are strongly deflected from one another and from the central axis when the axial displacement D is large (picture on the right). This produces large stresses within the model proteins or a large elastic energy for high twists. In order to relax it, the filament assumes a helical ground state as a compromise with reduced torsion and some curvature" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001394_tac.2017.2684459-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001394_tac.2017.2684459-Figure1-1.png", + "caption": "Fig. 1. Illustration of locally-polyhedral and locally-smooth functions", + "texts": [ + " From the non-expansive property, we have that \u2016\u03bb(t+ 1)\u2212 \u03bb\u2217\u20162 \u2264 \u2225 \u2225\u03bb(t) + 1 V h(t)\u2212 \u03bb\u2217 \u2225 \u2225 2 = \u2016\u03bb(t)\u2212 \u03bb\u2217\u20162 + 1 V 2 \u2016h(t)\u20162 + 2 V [\u03bb(t)\u2212 \u03bb\u2217]\u22a4h(t) \u2264 \u2016\u03bb(t)\u2212 \u03bb\u2217\u20162 + 2C V 2 + 2 V [d(\u03bb(t))\u2212 d(\u03bb\u2217)], (23) where the last inequality uses the definition of C and the concavity of the dual function (6), i.e, d(\u03bb1) \u2264 d(\u03bb2) + \u2202d(\u03bb2) \u22a4[\u03bb1 \u2212 \u03bb2] for any \u03bb1, \u03bb2 \u2208 \u03a0, and \u2202d(\u03bb(t)) = h(t). Throughout Section IV-A, the dual function (6) is assumed to have a locally-polyhedral property, introduced in [15], as stated in Assumption 2. A dual function with this property is illustrated in Figure 1. The property holds when f and gj for every j are either linear or piece-wise linear. Assumption 2: There exists an Lp > 0 such that the dual function (6) satisfies d(\u03bb\u2217) \u2265 d(\u03bb) + Lp\u2016\u03bb\u2212 \u03bb\u2217\u2016 for all \u03bb \u2208 \u03a0 (24) 0018-9286 (c) 2016 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. where \u03bb\u2217 is the unique Lagrange multiplier. The \u201cp\u201d subscript in Lp represents \u201cpolyhedral.\u201d Furthermore, concavity of dual function (6) ensures that if this property holds locally about \u03bb\u2217, it also holds globally for all \u03bb \u2208 \u03a0 (see Figure 1). The behavior of the generated dual variables with dual function satisfying the locally-polyhedral assumption can be described as follows. Define Bp(V ),max { Lp 2V , 2C V Lp } . Lemma 3: Under Assumptions 1 and 2, whenever \u2016\u03bb(t)\u2212 \u03bb\u2217\u2016 \u2265 Bp(V ), it follows that \u2016\u03bb(t+ 1)\u2212 \u03bb\u2217\u2016 \u2212 \u2016\u03bb(t)\u2212 \u03bb\u2217\u2016 \u2264 \u2212 Lp 2V . (25) Proof: Suppose the following condition holds 2 V [d(\u03bb(t))\u2212 d(\u03bb\u2217)] + 2C V 2 \u2264 \u2212Lp V \u2016\u03bb(t)\u2212 \u03bb\u2217\u2016+ L2 p 4V 2 , (26) then the inequality (22) in Lemma 2 becomes \u2016\u03bb(t+ 1)\u2212 \u03bb\u2217\u20162 \u2264 \u2016\u03bb(t)\u2212 \u03bb\u2217\u20162 \u2212 Lp V \u2016\u03bb(t)\u2212 \u03bb\u2217\u2016+ L2 p 4V 2 = [ \u2016\u03bb(t)\u2212 \u03bb\u2217\u2016 \u2212 Lp 2V ]2 ", + " The deviation from the optimality value (29) is bounded above by O(1/V +1/T ). The constraint violation (30) is bounded above by O(1/T ). To have both bounds be within O(\u01eb), we set V = 1/\u01eb and T = 1/\u01eb, and the convergence time of Algorithm 1 is O(1/\u01eb). Note that both bounds consider the average starting after reaching the steady state at time Tp, and this transient time Tp is at most O(1/\u01eb). Throughout Section IV-B, the dual function (6) is assumed to have a locally-smooth property, introduced in [15], as stated in Assumption 3 and illustrated in Figure 1. Assumption 3: Let \u03bb\u2217 be the unique Largrange multiplier, there exist S > 0 and Ls > 0 such that whenever \u03bb \u2208 \u03a0 and \u2016\u03bb\u2212 \u03bb\u2217\u2016 \u2264 S, dual function (6) satisfies d(\u03bb\u2217) \u2265 d(\u03bb) + Ls\u2016\u03bb\u2212 \u03bb\u2217\u20162. (34) Also, there exists Ds > 0 such that whenever \u03bb \u2208 \u03a0 and d(\u03bb\u2217)\u2212 d(\u03bb) \u2264 Ds, dual variable satisfies \u2016\u03bb\u2212 \u03bb\u2217\u2016 \u2264 S. The \u201cs\u201d subscript in Ls represents \u201csmooth.\u201d Using a similar proof process as in Section IV-A, the convergence result under the locally-smooth property is as follows. Define the smooth counterparts of Bp(V ) and Tp(V ): Bs(V ),max { 1 V 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003266_4243_2019_10-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003266_4243_2019_10-Figure2-1.png", + "caption": "Fig. 2 Different sensor formats for contactless oxygen measurements, applied to cultivation systems in bioprocesses: sensor spots (a), sensor nanoparticles (b), sensor nanoparticles in microfluidic channel (c). These figures were kindly provided by PyroScience GmbH (Germany)", + "texts": [ + " In this way, by embedding the sensitive dyes in an O2-permeable polymer matrix, the sensor remains stationary during the bioprocess causing the least possible disruption and leaving valuable space for other required devices (e.g., stirrer, reagents, or gas feedthroughs). This sensor format has made miniaturization of the sensor possible, to their use not only in bioreactors and shake flasks but also in small-scale plastic vessels (typically 50\u2013250 mL), and from microwell plates to microfluidic devices (Fig. 2). Many examples are described in the literature for different applications, and, as of today, these oxygen sensor formats are commercially available from various manufacturers. For instance, Tolosa et al. [26] have described a DO luminescent thin sensing layer, based on a platinum complex as indicator dye, immobilized in a polymer and T ab le 1 S om e co m m er ci al lu m in es ce nt O 2 se ns or s S en so r O pe ra tin g ra ng e A cc ur ac y (% ) t 9 0 (s ) \u2205 (m m ) A pp lic at io n R ef . M et tle r T ol ed o In P ro 68 70 i 8 pp b\u2013 60 % [ 1% + 8 pp b] < 20 12 C el l cu ltu re s ht tp s: //w w w " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001264_vppc.2012.6422500-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001264_vppc.2012.6422500-Figure7-1.png", + "caption": "Fig. 7: Examples of natural frequencies and modal shapes at 2816Hz (left) and 5204Hz (right)", + "texts": [], + "surrounding_texts": [ + "[1] Gieras, 1. F., Wang, c., and Lai, 1. C, Noise of Polyphase Electric Motors, CRC Press., 2006 [2] Swann, S. A. and Nasar, S. A., \"Eflect of rotor eccentricity on magnetic field in air gap of a non-salient-pole machine,\" Proc. lEE, vol. 110, No. 5,pp. 903-915,1964 [3] Kim, D., Kim, J., and Hong, J, \"A study on the reduction of noise and vibration of SPM according to reduction of Permanent Magnet,\" Int, Conf on Electrical machines and Systems, pp. 1252-1255,2010 [4] Sun, T., Kim, 1., Lee, G., and Hong l., \"Effect of pole and slot combination on noise and vibration in pennanent magnet synchronous motor,\" IEEE Magnetics, vol. 47, pp. 1038-1041,2011 [5] lung, 1. W., et aI., \"Reduction Design of Vibration and Noise in IPMSM Type Integrated Starter and Generator for HEY,\" IEEE Trans. Magnetics, vol. 46, no. 6, pp. 2454-2457, 2010 [6] He, G., Huang, Z., and Chen, D., ''Two-Dimensional Field Analysis on Electromagnetic Vibration-and-Noise Sources in Permanent-Magnet Direct Current Commutator Motors,\" IEEE Trans. Magnetics, vol. 47, no. 4, 2011 [7] LMS Yirtual.Lab Acoustics User Manual [8] Kim, S. and Park, S., \"Electric motors with an asymmetric rotor,\" patent applied for, 2011" + ] + }, + { + "image_filename": "designv11_33_0002503_1350650117753915-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002503_1350650117753915-Figure2-1.png", + "caption": "Figure 2. Fluid domain. (a) Entire fluid domain, (b) Oil supply pipe and (c) Oil films.", + "texts": [ + " Finally, using equations (11), (12), and (13), the volume fraction of air can be calculated as rair \u00bc 1 1\u00fe 1= Tloc T0 e 2:476\u00fe 250:8 Tsur psur ploc e 2:476\u00fe250:8 Tloc \u00f014\u00de where psur \u00bc 101, 325 Pa, Tsur \u00bc 40 C. The validation of this cavitation model has been checked in papers.34\u201336 This cavitation model can be inserted into the CFX software by CFX Command Language (CCL) expressions. The calculation domain of FRB consists of two oil films and an oil supply pipe. The calculation domain is shown in Figure 2(a). The parameters are shown in Table 1. The bearing parameters used in the simulation are exactly the same as with those in the experiment by Hatakenaka and Yanai.1 Given the complicated shape of the fluid domain, the entire domain should be divided to keep the mesh in good quality. In this work, the domain is divided into two parts, namely, oil films and oil supply pipe, as shown in Figure 2(b) and (c). These parts are discretized into a structured mesh as shown in Figure 3. This work employs the Grid Convergence Index (GCI) method for grid independence validation as recommended by American Society of Mechanical Engineers (ASME). This method provides a dependable method to estimate the uncertainty error of the grid. The ring-to-shaft speed ratio is regarded as an important parameter in FRB calculation. The torques on the inner and outer surfaces of the ring are the variables that determine the floating ring rotation speed; thus, they are selected as parameters to be verified" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002297_ldia.2017.8097244-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002297_ldia.2017.8097244-Figure5-1.png", + "caption": "Fig. 5. Initial mesh without the air layer.", + "texts": [ + " L is the pitch width of the mover, z is mover position, and I is the effective value of the current. In this paper, the vernier motor in Fig. 2 is called an 8- 6 vernier motor, and that in Fig. 3 is called a 14-12 vernier motor. The 14-12 vernier motor has twice as many magnets as the 8-6 vernier motor. By increasing Z1 and Z2, the 14-12 vernier motor is expected to generate a larger thrust. The static thrust characteristics of each model were investigated by using 3-D FEM. The amplitude of the current is 0.6 A, and a three-phase alternating current shown in (6) was applied. Fig. 5 shows the initial mesh of the 8-6 vernier motor and Table II shows the analytical parameter. The analyzed results are shown in Figs. 6, 7, and 8. Table III shows the average thrust, ripple rate, permanent magnet volume, and average thrust per permanent magnet volume of each model. The 14-12 vernier motor can generate almost the same thrust as the conventional model. We verified that the ripple rate is also lower than the conventional model. In this way, the vernier models can be expected to exhibit higher performance than the conventional model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000777_s0025654411040017-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000777_s0025654411040017-Figure1-1.png", + "caption": "Fig. 1.", + "texts": [ + "OI: 10.3103/S0025654411040017 Keywords: wheeled vehicle, equilibrium stability, rolling friction, dissipation. The vehicle (car) model under study (Fig. 1) consists of three rigid bodies, namely, two wheel pairs 1 and 2 of masses m1 and m2 and radii r1 and r2 whose centers of mass are at points O1 and O2, respectively, and car body 3 O1O2C of mass m and central radius of inertia \u03c1 whose center of mass is at point C. The system moves in the positive sense of the axis Ox. We study it in the vertical plane in the fixed coordinate system Oxz, where the axis Ox coincides with the support plane (the road) and the axis Oz is vertical and opposite to the gravity force", + ") The car body is assumed to be absolutely rigid, and the wheels are assumed to be deformable (with dissipation) in the vertical direction; this is conventionally modeled by a suspension of the wheel centers of mass O1 and O2 over the road on springs and dampers with stiffness and damping coefficients c1, \u03b31 and c2, \u03b32, respectively. The assumption that the wheels are deformable (with dissipation) naturally results in the appearance of friction torque resistance to their rolling; the friction torques Mk1 and Mk2 are equal to f1N1 and f2N2, respectively, and directed counterclockwise, i.e., opposite to the rolling direction (Fig. 1). Here N1 and N2 are the positive normal reactions at the points of contact between the wheels and the road and f1 and f2 are the wheel rolling friction coefficients depending on numerous factors, the main of which is the velocity of motion of the wheel centers. Following [1\u20133], we assume that fk = dk + ekv + hkv 2, k = 1, 2, (1.1) where v is the vehicle motion velocity and dk, ek, hk are phenomenological coefficients determined experimentally (see the State Standard for Pneumatic Tires [3]). Further, we assume that the wheels roll along the road without slipping. This assumption implies that, at the points of contact between the wheels and the road, there are tangential forces of static (adhesion) friction F1 and F2 directed along the axis Ox and satisfying the corresponding Coulomb inequalities (Fig. 1). *e-mail: zurav@ipmnet.ru **e-mail: gr51@mail.ru 495 Let v = const be the steady-state velocity of the vehicle motion under the assumptions described in Sec. 1. As a result of action of the moment M1 caused by the engine as an external energy source for overcoming the above-described resistance forces and because of the existence of some (small) initial deviations from the equilibrium and some initial velocities of system points, there arise vertical vibrations of the wheel centers and, as a consequence, of the car body ", + " Assume that (x1, z1) and (x2, z2) are the coordinates of the centers of mass of wheels 1 and 2, (xC , zC) are the coordinates of the car body center of mass C, a1 = xC \u2212 x1 and a2 = x2 \u2212 xC are the distances from the car body center of mass C to the axes of wheels 1 and 2 measured along the horizontal in equilibrium, b1 = z1 \u2212 zC and b2 = z2 \u2212 zC are the same quantities measured along the vertical in equilibrium, and \u03d5 is the angle of the car body rotation counted off from the horizontal, i.e., the angle between the line O1O2 (Fig. 1) and the negative direction of the axis Ox counted clockwise. In the triangle O1O2C (Fig. 1), we introduce the following notation: \u03c80 is the angle CO2O1; \u03c70 is the angle CO1O2; CO2 = l2; CO1 = l1; and O1O2 = l. We also assume that the wheel masses are negligibly small (m1 = m2 = 0) and the car body mass is equal to unity. Then the system under study has three degrees of freedom determined by the generalized coordinates xC , zC , and \u03d5. For this system, we can write out three equations of dynamics as follows: x\u0308C = F1 + F2, z\u0308C = \u2212g + N1 + N2, \u03c12\u03d5\u0308 = \u2212(F1 + F2)zC + [l1 cos(\u03c70 \u2212 \u03d5) \u2212 f1]N1 \u2212 [l2 cos(\u03c80 + \u03d5) + f2]N2, (3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000206_s0022112065001040-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000206_s0022112065001040-Figure4-1.png", + "caption": "FIGURE 4. Piston motion and shock waves.", + "texts": [ + " It is generally not difficult to obtain values near 100 for P,x/ra: in gun tunnel operation, and the curves therefore indicate that, except at very high velocities, the piston velocity may be expected to come within a few percent of its asymptotic value within the length of the tube. As outlined by Winter, the compression of the test gas to the first, and highest, peak of pressure is accomplished essentially by a series of reflexions between the piston face and the end of the tube, of the shock wave which initially precedes the piston. These shock reflexions are evident in the photograph in figure 4, which is a time-resolved schlieren record of the motion of a nylon piston, of mass 30 mg, as it rebounds from the closed end of a 0.25 in. diameter glass tube. The background of black and white lines is caused by inhomogeneities in the glass of the tube, against which the piston trajectory appears as a dark, curved line, with the shock waves as inclined light lines. The maximum piston velocity was 12OOft./sec. The process is represented alongside on an x-t diagram, where both the shock and the expansion waves emanating from the piston as it slows down are drawn", + " This ratio is seen to be consistently less than unity and, taking into account the further increase in the speed of sound as the peak pressure is approached, it is reasonable to assume that the expansion waves are fairly weak, and may be neglected for a first approximation. Finally, the pressure at the rear face of the piston rises t o high values only as the piston velocity becomes low, i.e. towards the very end of the compression stroke. It is therefore taken to be constant at its value before shock reflexion from the piston. Using these assumptions, an energy balance may be formulated, beginning when the piston is at 1, in figure 4, and equating the energy lost by the piston, 664 R. J . Stalker plus the work done on the piston by the driver gas, with the energy gained by the compressed gas. Defining a cycle as being completed by the passage of the shock from the piston to the closed end and back again, then, at the end of any where the subscript 1 refers to conditions when the piston is at 1, the subscript n + 1 refers to conditions at the end of the nth cycle, P,, is the peak pressure developed a t the closed end during this cycle, x is now the distance of the piston from the closed end, and yo the ratio of specific heats in the test gas", + " and a peak pressure of about 15,000 p.s.i. An increase of approximately 50 yo over the peak pressure for perfect gas was noted, but this disappeared in accounting for the increased entropy change across the shock wave when it first reflects from the piston. These conditions represenl; modest pressure levels, and real-gas effects may be expected to reduce as pressures rise, leading to the conclusion that errors induced in maintaining the assumption of perfect gases are small enough to be ignored here. It can be seen from figure 4 that, after achieving peak pressure in the first violent compression, the piston rebounds from the closed end and executes one or more oscillations before coming to rest. Though not a primary limitation on performance, these oscillations do imply a delay in establishing steady conditions in the test gas, and are therefore worthy of a brief, not necessarily very accurate, treatment. Such a treatment has been reported (Stalker 1961), and satisfactorily compared with experimental results. A resume of the analysis only will be presented here. The model used for the piston motion is shown in figure 4. The shock R, generated as the piston first reverses, is assumed to propagate at constant velocity into a region of uniform flow, conditions behind the shock are calculated by assuming a stationary piston, the velocity of the piston is assumed to be considerably lower than the speed of sound behind the shock, and the pressure is taken as constant between the piston and the end at any instant. The equation of motion for the piston is thus where subscriptfrefers to conditions in the test gas when the piston is stationary, and a, is then the speed of sound in the driver gas behind the piston" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001386_ssp.164.67-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001386_ssp.164.67-Figure2-1.png", + "caption": "Fig. 2. Loading of the specimen", + "texts": [ + " Five gear wheels 13 are mounted on the intermediate shaft 14. At the lever 4 there are the tensometers joined with the conditioning card DBK 16 with the tensometric amplifier. The required stress is obtained in the specimen by suitable set-up of the rotational column 5; the angle \u03b1 = 0 corresponds to bending, the angle \u03b1 = \u03c0/2 means torsion, and each intermediate value of the angle \u03b1 causes the biaxial stress state generated by the simultaneous bending and torsional moments. A scheme of the specimen loading is presented in Fig. 2. Values of the torsional moment Mt(t) and the bending moment Mb(t) loading the specimen are expressed by the relationship )( )( tM tM tg b t=\u03b1 , (1) the resultant moment is calculated from the relation )()()( 22 tMtMtM bt += . (2) The machine drive is based on gear wheels connected with toothed belts. The weights located at the disks of the vibrator determine the component amplitude of vibrations of each disk. Amplitudes of the bending and torsional moments Mba and Mta are equal to the product of the lever arm length l (4 in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003396_j.ijheatmasstransfer.2019.05.052-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003396_j.ijheatmasstransfer.2019.05.052-Figure3-1.png", + "caption": "Fig. 3. Verification of the model with (a) a lab-scale TC4 electron beam melting p", + "texts": [ + " Because of the difficulties in precisely controlling industrialscale TC4 electron beam melting experiments, the accuracy of the model was verified with the data reported in the literature. In Ref. [21], a circular-moving electron beam was used to melt the TC4 sample under vacuum, and the liquid pool profile in the etched sample was identified by the change in the grain size. Using the calculation domain and boundary conditions reported in Ref. [21], the present model predicts the pool profile of the TC4 sample with turbulence model off. A comparison of the numerical prediction of the pool depth and the experimental observation in Ref. [21] is shown in Fig. 3(a), which confirms the accuracy of the present model for solidification evolution. The Al concentration variation tendency due to evaporation is absent in Ref. [21]. To further verify the present model for evaporation prediction, a group of supplementary experimental data was selected from Ref. [24]. The data shows the change of the average Al concentration in a stirred TC4 alloy molten pool with continuous Al evaporation. A comparison of the data and the numerical prediction results using the reported boundary conditions is shown in Fig. 3(b). The results suggest that the numerical prediction with the present model for the change of the aluminum concentration due to evaporation loss is satisfactory given some of the uncertainties and assumptions involved and experimental error. Overall, these comparisons between simulation and experiment indicate that the results using present model have relatively high reliability. Thus, the present numerical models of solidification and aluminum evaporation were used for further calculations. To explore the effect of the Al evaporation rate on segregation of aluminum, the aluminum distributions in round ingots produced by EBCHR with a casting speed 10 mm/min and different pouring temperatures were determined" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002459_robio.2017.8324729-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002459_robio.2017.8324729-Figure4-1.png", + "caption": "Fig. 4: Catenary curve applied to the mathematical model for the cable", + "texts": [ + " (1) with the condition of xq < x < xp. However, in practice, the position of the lowest point of the loose cable changes while UAVs are flying because both or one side of the endpoint position change. Besides, it is technically difficult to measure the lowest point during the flight operation. Therefore, this paper derives a mathematical model of Catenary curve based on the coordinates \u03a3Base whose origin is located on the endpoint of one side instead of using the information of the lowest point position as shown in Fig. 4. Here let us express the position of the lowest point R of Catenary curve with respect to the coordinates \u03a3B as (xr, zr). From eq. (1), we can simply express Catenary curve as: C\u2032 : z = a ( cosh x\u2212 xr a \u2212 1 ) + zr (3) The position of the endpoint Q in \u03a3 and the position of the lowest point R in \u03a3Base has a relationship, which can be expressed as (xq, zq) = \u2212(xr, zr). Since the position of the endpoint Q in \u03a3 can be expressed as: zq = a ( cosh xq a \u2212 1 ) (4) eq. (4) is rewritten as follows. C\u2032 : z = a ( cosh x\u2212 xr a \u2212 cosh xr a ) (5) However, eq. (5) still includes the information of the lowest point of the cable, xr. To remove this, now we consider the relative distance between two endpoints (s, v) as shown in Fig. 4. Then, we can express the lowest position xr in eq. (6) with respect to \u03a3B(Fig. 4). xr = \u2212a tanh\u22121 v l + s 2 (6) where l represents entire length of the cable between the endpoints P and Q, v and s indicate relative distances between two endpoints in vertical and horizontal directions, respectively. Finally, the mathematical model of Catenary curve with respect to the coordinates \u03a3B can be obtained as in eq. (7). Eq. (7) indicates that the shape of the cable can be expressed from the UAV located on one side of the cable. CB : z = a { cosh (x a + tanh\u22121 v l \u2212 s 2a ) \u2212 cosh ( tanh\u22121 v l \u2212 s 2a )} (7) As expressed in Section III-A, Catenary number is a key parameter to determine Catenary curve, which generally requires an information of mass density, tension measurement, and angle of the cable at one endpoint during the operation When we have light linear density, we can not measure tension properly because of the sensor noise and it is very difficult to estimate proper shape of the cable" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001735_ac301951w-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001735_ac301951w-Figure1-1.png", + "caption": "Figure 1. (a) Schematic of the experimental setup. (b) Bright-field image of the spot formed on the glass substrate after focusing the laser beam. (c, d) Scanning electron micrographs of the spot: (c) top view and (d) 45\u00b0 view.", + "texts": [ + " Horseradish peroxidase (HRP) (Wako, >100 units/mg) and glucose oxidase (GOD) (Toyobo Enzyme, >100 units/mg) were dissolved in citrate buffer (pH 4.6) at 0.01 and 0.1 mg/ mL, respectively. HRP and GOD stock solutions were mixed in the ratio 1:1 immediately before use. The absorption spectra of sample solutions were measured by using a UV\u2212visible absorption spectrometer (Shimadzu, UV-2550). The topography and morphology of the produced nanostructures were evaluated by using an atomic force microscope (AFM) (SII, SPI-4000) and a scanning electron microscope (SEM) (FEI, DB-235). Figure 1 shows a schematic of the experimental setup. A green (wavelength 532 nm) diode-pumped solid-state (DPSS) laser beam (Shanghai Dream Lasers, SDL-532-020TL) was made incident on an optical inverted microscope (Olympus, IX70) via a beam expander. The laser beam was focused by using an objective lens (60\u00d7, N.A. 0.9). The x\u2212y position of the sample and the laser irradiation time were controlled by a motorized stage and a mechanical shutter, respectively. The Rayleigh scattering spectrum of a single aggregate was measured by using a multichannel spectrometer (B&W Tek i-trometer) under the dark-field illumination of a halogen lamp. The backreflected light intensity of the laser beam was detected by using a photomultiplier (Hamamatsu Photonics, R1166). \u25a0 RESULTS AND DISCUSSION o-PD is a reagent that forms a colored product upon oxidative polymerization.12 First, we investigate the effect of focusing a laser beam on an o-PD aqueous solution. The solution was dropped on a microscope cover glass, and a 2-mW green laser beam was focused on the glass\u2212solution interface, as shown in Figure 1a. Figure 1b shows an optical microscope image of a small spot formed at the laser focus position after laser irradiation of the 1 mM o-PD solution for 80 s. Figure 1c,d shows scanning electron microscope (SEM) images of this spot. Interestingly, it was seen that the aggregates comprised smaller nanoparticles. The surface of the aggregate was covered by the nanoparticles, but the outline had a regular domelike shape. The radial size of the aggregates corresponds to that of the laser focus (\u223c800 nm), indicating that the chemical reaction is significantly accelerated only at the laser-irradiated spot. If this reaction were to be attributed to the thermal effect, aggregates much larger than the laser spot would be formed because of the thermal conductivity of the solvent and the glass substrate" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003921_b978-0-12-803581-8.11722-9-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003921_b978-0-12-803581-8.11722-9-Figure1-1.png", + "caption": "Fig. 1 Viscous clutches; a \u2013 disk clutch, b \u2013 cylindrical clutch: 1 \u2013 driving part, 2 \u2013 driven part, 3 \u2013 working fluid.", + "texts": [ + " Basic Information Concerning Hydraulic Clutches and Brakes Efficiency of a hydraulic clutch Z, which is the ratio of the received power to the delivered power, is described by the following relation: Z\u00bc M o2 Mo1 \u00bc o2 o1 \u00bc ik \u00f01\u00de where: M \u2013 hydraulic torque transmitted through the hydraulic clutch, o1 \u2013 angular velocity of the input shaft of the hydraulic clutch, o2 \u2013 angular velocity of the output shaft of the hydraulic clutch, ik \u2013 speed ratio. For hydraulic brakes Z\u00bc 0 and ik \u00bc 0, as o2 \u00bc 0. Hydraulic clutches and brakes are classified into two basic groups: viscous and hydrodynamic. In viscous clutches and brakes the driving and driven parts are connected by friction caused by shear stress present in the fluid. According to the shape of working surface, one can distinguish two basic types of viscous clutches and brakes: disk and cylindrical ones, Fig. 1. Hydrodynamic clutches and brakes consist of two rotors (a pump connected to input shaft, and a turbine connected to output shaft), a housing in which there are bearing supported shafts, and a working fluid which circulates in the canals of the working space within the rotors and housing. In hydrodynamic clutches and brakes the rotors are connected by the fluid\u2019s hydrokinetic impact on the rotor blades. Depending on the build of the rotors, one can distinguish between hydrodynamic clutches and brakes whose rotors contain inner rings, and those whose rotors have no inner rings" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001319_icelmach.2012.6350265-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001319_icelmach.2012.6350265-Figure8-1.png", + "caption": "Fig. 8. Description of algorithm with magnetic/electric coupling.", + "texts": [ + " Various software (Flux2D, ANSYS Maxwell, Opera-2d...) provide this feature that facilitates the user\u2019s work. However, model of switching remains complex and unusual. We have developed an approach using the FLUX2D software controlled with Python code to reflect the switch as defined at the previous section. The main program uses the results of the previous iteration to determine the state of the circuit at the moment while providing the appropriate values of contact resistances and arc voltage. The principle of the algorithm is described in Fig. 8. B. 11BSolve and mesh description We solve as a transient magnetic time-stepping coupled to an electrical circuit at constant speed. The electrical equation solved is (4). .. IRV c (4) where \u03a6 is the flux vector computed by the finite element software in every coil, V and I are the voltage and current vectors of the circuit branches, cR is the resistance matrix of the circuit. The speed \u2126 is assumed to be constant. The angular step \u0394\u03b8 has to be related to the size of the mesh in the rotating air gap", + " The mesh description of the entire machine is shown in Fig. 9 and in Fig. 10 for the air gap. Mesh in the stator must be particularly refined to take into account eddy currents in the solid iron stator. C. 12BImplementation of the brush/segment contact resistance and the arc voltage in the coupled circuit As mentioned in the previous section, the knowledge of the contact brushes/segment is essential for the study of commutation. The brush-segment contacts are modeled in the circuit by Nb \u00d7 Ns switches as shown on Fig. 8 where Nb and Ns are respectively the number of brushes and segments. We isolate Fig. 11 the model of one segment contact with the Nb brushes. The switch-on / switch-off orders of the switches are represented by the A(\u03b8) matrix given in (5). This matrix of Nb rows and Ns columns contains only 0 or 1 and defined as follows. )(, jiA (5) Note that the matrix A(\u03b8) can contain only one \u201c1\u201d per column. Switch resistances are located in a matrix also depending on the position, written R(\u03b8). The current term of this matrix is calculated according to the mutual surface between the brush i and the segment j ", + " The relation between resistances and surfaces has already been developed in the previous section. Finally, every term ijr of the resistance matrix R(\u03b8) can also be expressed from a single term )(11 r ; the other terms are obtained by circular permutation. The )(11 r term has been calculated and measured and shown on Fig. 7. s ji N i p jrr 2 )1()1()( 1,1, (6) When no mechanical contact exists and if the current in the related coil in not totally reversed, an arc can occur at the brush trailing edge. In such situation an additional arc voltage, which is not represented in Fig. 8, is added to the model in series with the corresponding switch. The switch is then stuck in the \u201cclosed\u201d position and an arc voltage is added to the contact resistance until the arc vanishes, i.e. the current in the coil is equal to the current in the main path considered (the currents comparison is ensured by the algorithm). The arc voltage characteristic has been explained in the previous section. IV. 3BRESULTS AND COMPARISON WITH MEASUREMENTS This model provides, with a good accuracy, the knowledge of internal variables which can be difficult to measure such as the current in a given coil in the armature" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003222_ffe.12997-Figure12-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003222_ffe.12997-Figure12-1.png", + "caption": "FIGURE 12 Meshed finite element model of the gear pair in contact with the position and orientation of the model's global coordinate system5,33", + "texts": [ + " In the following subsections, three separate cases, ie, three different gearings, have been analysed in order to validate three main features of the developed calculation model and procedure\u2014the accuracy of calculation of stress and strain components and their evolution during the meshing cycle, calculation of fatigue life, and accuracy of prediction of location, orientation, and morphology of initiated cracks. As a part of an earlier study,5 accuracy of proposed model for determination of individual stress components and their evolution during the single loading cycle was evaluated by comparing calculation results with results of photoelastic experiments and results obtained through finite element nonlinear contact analysis. Figure 12 shows the numerical model prepared in ANSYS. To make the model more efficient, only part of the rim and five teeth of the pinion and part of the rim and six teeth of the wheel were modelled. For meshing of gears models, three types of finite elements were used. Main part of the geometry was meshed with six\u2010node triangle parabolic elements (PLANE82). As loading was modelled with contact, parts of teeth flanks in contact were meshed with contact elements (TARGE169 and CONTA172)\u2014parabolic elements with 2 end and 1 midside node, each with two degrees of freedom (translation in x and y direction) and very suitable for analysis of plane stress or plane strain problems" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002561_978-3-319-69748-2_2-Figure2.9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002561_978-3-319-69748-2_2-Figure2.9-1.png", + "caption": "Fig. 2.9 Principle of RE (Ag/Ag/Cl reference electrode), 1: metal lead (Ag), 2: metal salt (AgCl), 3: aqueous solution of KCl, 4: hydrogel-containing KCl solution, 5: solid melt of KCl, 6: junction containing diaphragm, porous ceramic or opening, 7: solid state reference element, 8: insulating encapsulation. Reprinted from [191] with permission of Springer. \u00a9 2008", + "texts": [ + " Other reported ionic liquids are tributyl (2-methoxyethyl) phosphonium bis(penthafluoroethanesulfonyl)amide [190] and bis(tritrifluoromethane sulfonyl) amides with 1-alkyl-3-methyl-imidazolium, phosphonium or ammonium cations [175]. Another type of RE was proposed using poly(n-butyl acrylate)-based membrane on top of glassy carbon electrode with polypyrrole as a transducing layer [177]. REs based on a polymer, an inorganic salt, and an Ag/AgCl element showed excellent stability, long lifetime, and insensitivity towards the ions in the sample. The basic principles of REs are discussed in [191] and shown in Fig. 2.9. In [78], an Ag/AgCl RE was coated with a polymer containing chloride ions, which was encapsulated by a Nafion outer layer. Nafion blocks the chloride anion diffusion to the test solution maintaining a constant chloride concentration on the AgCl electrode surface. An Ag/AgCl RE with a Nafion coating to increase its stability was reported in [124, 130]. Despite all the aforementioned effort in the development of a miniature RE, many issues related to the long-term behavior of these have not been solved yet. The electrochemical kinetics need to be investigated further to clarify the processes related with the mobility and exchange between the solids and dissolved ions. Figure 2.9 summarizes and illustrates the various methodologies followed for RE development. In the preceding sections, the technology regarding the potentiometric sensing element was discussed. It is obvious that these are not standalone devices, but they rather require additional circuitry in order to record their output signal. The following is a very brief discussion of the challenges in recording from potentiometric sensors and some examples of recording electronics used in the literature using commercially available discreet electronic components" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001718_s11431-011-4499-5-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001718_s11431-011-4499-5-Figure9-1.png", + "caption": "Figure 9 The 7-bar mechanism.", + "texts": [ + " Since the axes of M and K are always parallel to each other and JP=TH, PT is always parallel to JH, is equal to , which also means that is not an independent parameter. Therefore, gz IIIm (x0 0)=4. X IIIm = 12,8 IIm ( x 0 0)+ gz IIIm (x 0 0)= X IIIm (x 0 0)=4. FIII =6nIII X III P i i mpIII 1 =6\u00d715\u00d72+4=0, F=FI+FII+FIII=4+(1)+0=3. It can be seen that the chain PR(Pa)R generally considered as a hybrid chain formed by parallel and serial chains is actually a composite chain with an independent loop. It can be decomposed. The mobility can be also obtained from eq. (10). F=6n 1 ZP ii P L j X jm1 =6\u00d7125\u00d715+(2+0+4)=3. Example 6. Figure 9 illustrates the 3-loop 7-bar mechanism proposed in ref. [7]. In loop I, the four axes of link group RCCR are parallel to each other, every link can not rotate around z-axis or y-axis, the constraint of the virtue loop is X Im = Im (000 0)=2, FI= In6 1 IP X i Ii p m =6\u00d73(5\u00d72+4\u00d72)+2=2, 7,2 Im (0 0 0 0)=2. Loop II is formed by adding link group SRP at points E, F and G to links 2 and 7 of loop I. The common constraint of link group EFG is gz IIm (0 0 0 x 0 z)=2. X IIm = 7,2 Im (0 0 0 0)+ gz IIm (0 0 0x 0 z)= X IIm (0 0 00 0 0)=0, IIF = nII IIP i ip1 X IIm =6\u00d72(5\u00d72+3\u00d71)+0=1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000191_aim.2010.5695953-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000191_aim.2010.5695953-Figure4-1.png", + "caption": "Fig. 4. Point at minimum distance form the action lines of the normal vectors in the contact points", + "texts": [ + " To check the uniformity of the angular distribution of the normal vectors, the measure defined \u201carrangement of force directions\u201d in [10] can be computed, defined as \u0394\u03d5 = 1 n n\u2211 i=1 \u2223\u2223\u2223\u2223\u03b8i \u2212 2\u03c0 n \u2223\u2223\u2223\u2223 , where \u03b8i is the angle between the normal vectors for two successive points of the grasp polygon. The off-centering of the system of forces can be evaluated as cm\u2212cf , where cf is the point at minimum distance from the action lines of the normal vectors computed as cf = min p n\u2211 i=1 d(p, ln(pci)), where d(p, ln(pci)) is the distance between the point p and the line ln(pci), which is the line of action of the normal vector to the surface in the contact point pci (see Fig. 4). The object reported in Fig. 5 is considered, and an uniform mass distribution is assumed. On the top-right of the figure, the minimal inertia regions RI are shown, assuming a friction coefficient \u03bc = 0.4. The red parts denotes the selected regions while the regions in yellow are minimal inertia regions discarded due to their insufficient size with respect to the considered fingertip size, inaccessibility or excessive curvature. The center of mass cm is represented with the blue plus symbol. On the bottom-right of the figure, the resulting grasp regions RG are shown using alternatively black and green color to distinguish adjacent regions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001616_ut.2011.5774088-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001616_ut.2011.5774088-Figure1-1.png", + "caption": "Figure 1 Proposed autonomous amphibious vehicle", + "texts": [ + " In the vehicle fusion, sea-air fusion is the most difficult one. DARPA (Defense Advanced Research Project Agency) has been working on \u2018Submersible aircraft\u2019 or \u2018Flying submarine\u2019[8]. The submersible function requires neutral buoyancy for diving and strong enough chamber to protect the vehicle body in the water pressure. The flying function requires lighter density of the vehicle than a lifting force to fly. This confliction is extremely difficult to compromise. An alternative is required to realize the amphibious vehicle. As shown in Fig. 1, we propose a autonomous amphibious vehicle which covers sea-land-air;3-in-1 amphibious vehicle. In terms of the vehicle technology, the fusion of sea-land-air vehicles\u2019 functions could cover the most kind of the vehicle fusion. It has a dry cargo unit to carry UAVs and AGVs. The lateral thrusters are mounted to move as a hovering type AUV . In the bottom, a crawler unit is installed and that has two tracks with a camera dome. A funnel shape communication tower is mounted. As shown in Fig. 2, the vehicle could separate into two parts in underwater" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003345_aicai.2019.8701333-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003345_aicai.2019.8701333-Figure1-1.png", + "caption": "Fig. 1. Schematic diagram of x-z inverted pendulum", + "texts": [ + " Robustness of the controller is validated by comparing the results of proposed controller with PID controller suggested by Wang [7] for different pendulum masses.The work is organized in V sections. Section II gives the modelling of x-z inverted pendulum. The designing of RLQR-ANFIS controller is explained in Section III. The simulation results under different conditions are presented in Section IV and finally the paper is concluded in section V. II. DYNAMIC MODELLING OF X-Z INVERTED PENDULUM The schematic diagram of x-z inverted pendulum is shown in Figure 1. In Figure 1, \u2113 is the distance between cart hinge point to the centre of mass of the pendulum, M and m are the cart and the pendulum mass respectively. Further Fx and Fz represents forces on the cart in x and z directions respectively. Angle \u03b8 is the angle made by the inverted pendulum with the vertical axis. The Lagrange\u2019s equations of motion for XZIP is given as [7]: xFmlmlxmM 2sincos)( (1) zFgmMmlmlzmM )(cossin)( 2 (2) 0sinsincos gzxl (3) Using Eq. (1), (2) and (3) the state space representation of the system is given as: )mM(gF F 0Ml 1 00 )mM( 10 00 0M 1 00 z z x x 0 Ml )mM(g 0000 100000 000000 001000 0g M m 0000 000010 z z x x z x (4) Incorporating the parameters of x-z inverted pendulum (given in Table 1) in Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002772_ilt-12-2017-0368-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002772_ilt-12-2017-0368-Figure1-1.png", + "caption": "Figure 1 Schematic of deep groove ball bearing", + "texts": [ + " In the present study, Fast Fourier Transformation (FFT) method is used for accelerating the calculation of the elastic deformation, as this method has high computation efficiency (Meng and Chen, 2016, 2017; Ju and Farris, 1996; Hu et al., 1999; Wang et al., 2012). Meanwhile, the point heat source integration method is used to calculate the surface temperature of the bearing, as this method can analytically analyze the surface temperature rise (Tian and Kennedy, 1994; Chen et al., 2008; Meng, 2012). Then, the conclusions are drawn about effects of the external load, sliding speed and slide-roll ratio on TEHL performances for the deep groove ball bearing. The schematic of the deep groove ball bearing is shown in Figure 1(a). Points A and B in Figure 1(b) represent the contact edge of the enlarged inner raceway groove. In the present study, the TEHL model for the deep groove ball bearing is modeled based on the unified Reynolds equation and the point heat source integrationmethod. In solving this model, the friction heat is assumed to be absorbed by the contacting body, and the thermal convection and compression work of the fluid are ignored. For the film thickness from the middle layer of the lubrication film to the two solid surfaces, the temperature variation is assumed to be linear", + " As seen from Figure 6, the fluctuating pressure in the central contact region can be found because of the interference of the asperities on the surfaces of the rolling elements and raceway. When the load increases, the dimensionless pressure peak hardly varies but the actual dimensional pressure increases, as the value of ph is different at different loads and they have a positive correlation. Another phenomenon is that the maximum pressure occurs at the two edges of the inner raceway along the bearing width, which is perpendicular to the rolling direction, marked off with points A and B in Figure 1(b). For the deep groove ball bearing, the two edges of the raceway have upward curvatures, so that they contact easily with the rolling elements. In this case, the Figure 9 Dimensionless film pressures at different relative velocities Deep groove ball bearing Xuefang Cui, Fanming Meng, Delong K and Zhitao Cheng Industrial Lubrication and Tribology D ow nl oa de d by U ni ve rs ity o f Su ss ex L ib ra ry A t 0 7: 56 1 1 A ug us t 2 01 8 (P T ) lubricant flow is impeded because of the narrowed film thickness; thus, the hydrodynamic action of the lubricant is weakened and the large contact pressure appears", + "When the load increases, the clearance between the rolling element and inner raceway becomes small, and thus, the probability of the contact between them increases. In this case, the frictional heat\u2019 outflow from the contact region is rewarded because of the narrowed clearance; thus, the temperature increases. Therefore, appropriately choosing the external load for a deep groove ball bearing is necessary. It should be pointed out that at the two edges of the raceway illustrated with points A and B in Figure 1(b), there do not exist temperature peaks similar to the pressure peaks shown in Figure 6. This is due to the phenomenon that the temperatures at the two edges are equal to the surrounding temperature. In a similar behavior, the surface temperature of the rolling element increases with increasing load. This can be observed from Figure 8. Moreover, the maximum temperature occurs at the lubricant outlet. By the comparison between temperatures shown in Figures 7 and 8, one can found that the temperature of the rolling element is less than that of the inner raceway at the same load" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002455_1.4039395-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002455_1.4039395-Figure9-1.png", + "caption": "Fig. 9 UR10 robot", + "texts": [ + " We plan to incorporate efficient formulation of collision checking with other segments of the robot such as the manipulator links as our future work to augment the capability of the optimal motion planner. Journal of Mechanisms and Robotics JUNE 2018, Vol. 10 / 031010-7 Downloaded From: http://mechanismsrobotics.asmedigitalcollection.asme.org/ on 05/07/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use Funding Data Science and Engineering Research Council (Grant No. SERC 12251 00005). Singapore University of Technology and Design (Grant Nos. IDG31700107 and SUTD-ZJU/RES/03/2013). For robot with nonspherical wrist, we use Universal Robot UR10 [39] (see Fig. 9) as an illustration. Its manipulability measure is obtained as w \u00bc a2a3s3\u00f0a2c2 \u00fe a3c23 \u00fe d5s234\u00des5 (A1) To reduce the number of variables, denote a new variable C \u00bc a2c2 \u00fe a3c23 \u00fe d5s234 (A2) Then the manipulability becomes w \u00bc a2a3s3Cs5 (A3) Step 1: Solve for C From the robot\u2019s forward kinematics [2] and equating it with Eq. (10), we get Px X \u00bc \u00f0d4 \u00fe d6c5\u00des1 \u00fe c1\u00f0C d6c234s5\u00de (A4a) d6R13 \u00bc d6c5s1 d6c1c234s5 (A4b) Subtracting Eq. (A4b) from Eq. (A4a) gives Px X d6R13 \u00bc d4s1 \u00fe c1C (A5) Similarly Py Y d6R23 \u00bc d4c1 \u00fe s1C (A6) Squaring both Eqs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002077_ffe.12654-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002077_ffe.12654-Figure7-1.png", + "caption": "FIGURE 7 The equivalent stress field and functions of the equivalent and coordinate stress fields are plotted for the most critical loading case: Mx = 514.3 kNm, My = 2064.9 kNm, Mz = \u2212453.9 kNm, Fx = 101.3 kN, Fy = \u221234.1 kN, and Fz = \u2212329.5 kN. [Colour figure can be viewed at wileyonlinelibrary.com]", + "texts": [ + "21 Design loading case is defined by the geometry, the wind conditions (yaw position, pitch position of blade, etc), and the design situation (power of the wind turbine, start, stop, etc). The 3 moments and forces were prescribed on the analytical rigid body at the reference point as follows: Mx = 514.3 kNm, My = 2064.9 kNm, Mz = \u2212453.9 kNm, Fx = 101.3 kN, Fy = \u221234.1 kN, and Fz = \u2212329.5 kN. The most critical regions in the axle pin were detected from the stress field after computation. Bending loading dominates in this loading case. Figure 7 shows the equivalent von Mises stress field, and 2 profiles of the equivalent von Mises stress and the longitudinal \u03c322 are plotted in diagrams, as shown in Figure 8. Both stress profiles are taken for the outer surface of the axle pin and are plotted as a function of the y coordinate. The highest peak of the crack opening stress appears at 600 mm from the end of the axle pin around the axial bearing position, as shown in Figure 8. The most critical sections for crack initiation and propagation are defined from stress profiles. The highest stress \u03c322 is 135 MPa on the surface of the axle pin in the region of the detected defects, as shown in Figure 2B. The equivalent von Mises stresses are much lower than the yield stress. Figure 7 shows all the above\u2010mentioned stresses and the equivalent von Mises stress field at the FIGURE 8 The various stress distribution functions through the wall thickness at the most critical section. [Colour figure can be viewed at wileyonlinelibrary.com] critical location. The most critical crack opening\u2010bending stress \u03c322 appears on the upper surface of the axle pin. The crack opening stress is 135 MPa on the surface and on the inner surface \u221210 MPa in compression; therefore, the total stress amplitude is 145 MPa" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000013_j.proeng.2010.04.064-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000013_j.proeng.2010.04.064-Figure2-1.png", + "caption": "Fig. 2 (a) schematic and (b) actual traction rig.", + "texts": [ + " The traction features have a base diameter of 6 mm and a height of 4 mm. There are a total of 30 traction features per sole unit. Both sole units were assembled with a New Balance SDS 606 sprint shoe upper, UK size 9. The commercially available sprint shoes selected for testing were marketed towards 100m sprinters. The test shoes, UK size 9 (28 cm) are detailed in Table 1. The traction properties of both commercially available and SLS sprint shoes were evaluated using a purpose built test fixture (see Fig. 2). The test fixture was designed based on the ASTM standard for testing traction characteristics of the athletic shoe-sports surface interface (ASTM F 2333-04). The main features of the apparatus follow the guidelines of the ASTM standard. Some aspects were, however, modified. One modification is the manner in which the test shoe is secured to the apparatus. The standard test method ASTM F 2333-04 specifies that the test shoe be mounted on a foot form, creating a tight fit capable of properly transmitting forces through the shoe material to the outsole-playing surface interface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001232_tmag.2012.2198203-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001232_tmag.2012.2198203-Figure5-1.png", + "caption": "Fig. 5. Analyzed model (1/2 region).", + "texts": [ + " The motion equation of the mover is shown as follows: (8) (9) where and are the mass of mover, and are displacement of the mover in - and -directions, and are the viscous damping coefficient, and are the spring constants in - and -directions, and and are the electromagnetic force acting on the mover in - and -directions, respectively. Now because of difference of support structure mass have and component. The thrust of the mover is calculated using theMaxwell stress tensor method, and is substituted into (8) and (9). The position of mover is calculated by the time step. Fig. 4 shows the flowchart for this coupled analysis. The vector control is taken into consideration in this analysis. Fig. 5 shows the FEM model without air regions. The analyzed region is of the whole region because of the symmetry. The number of tetrahedron elements, edges, and unknown variables are about 306 400, 367 200, and 347 800, respectively. Table I shows the analysis conditions. The number of steps is 9900, time division is 100 s, and total CPU time is about 240 hours. Fig. 6 shows the prototype. As described at Section II-A, the -axis is supported with a flat spring and the -axis is supported by the coil spring and the mover can be moved independently on each axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002099_amm.868.124-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002099_amm.868.124-Figure9-1.png", + "caption": "Fig. 9 Nephogram of thermal deformation when a thermal equilibrium state is reached", + "texts": [ + " Namely, the temperatures increase sharply at the beginning and then gradually stabilize to the thermal equilibrium temperature when the heat generated by the components is equal to the heat dissipation into ambient. The temperatures measured, the simulated results with TCCs and without TCCs are compared. It can be seen that the temperatures with TCCs are about 6\u2103and 7\u2103 higher than the measured results for the front and rear bearings, respectively. The dynamic variation process of nut\u2018s temperature shows a similar trend. The thermal deformation of the ball screw feed drive system is shown in Fig. 9. It can be seen that the maximum deformation is 54.2\u00b5m and happens at the supported end of the feed drive system. The thermal deformation of the right side is larger than that of the left side because the left end of the screw is the fixed end and the right end of the screw is the supported end. The thermal deformation of the fixed end is restricted by the axial and radial loads and the supported end is only subjected to radial load. So the fixed end of the screw shaft cannot expand along the axial direction, while the supported end can expand along the axial direction, resulting in the accumulation of the axial thermal elongation at the supported end" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003921_b978-0-12-803581-8.11722-9-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003921_b978-0-12-803581-8.11722-9-Figure4-1.png", + "caption": "Fig. 4 Cylindrical viscous clutch with ER fluid: 1 \u2013 driving part, 2 \u2013 driven part, 3 \u2013 sealing ring, 4 \u2013 bearing, 5 \u2013 slip ring for supplying high voltage, 6 \u2013 insulator, 7 \u2013 bolt.", + "texts": [ + " Disk viscous brake with ER fluid works in such a way that the torque is transferred from the driving part to the driven part with the friction occurring as a result of shear stress in the working fluid. The shear stress value can be altered by changing the electrical field. An increase in voltage U applied to electrodes causes an increase in the intensity of the electric field E, which subsequently causes an increase in shear stress t in the ER fluid, simultaneously increasing the transferred torque M. Cylindrical Viscous Clutch With ER Fluid A cylindrical viscous clutch with an ER fluid shown in Fig. 4 is designed to be attached to a shaft of an electrical engine (K\u0119sy et al., 2008). The outer radius of the clutch is 160 mm, and its length is 90 mm. The clutch consists of two parts, driving and driven, consisting of connected rings of different radii. These parts, as electrodes, are isolated from each other and connected electrically with terminals of an electric high voltage power supply using electrical wiring and slip rings cooperating with carbon brushes. The driving part is set directly on the engine shaft" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001718_s11431-011-4499-5-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001718_s11431-011-4499-5-Figure2-1.png", + "caption": "Figure 2 The non-coincident links between every two adjacent loops.", + "texts": [ + " A link group is regarded as the combination of non-coincident links between any independent loop and its adjacent loop in a spacially multi-loop mechanism. The mobility of a link group can be equal to 0, or less than 0, or more than 0. The links of a link group can be driven link or driving link . The link group defined here is a generalized group, not the Assur group with zero freedom. For the mechanism shown in Figure 1, it contains four independent loops, and has four link groups, ABCD, EFG, HK and PRM, as shown in Figure 2. To describe the motion transmission manner of two ad- jacent loops by a unifying concept, the concept of virtual kinematic pair is defined here. Assume that links M and J are responsible for the motion transmission between two adjacent loops. No matter whether they are adjacent, it is supposed that they are connected by a kinematic pair, which is regarded as a virtual kinematic pair, short for virtual pair, denoted as , 1 . M J jG The virtual pair is a generalized pair. When M and J are adjacent, the virtual pair is a real pair and when they are not adjacent, the virtual pair is a generalized pair, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002466_tvt.2018.2800777-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002466_tvt.2018.2800777-Figure6-1.png", + "caption": "Fig. 6 Wheel testing bed", + "texts": [ + " IEEE TRANSACTION ON VEHICULAR TECHNOLOGY, VOL. X, NO. X, X 2017 4 m 2 1 m 2e 2e DPe e e 1e 1e cos sin d cos sin d F b r (12) m 1 2 m 2 e e e 2e 1e dT r b d (13) where We, FDPe, and Te are the vertical load, net traction force, and driving torque of the deformable mesh patterned wheel, respectively. In the experiment, the testing bed used for the deformable, mesh surface wheel contains two parts: the loading and testing unit, and the soil bin, as shown in Fig. 6. Experimental Wheel Loading and Testing Unit The parameters of the experimental soil are shown in Table I. The equivalent coefficients for the deformable, mesh surface wheel on the experimental soil can be determined through experiments, and the results are shown in Table II. equations for We, FDPe, and Te are defined as functions of s and \u03b8, and Eqs. (11)-(13) can be written as follows e e 1 DPe DPe 1 e e 1 , , , W W s F F s T T s (14) The four wheels of the manned lunar rover are driven by four independent motors" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002860_icuas.2018.8453322-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002860_icuas.2018.8453322-Figure2-1.png", + "caption": "Fig. 2. Relative geometry of the follower and the virtual target.", + "texts": [ + " Accordingly, the kinematics of the virtual target is calculated by[ xvt yvt ] = [ xl yl ] + [ cos\u03c8l sin\u03c8l \u2212 sin\u03c8l cos\u03c8l ] [ l\u0303 sin \u03b3\u0303 l\u0303 cos \u03b3\u0303 ] vvt = \u221a x\u03072vt + y\u03072vt \u03c8vt = arctan(x\u0307vt/y\u0307vt) \u03c6vt = arctan(\u03c8\u0307vtvvt/g) (5) where the subscript l denotes the parameters related to the leader, and the subscript vt denotes the parameters related to the virtual target. To accomplish maneuvering decision in the formation flight, a tailored scoring function is constructed to encourage the follower to maneuver to the tail of the virtual target. The scoring function is composed of contributions of an orientation score and a relative range score. Fig. 2 illustrates the geometric relationship of the follower and the virtual target during the flight. The vector connecting the centers of mass from the follower to the virtual target refers to the line of sight (LOS). The angles are shown from the point of view of the follower. Conventionally, the relative position of the UAVs is described in terms of the antenna train angle (ATA), the aspect angle (AA), and the range (R) between the centers of mass of the UAVs [16]. ATA is measured from the nose of the follower to the LOS, which informs the follower where the virtual target is" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002248_s00170-017-1085-4-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002248_s00170-017-1085-4-Figure7-1.png", + "caption": "Fig. 7 Equivalent parallel spring model", + "texts": [ + " The overall stiffness of the ith robot can be obtained by the addition of compliance matrix [23] K\u22121 i \u00bc K\u03b8 i \u22121 \u00fe KL i \u22121 \u00f028\u00de To obtain the stiffness of the entire multiple robot system, it is necessary to establish the deflection equation for the object held by the robots. Here, the system is decomposed into n single robots as shown in Fig. 3. In the decomposed model, the observation point C is shared by all robots expressed as Ci (i = 1, 2,\u22ef, n). Since the same observation point is used for all robots, a parallel spring system as shown in Fig. 7 can be used under which the deflection that each robot undergoes would be identical, that is DC \u00bc DC1 \u00bc DC2\u22ef \u00bc DC n\u22121\u00f0 \u00de \u00bc DCn \u00f029\u00de Then the total force at point C should be the summation of all the robot\u2019s end effector forces contributing to point C, i.e., Kall\u2022Dc \u00bc K1\u2022DC1 \u00fe K2\u2022DC2\u22ef\u00fe Kn\u2022DCn \u00f030\u00de In light of Eq. (26) and Eq. (27), the system stiffness of the coordinated multiple robot system can be obtained as Kall \u00bc K1 \u00fe K2 \u00fe\u22efKn \u00f031\u00de In this section, two coordinated robots are used to validate the system stiffness of a coordinated multiple robot system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002099_amm.868.124-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002099_amm.868.124-Figure7-1.png", + "caption": "Fig. 7 Nephogram of temperature field distribution of the feed drive system when thermal equilibrium", + "texts": [ + " The ball screw feed drive system continuously runs back and forth at a feed rate 8m/min till a thermal equilibrium state is reached. The heat generated by the ball screw and bearings, convective heat transfer, TCCs of solid joints and bearing stiffness can be obtained based on the methods discussed in Section 4 . Then the above boundary conditions are applied to the thermal characteristics analysis model to simulate the temperature field and thermal deformation of the ball screw feed drive system, as shown in Fig. 7. The maximum internal temperature of the ball screw feed drive system is about 49.1\u2103 (Front bearing). In addition, the temperatures of the bearings and nut are higher than that of other components because the bearings and nut are the main heat sources of the feed drive system. The temperature gradients of bearings / bearing housings, bearings / screw shaft, nut / worktable and nut / screw shaft are obvious. The cause is that the TCCs of solid joints restrict the heat generated by the bearings and nut flowing to bearing housing, screw shaft and worktable, which deteriorate the heat dissipation condition of the bearings and nut" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002636_j.mechmachtheory.2018.05.003-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002636_j.mechmachtheory.2018.05.003-Figure1-1.png", + "caption": "Fig. 1. Lima\u00e7on machine embodiments. Note how the rotor profiles differ in each embodiment: (a) Lima\u00e7on-to-lima\u00e7on machine, (b) Circolima\u00e7on machine, (c) Lima\u00e7on-to-circular machine.", + "texts": [ + " Background on the lima\u00e7on positive displacement machines Belongs to the rotary positive displacement machine category, the lima\u00e7on machine has gained its name from the unique motion of the machine rotor inside the housing; such motion always follows the lima\u00e7on curve (also referred to as Pascal\u2019s snail or snail curve), which can be produced by a number of mechanisms, some of which have been described to a reasonable level of detail by Sultan [10] . In further work by Sultan [11,13] and Phung et al. [7] it was pointed out that the rotor and housing of lima\u00e7on machines can be manufactured to either lima\u00e7on or circular curves; hence, the names: lima\u00e7on-to-lima\u00e7on machines [10\u201312] , circolima\u00e7on machines [13] , and the lima\u00e7on-to-circular machines [7] . As shown in Fig. 1 , the three machine embodiments share some similar operational characteristics. The rotor chord, p 1 p 2 , of length 2 L rotates and slides about the lima\u00e7on pole, o . As for the cases of the lima\u00e7on-to-lima\u00e7on and the lima\u00e7on-tocircular machines, the housing are the curves formed by the traces of point p 1 ( or p 2 ), namely lima\u00e7on curve, while the housing of the circolima\u00e7on is a circular curve. The machines\u2019 rotor takes the two-lobe form. The two lobes are mirrored images of each other in which each lobe is manufactured to either lima\u00e7on or circular curves. The centre point of the rotor chord, m , is restricted to move on a stationary circle of radius r , this circle is referred to as the lima\u00e7on base circle. A stationary Cartesian frame X o Y o is introduced and attached to the pole o; at the same time, a rotating frame X r Y r is attached to the rotor chord, p 1 p 2 , at its centre point, m , as shown in Fig. 1 . The rotor angular displacement, \u03b8 \u2208 [0, 2 \u03c0 ], is the angle rotated by the chord as the rotor performs its rotational motion, this angle \u03b8 is measured from the X o -axis to X r -axis in the anticlockwise direction. When the chord is sliding, its centre point, m , slides a distance s as a function of the rotor angular displacement, \u03b8 , with respect to the pole, o . This sliding distance can be formulated as follows [10,13] : s = 2 r sin \u03b8 (1) The radial distance of the housing, R h , measured from the pole, o , to the apex, p 1 , can be expressed as: R h = 2 r sin \u03b8 + L (2) which can also be described in the Cartesian coordinates, X o Y o , as: { x p 1 = R h cos \u03b8 = r sin 2 \u03b8 + L cos \u03b8 y p 1 = R h sin \u03b8 = r ( 1 \u2212 cos 2 \u03b8 ) + L sin \u03b8 (3) In order for the lima\u00e7on curve to be looping and dimple free, Costa et al", + " (3) or (4) as follows: \u03b80 = sin \u22121 ( \u2212 L 4 r + \u221a L 2 ( 4 r ) 2 + W g 4 r ) = sin \u22121 ( \u2212 1 4 b + \u221a 1 ( 4 b ) 2 + W g 4 Lb ) (7) With the case of the circolima\u00e7on and the lima\u00e7on-to-circular machines, where the rotor lobes are manufactured of circular arcs, the truncation, z , can be calculated as: z = L \u2212 C r \u2212 \u221a R lc 2 \u2212 ( k + W g 2 )2 (8) Let k s be the seal spring stiffness, and k w be the equivalent stiffness of the wall of the machine housing ( k w k s ); and let I be the instantaneous centre as shown in Fig. 4 . Of note is that local deformation at the seal-housing contact point has been taken into account in the equivalent stiffness coefficient, k w . Given that the initial deflection of the seal spring, \u03b4s , is known at the rotor radial displacement \u03b8 = 0 ; \u03bb measures the angle between the rotor chord, p 1 p 2 , (as shown in Fig. 1 a) and a line that connects the correspondent point of p 1 on the housing to the instantaneous centre, I (which is shown in Fig. 4 ). If \u03bc was the coefficient of friction between the housing and the seal, the initial forces exerted by the machine 1 housing and the spring on to the seal can be calculated as: { F cos \u03bb( 1 \u2212 \u03bc1 tan \u03bb) = k w \u03b4w F s = k s \u03b4s (9) where \u03bb = tan \u22121 ( 2 r cos \u03b8 2 r sin \u03b8+ L ) . At the position \u03b8 = 0 , \u03bb = \u03bb0 = tan \u22121 ( 2 r L ) \u03b4w and \u03b4s are the initial deflections of the housing wall and the spring, respectively", + " As a result of that, it is critical to reinforce the housing at these angular positions and strengthen the contact surface between the seal and the seal groove, to prevent failure during the machine operation and increase the reliability of the lima\u00e7on machine. It is also noticeable that the seal displacement in X direction ( Fig. 12 a) is of significant magnitude when the apex seal is approximately at the 3 \u03c0 2 (270 o ) position on the housing. This is due to the geometry of the lima\u00e7on machine. Based on Fig. 1 c, it is apparent that the rotor experiences a sliding motion in the positive X direction in the first quarter of the X o Y o frame. When the rotor chord approaches the second quarter of the X o Y o frame, the direction of the sliding motion shifts towards the negative X direction. At this instance, the left-hand side seal is now located at the 3 \u03c0 2 position on the housing. Due to inertia, the seal tends to move in the positive X direction and dig into the housing wall. This results in substantial seal displacement along X direction, which would require reinforcement to the correspondent section of the housing wall" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001425_1.4006251-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001425_1.4006251-Figure1-1.png", + "caption": "Fig. 1 Analytical model of the three-link underactuated manipulator", + "texts": [ + " In addition, we demonstrate through nonlinear bifurcation analysis that the two free links can be swung up and are stabilized at some configurations by the production of nontrivial and stable equilibrium points. Finally, we theoretically and experimentally clarify that the number of stable and nontrivial equilibrium points increases from 2 to 4 by increasing the excitation frequency. 2.1 Analytical Model and Equation of Motion. The analytical model of a three-link underactuated manipulator is shown in Fig. 1. An actuator of the manipulator is attached to the first joint, but the second and third joints do not have not only any actuators but also sensors. The manipulator moves on the x-y plane, which is inclined at angle a from the horizontal plane. Then, the manipulator experiences the gravitation g0 (\u00bcg sin a) in the positive x-direction. Parameters mi, li, and lig, in Fig. 1 are the ith mass, the length of the ith link, and the distance between the ith joint and the center of gravity of the ith link, respectively. Parameter Ii is the mass moment of inertia about the center of the ith link. Parameter s is a torque applied to the first joint. hi is the relative angle of the ith link with respect to the (i \u2013 1)th link (i\u00bc 2, 3). The parameter values corresponding to the experimental apparatus are as follows: m1 \u00bc 43:8 10 3 kg; m2 \u00bc 30:2 10 3 kg; m3 \u00bc 3:6 10 3 kg; l1 \u00bc 1:16 10 1m l2 \u00bc 0:51 10 1m; l3 \u00bc 0:525 10 1m; l1g \u00bc 5:53 10 2m; l2g \u00bc 2:54 10 2m l3g \u00bc 2:65 10 2m; I1 \u00bc 2:27 10 4 kg m2; I2 \u00bc 7:11 10 5 kg m2; I3 \u00bc 2:49 10 6 kg m2 g0 \u00bc 0:522 m=s2; a \u00bc 5:3 10 2rad The equations governing the relative angles of the second and third links can be expressed using Lagrange\u2019s formulation as follows: \u00f0A2 \u00fe b2 cos h3\u00de\u20ach2 \u00fe \u00f0A2 \u00fe b1 cos h2 \u00fe b2 cos h3\u00de\u20ach1 \u00fe \u00f0b2 cos h3\u00de\u20ach3 \u00fe b1 _h2 1 sin h2 b2\u00f0 _h1 \u00fe _h2 \u00fe _h3\u00de2 sin h3 \u00fe c2 sin\u00f0h1 \u00fe h2\u00de \u00bc 0 (1) A3 \u20ach3 \u00fe fA3 \u00fe b3 cos\u00f0h2 \u00fe h3\u00de \u00fe b2 cos h3g\u20ach1 \u00fe \u00f0A3 \u00fe b2 cos h3\u00de\u20ach2 \u00fe b2\u00f0 _h1 \u00fe _h2\u00de2 sin h3 \u00fe b3 _h2 1 sin\u00f0h2 \u00fe h3\u00de \u00fe c3 sin\u00f0h1 \u00fe h2 \u00fe h3\u00de \u00bc 0 (2) where constant parameters Ai, bi, and ci, are A2 \u00bc m2l2 2g \u00fe m3l22 \u00fe I2; A3 \u00bc m3l2 3g \u00fe I3; b1 \u00bc \u00f0m2l2g \u00fe m3l2\u00del1; b2 \u00bc m3l2l3g; b3 \u00bc m3l3gl1 c2 \u00bc \u00f0m2l2g \u00fe m3l2\u00deg0; c3 \u00bc m3l3gg0 For swing-up control, we set the configuration of the first link to be similar to that in the previous research [14,22] as follows: h1 \u00bc ae cos t (3) This provides high-frequency excitation to the second link (the frequency is referred to as the excitation frequency)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001718_s11431-011-4499-5-Figure10-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001718_s11431-011-4499-5-Figure10-1.png", + "caption": "Figure 10 The 3-RPS mechanism.", + "texts": [ + " The common constraint of link group EFG is gz IIm (0 0 0 x 0 z)=2. X IIm = 7,2 Im (0 0 0 0)+ gz IIm (0 0 0x 0 z)= X IIm (0 0 00 0 0)=0, IIF = nII IIP i ip1 X IIm =6\u00d72(5\u00d72+3\u00d71)+0=1. Loop III is formed by adding the link group 2P to links 5 and 7 of loop II. The common constraint of the link group KH is gz IIIm ( x 0 0)=4, 7,5 IIm ( x 0 z)=5. X IIIm = 7,5 IIm ( x 0 z)+ gz IIIm ( x 0 0)= X IIIm ( x 0 0)=4, IIIF = 6nIII IIIP i ip1 X IIIm =6\u00d715\u00d72+4=0, F=FI+FII+FIII=21+0=1, we can obtain from eq. (10) that F=6\u00d76(5\u00d76+4\u00d72+ 3\u00d71)+(2+0+ 4)=1. Example 7. Figure 10 is a 3-RPS mechanism with three uniformly distributed link groups. In each link group, the axis of R-pair is perpendicular to the axis of prismatic pair P. In loop I, there is no common constraint, X Im =0, FI = In6 IP i ip1 X Im =6\u00d75(5\u00d74+3\u00d72)0=4. The moving platform has three rotations, three translations and a local freedom around CD-axis, therefore, 6,3 Im (0 0 00 0 0)=0. In loop\u2161, link group GHK can not translate along x-, y-axes and the common constraint of link group GHK is gz IIm (0 0 0x0 0)=1, so X IIm = 6,3 Im (0 0 00 0 0)+ gz IIm (0 0 0x0 0)= X IIm (0 0 00 0 0)=0, IIF = IIn6 IIP i ip1 X IIm = 6\u00d72(5\u00d72+3\u00d71)+0=1, F=FI+FII=4+(1)=3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002149_gt2017-63495-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002149_gt2017-63495-Figure5-1.png", + "caption": "Figure 5 Schematic of the three-pad HAFB", + "texts": [ + " Each top foil has an arc angle of 115 degree, and hydrodynamic preload is 35 .m At the center of each pad, an orifice hole is drilled (1 mm diameter), and orifice tubes (stainless steel) are welded on the backside of the top foil concentric to the injection holes. Externally pressurized air is injected into the bearing clearance through those orifice tubes. Figure 4 shows the top foil with welded orifice tube. When the top foils are assembled on to the bearing sleeve, the orifice tubes are located at 60 , 180 (direction of gravitational loading) and 300 . Figure 5 shows orifice tubes configuration for the three-pad HAFB. Orifice tubes are connected to the main supply pressure line, and the controlled hydrostatic injection is achieved by controlling the air pressure to the orifice tube located at 180 through a separate on/off valve. Figure 6 shows the three-pad HAFB. 4 Copyright \u00a9 2017 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 08/23/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use The rotor is constructed with several components" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002870_joe.2018.8345-Figure12-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002870_joe.2018.8345-Figure12-1.png", + "caption": "Fig. 12 Vibration acceleration of the 4 pole 30 slot motor at 750 Hz (a) Motor C, (b) Motor D", + "texts": [ + " The electromagnetic force calculated by Maxwell was transferred to the harmonic response calculation module of the structure field, so as to perform the coupling analysis of the electromagnetic field and the structure field. The harmonic response analysis module was applied to solve the vibration displacement. By selecting the modal superposition method for calculation, the frequency response, vibration displacement, velocity, acceleration of the entire vibration of the stator, and vibration displacement, velocity, and acceleration at a certain node can be obtained. Fig. 11 shows the overall vibration acceleration of the stator at the slot frequency of 900 Hz for model motor A and model motor B. Fig. 12 shows the overall vibration acceleration of the stator at the slot frequency of 750 Hz for model motor C and model motor D. According to the simulation results, the motor C was a typical second-order vibration mode with a maximum amplitude of 2.9 \u00d7 10\u22128 m, located between the two magnetic poles. Motor D was also a typical second-order mode with a maximum amplitude of 9.1 \u00d7 10\u22129 m, centred on each pole. In the above two motors, the spatial distribution of the excitation force obtained by the analytical calculation of the electromagnetic excitation force was completely consistent with the vibration pattern obtained by the simulation calculation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001682_speedam.2010.5544954-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001682_speedam.2010.5544954-Figure2-1.png", + "caption": "Fig. 2 : Windings on the stator part", + "texts": [ + " This simplified radial structure is then used for the study of axial heat transfer, in the r-z plane. The study focuses on a thermal machine closed, cooled by water at the external frame. An internal convection is ensured by shaft mounted fans on both sides of the rotor (Cf. Fig.4). II. INVESTIGATED MACHINE The synchronous machine studied in this article is a new structure of flux switching synchronous machine with a hybrid excitation. This particular structure uses the principle of both flux switching and flux concentration. We developed a prototype (see Fig. 1 and Fig. 2). This machine is composed of a stator that includes armature coils, permanent magnets and a wound inductor. The salient rotor is simply made of stacked, soft iron steel. The prototype is a three-phase machine containing twelve magnets, with each phase composed of four magnets and four concentrated coils. The rotor contains Nr teeth (with Nr=10), and the relation between the mechanical rotation frequency F and the electrical frequency f can be expressed as: f = Nr F. We can see from the picture presented in Fig.2 that the length of the end-windings is as small as possible. Unlike conventional machines, it is not in the end winding that overheating will be most critical. In this structure, the magnetic losses in the rotor are difficult to evacuate, furthermore, we wish to study at high speed. We must add to the magnetic losses the air friction losses in the rotor-stator airgap. III. THERMAL ANALYSIS A. Cooling elements The machine is totally enclosed, which means that there is no air movement entering or leaving the middle between ambient and inside the machine" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001395_jsc.0b013e31821d97c0-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001395_jsc.0b013e31821d97c0-Figure1-1.png", + "caption": "Figure 1. Superior (top) and inferior (below) illustrations of the triangular-shaped platform. The top illustration depicts load cell placement; the bottom illustration shows the length of each side of the platform.", + "texts": [ + " On-line data collection produces a characteristic waveform or jump signature associated with each subject\u2019s effort and offers the athlete instant visual feedback on their performance. For each jump, the instrumented platform measures the maximum vertical height attained in 3 ways: as subjects take off from the platform, the amount of time spent in the air (hang times), and when they land on the device. High levels of data reliability and reproducibility were previously verified for jump heights measured by using the platform (4,5). Given its level of technology and sophistication, the platform (Figure 1) may yield more accurate vertical jump height values than a Vertec or a commercial force plate does. In turn, a more accurate measurement device such as the platform may alter what we know about the vertical jumpanthropometry relationship and yield more precise prediction equations. Another weakness with prior studies that examined the jump height-anthropometry relationship is a sample that comprised only one gender or a small size (1,22,24). Results from such studies, and inferences that can be drawn, are limited by a lack of data heterogeneity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003547_rnc.4696-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003547_rnc.4696-Figure1-1.png", + "caption": "FIGURE 1 The inertial and body-fixed frames of manned submersible [Colour figure can be viewed at wileyonlinelibrary.com]", + "texts": [ + " Finally, some conclusions are drawn in Section 6. In this section, the dynamic model of the manned submersible system is briefly analyzed. Generally speaking, the manned submersible system is treated as a six-degrees-of-freedom (6-DOF) rigid-body dynamic model subject to external disturbances and model uncertainties. First of all, two reference frames are defined, which are inertial reference frame Fn = {On, in, jn, kn} and body-fixed reference frame Fb = {Ob, ib, jb, kb}. The reference frames are depicted in Figure 1. The kinematics of the manned submersible is described by . \ud835\udf02 = J(\ud835\udf02)\ud835\udf10, (1) where \ud835\udf02 = [ x \ud835\udc66 z \ud835\udf11 \ud835\udf03 \ud835\udf13 ]T \u2208 R6 denotes the position and Euler angle vector in the inertial frame, and \ud835\udf10 = [ u v w p q r ]T \u2208 R6 denotes the linear velocity and angular velocity vector in the body-fixed frame. In addition, J(\ud835\udf02) \u2208 R6\u00d76 represents the rotation matrix defined by J(\ud835\udf02) = \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 C\ud835\udf03C\ud835\udf13 S\ud835\udf11S\ud835\udf03C\ud835\udf13 \u2212 C\ud835\udf11S\ud835\udf13 C\ud835\udf11S\ud835\udf03C\ud835\udf13 + S\ud835\udf11S\ud835\udf13 0 0 0 C\ud835\udf03S\ud835\udf13 S\ud835\udf11S\ud835\udf03S\ud835\udf13 + C\ud835\udf11C\ud835\udf13 C\ud835\udf11S\ud835\udf03S\ud835\udf13 \u2212 S\ud835\udf11C\ud835\udf13 0 0 0 \u2212S\ud835\udf03 S\ud835\udf11C\ud835\udf03 C\ud835\udf11C\ud835\udf03 0 0 0 0 0 0 1 S\ud835\udf11T\ud835\udf03 C\ud835\udf11T\ud835\udf03 0 0 0 0 C\ud835\udf11 \u2212S\ud835\udf11 0 0 0 0 S\ud835\udf11\u2215C\ud835\udf03 S\ud835\udf11\u2215C\ud835\udf03 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 , (2) where the compact notation C denotes for cos(\u00b7), S for sin(\u00b7), T for tan(\u00b7)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001718_s11431-011-4499-5-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001718_s11431-011-4499-5-Figure6-1.png", + "caption": "Figure 6 The 2-RPC/RRC mechanism.", + "texts": [ + " One axis at point D3 is attached to the moving platform, thus the rotation around the axis has two components, one is parallel to x-axis, the other is parallel to y-axis. The rotation around axis C3 has two components too, one is parallel to y-axis, the other is parallel to z-axis for the axis intersects the movable platform at an angle. As a result, link group 3 has three rotations and three translations. gz IIm (0 0 0, 0 0 0)=0. X IIm = JM Im , ( 0 x y z)+ gz IIm (0 0 0, 0 0 0)= X IIm (0 0 0, 0 0 0)=0, IIF = IIn6 X II P i i mpII 1 =6\u00d745\u00d75+0=1. F= FI+FII=4+(1)=3. We can obtain from eq. (10) that F=6\u00d7135\u00d715+(0+0)=3. Example 4. Figure 6 illustrates the 2-RPC/RRC mechanism generated by transforming one RPC of the 3-RPC mechanism into RRC in ref. [21]. The 2-RPC/RRC mechanism includes two independent loops. Loop I denoted by ABCGHM has no rotation around z-axis, 1X Im . From eq. (8), we get IF = In6 IP i ip1 + X Im =6\u00d75(5\u00d74+4\u00d72)+1=3. Link 3 has only a translating motion for its two rotation axes are parallel to plane O-xy, so 3,6 Im 6,3 Im ( 0 0 0). Loop II is formed by the link groups RPC and RRC, and the common constraint of RRC is gz IIm (0 0 0 0), X IIm = 6,3 Im (0 0 0)+ gz IIm (00 0 0)= X IIm (00 0 0)=2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001113_cca.2010.5611189-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001113_cca.2010.5611189-Figure6-1.png", + "caption": "Fig. 6. Kinematics of the traverse.", + "texts": [], + "surrounding_texts": [ + "Constant accelerations and decelerations for all actuators are assumed. This means we need to limit all speed changes. In case of maximal acceleration/deceleration at least one speed has maximal rate. We need to compute the scalar factor, such that at least one has maximal change if a maximal change per cycle is exceeded. The largest possible speed change within one cycle is: \u2206s\u0307maxi = \u2206T s\u0308maxi . (8) The actual speed change is \u2206s\u0307i = s\u0307i \u2212 s\u0307oldi . (9) The proportion of the actual possible change is \u03bbi = \u2206s\u0307maxi \u2206s\u0307i . (10) If \u2206s\u0307i = 0 then we set \u2206s\u0307i = \u2206s\u0307max, hence \u03bbi = 1. The smallest proportion factor will make sure that for desired maximal acceleration/deceleration at the end-effector/hook, at least one can have maximal rate: \u03bbmin = min ( 1, |\u03bbi| ) , (11) i.e. \u03bbmin is at most 1. Therefore, the actuator space parameters are: s\u0307i = s\u0307iold + \u03bbmin\u2206s\u0307i. (12) This causes a linear change in the change of the endeffector/hook velocity: v = vold + \u03bbmin\u2206v, (13) but nonlinear absolutely spoken. This holds, because the Jacobian is a linear transformation. I.e., that the same factor which reduces the speed changes in the actuator speed space also changes the change of end-effector/hook velocity." + ] + }, + { + "image_filename": "designv11_33_0002089_0954406217718857-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002089_0954406217718857-Figure6-1.png", + "caption": "Figure 6. Analysis of the forces acting on the gear shaft.", + "texts": [ + " The hydraulic force on a TSV yields F tsv h \u00f0 \u00de \u00bc p \u00f0 \u00de b azi\u00feaxk\u00f0 \u00de \u00f015\u00de The hydraulic force on the tooth tip yields the same as equation (15), and thus the hydraulic force yields F h \u00f0 \u00de\u00bc X F tsv h \u00f0 \u00de\u00fe X F tip h \u00f0 \u00de \u00f016\u00de As shown in Figure 2, the meshing force is in the direction along the meshing line, tangent to the base circle, which can be evaluated based on the torque of the ring gear4 Fm \u00bc M rb2 \u00bc poutb\u00f0r 2 f2 r2a2\u00de 2rb2 cos \u00f0 \u00dei sin \u00f0 \u00dek\u00f0 \u00de \u00f017\u00de Therefore, the radial force acting on the gear yields F r \u00bc F h\u00feFm \u00f018\u00de Driven by an electric motor, the gear shaft and the motor shaft are connected by a flexible coupling that allows small misalignments between them. Therefore, the shaft-motor coupling effect is not considered in this work, and the radial force Fr is assumed to be balanced by the bearing supporting forces as in previous studies21,25 as shown in Figure 6. The equilibrium of the gear shaft yields Fr \u00fe Fs1 \u00fe Fs2 \u00bc 0 Fs1 ds1 Fs2 ds2 \u00bc 0 \u00f019\u00de Note that the two journal bearings are symmetrically distributed; thus equation (19) yields Fs1\u00bc Fs2 \u00bc Fr=2 \u00f020\u00de The film part addresses the water film in the gear shaft/journal bearing interface involving the film geometry that is altered by the elastic deformation of the interface and the film pressure that generates the supporting force to balance the radial force. Film modelling. Figure 7 depicts the wedged film in the gear shaft/journal bearing interface in an exaggerated way in a bearing Cartesian reference system (xbybzb) with respect to the bearing centre" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000300_978-3-642-03895-2_52-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000300_978-3-642-03895-2_52-Figure4-1.png", + "caption": "Fig. 4 Dynamic tilt table Erigo with the input tilt angle \u03b1tilt and resulting ECG signals", + "texts": [ + " Cardiovascular adaptation depends on the proper interplay of the hemodynamic circulation system and the reflex mechanisms that maintain blood pressure homeostasis [4, 5]. Thus, heart rate (HR) and heart rate variability (HRV) change with body posture. Controlling HR (and HRV) can be important to limit stress and arousal during motor treatment. The dynamic tilt table Erigo IFMBE Proceedings Vol. 25 (Hocoma, Switzerland) is used to investigate and control the relationship between body posture and heart rate. Subjects are tightly fixed to the Erigo by a belt system, and they can be tilted between an inclination angle \u03b1tilt of 0\u00b0 and 76\u00b0 (Fig. 4). Using both an inverse human cardiovascular model and a proportional controller, the \u201cmodel-based controller\u201d determines the required angle \u03b1tilt to track a desired heart rate (Fig. 5). The HR is extracted online from ECG, which is acquired by a PowerLab amplifier (ADInstruments, Germany) and an ECG recording system. The HR controller was evaluated in an experimental study with a single healthy subject. The desired HR is a step function, with a step from 0 to 72 bpm after 30 s (Fig. 5). Following the inverted model within the controller, the tilt angle of the Erigo initially increases to 36\u00b0, and the HR responds with an overshoot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000840_ipemc.2012.6258967-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000840_ipemc.2012.6258967-Figure1-1.png", + "caption": "Fig. 1 Cross-section diagram of SRM", + "texts": [ + " This paper presents an improved this CSI circuit to apply the SRM driving. This CSI can flow the square current wave form without the voltage spike and it is easy to expand the multiphase SRM driving. In this paper, the brief characteristics of the proposed CSI circuit are shown by the circuit simulation and some SRM drive experimental results are also shown. 1087 978-1-4577-2088-8/11/$26.00 \u00a92012 IEEE II. DRIVING METHOD OF SRM A cross-section diagram of SRM which has 4 poles and 6 slots is shown in Fig. 1. Transition diagram of inductance with respect to each rotor position is shown in Fig. 2. From Fig. 2, the inductance value is linearly varied with respect to each rotor position. However actually, this value is varied non-linearly. For this reason, it is difficult to control the current with conventional linear feedback control. B. Voltage source inverter (VSI) Conventional SRM drive method is the current hysteresis control by using the voltage source inverter. The circuit diagram, control block diagram and load current waveform of this drive method are shown in Figs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001620_saci.2011.5873037-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001620_saci.2011.5873037-Figure1-1.png", + "caption": "Figure 1. The OBBs models for the articulated arm robots and for the moving object; a) moving object; b) link 2 of the arm robot.", + "texts": [ + " The Cartesian position of the moving obstacles is detected using the vision architecture developed by Grigorescu [7] that has the scene geometry estimation and the scene understanding modules. It will be considered that the moving obstacle is in collision with the articulated arm robot if the smallest distance between robot components and the obstacle is smaller than min_d . The shape of the moving obstacle, and also the shape of the arm robot components, can be very complex. To simplify the collision detection procedure, the arm robot components and the moving objects were represented using oriented bounding boxes (OBBs) [6]. As is represented in figure 1 a), it was created an OBB model for the moving object. The model contains an array of small triangles that represents the surface of the object. To detect if the distance between the robot and object is smaller than min_d the surface of the object model is bigger, with min_d , than the real surface. The links, respective the joints, of the arm robot have their own OBB models, as is represented in figure 1 b). Having the position of the OBB models, it was detected, using the interference detection method [6] elaborated by Gottschalk, if there is a collision. If the collision occurs the method determines which triangles of the two OBB models have common points. Each triangle of the moving obstacle OBB model has a normal vector that respects the following constraints: \u2022 it starts from the centre of the incircle; \u2022 it is perpendicular on the surface of the triangle; \u2022 it is orientated from the inside to the outside of the obstacle. It will be considered that the vector of the m moving obstacle OBB triangle has the orientation: kzjyixv mmmm ++= (1) and it passes through the point mS of the arm robot base Cartesian frame (which is considered as reference frame). The vectors i , j and k are the normal vectors of the reference frame. The surface of the moving can be very complex and, as a result of that, it contains many small triangles. In the example from figure 1, the surface of the moving obstacle is very simple and in OBB models collision are involved only two triangles. III. COLLISION AVOIDANCE When the control structure know that the arm robot it is in collision, it determines the triangles of the moving obstacle OBB model that have common points with the arm robot OBB model. In the next step, it determines the orientation of the normal vector from the collision surface of the moving obstacle, as: n v v m c = (2) where mv represent the moving obstacle OBB triangles that are in collision and n is the total number of collision triangles" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002165_j.jmapro.2017.08.008-Figure14-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002165_j.jmapro.2017.08.008-Figure14-1.png", + "caption": "Fig. 14. (a) A fabricated portion of example object 3, (b) Support material removed easily from model; Support generated by (c) MakerBot, (d) Meshmixer, (e) Proposed method.", + "texts": [ + "0 min to fabricate the whole object along ith support as shown in Fig. 13(f). To demonstrate the support emovability, a portion of this object containing the most free form urface is printed with model and support material as shown in ig. 14(a and b). The support structure is easy to remove due to he effectiveness of the partial contact support technique as disussed in section 2.5. Comparing with the commercial machine, he proposed methodology shows 14% improvement of total build ime (TBT) where 53% savings of support material as compared in able 4. Fig. 14(c\u2013e) shows the comparison of the support structure etween commercial support generators and proposed methodlogy. The comparison with respect to support volume, contour umber, and build time between commercial support generators nd proposed methodology is shown in Table 5. The implementation sequence of the proposed methodology on xample 4 is shown in Fig. 15(a\u2013e). The normal vector parameter ange is calculated as 0 \u2212 0.85, which is divided into four groups o generate coherent region as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003894_ecce.2019.8913269-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003894_ecce.2019.8913269-Figure7-1.png", + "caption": "Fig. 7: Magnetic flux density distribution at time zero under 68.75% dynamic eccentricity, (a) conventional TSFE analysis, (b) VBR method.", + "texts": [ + " permeabilities calculated in the frequency domain are valid for the given static eccentricity fault severity and corresponding induction motor operating point. Static eccentricity fault frequency components (SE1, SE2) calculated by (1) extracted from stator currents using Fast Fourier Transform (FFT) are given in Fig. 6. V. DYNAMIC ECCENTRICITY The Dynamic eccentricity with 68.75% (0.55 mm displacement of rotor axis in the positive \u201cX axis\u201d direction in the FE grid) was applied. The magnetic flux density distribution of the case-study induction motor at time zero simulated by means of conventional TSFE and VBR method are given in Fig. 7a and Fig. 7b, respectively. Again, all MVPs are assumed to be zero in the conventional TSFE analysis which result-in zero magnetic flux density in the entire geometry of the motor. This leads to a lengthy numerical transient response. It can be seen in Fig. 7b that the permeabilities of the induction machine were successfully calculated considering the dynamic eccentricity and operating point (see the asymmetry between magnetic flux density distribution in areas 1 through 4). The torque profile extracted form TSFE simulation with and without imported permeabilities from the VBR method is given in Fig. 8, while their required number of AC cycles of TS-FE simulation and the corresponding CPU times to convergence are listed in Table III. Accordingly, the required number of AC cycles of TS-FE simulation and the corresponding CPU times to convergence are decreased by 97" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003677_i2mtc.2019.8826954-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003677_i2mtc.2019.8826954-Figure2-1.png", + "caption": "Fig. 2. Gearbox fault simulation platform and data acquisition system", + "texts": [ + " EXPERIMENT EVALUATION AND RESULT DISCUSSION The experiments in this study are carried out on the INV1618 gearbox fault simulation test platform of the China Eastern Noise and Vibration Research Institute. A highsensitivity resonant acoustic emission sensor is mounted on the surface of the gearbox for actual data acquisition. The data acquisition system is built by using NI's c-DAQ module. The INV1618 gearbox test bench can be simulated by the knob to simulate three gear working modes: normal gear, weak gear and broken gear. The gearbox fault simulation platform and acoustic emission data acquisition system are shown in Fig. 2. The acoustic emission sensor is mounted on the side of the gearbox as shown in Fig. 3. Before starting the acoustic emission signal acquisition, the parameters are set by the INV1618 gearbox fault simulation platform, including the working mode, gearbox speed, and sampling rate setting. The parameter setting for this experiment is that the gearbox speed is 1000 rpm and the data acquisition system has a sampling rate of 1 MSa/s. Then the working mode knob of the gearbox test platform is adjusted to work in three modes: normal, weak gear and broken gear" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000907_s0361521912040039-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000907_s0361521912040039-Figure1-1.png", + "caption": "Fig. 1. Schematic diagram of the laboratory setup for the preparation of globular carbon: (1) piston pump (PMS 50), (2) com pressor, (3) tubular quartz reactor, (4) nozzle, (5) electric furnace, (6) air cooler, and (7) filtering bag.", + "texts": [ + " Kryazhev and coauthors reported [3, 4] that poly vinylene chlorides form amorphous, in particular, microporous carbon materials as a result of thermal dehydrochlorination with intermolecular condensa tion under unusually mild thermal treatment condi tions (about 200\u00b0\u0421) according to the following reac tion scheme: . Data obtained with the use of the dehydrochlorina tion of CPVC (Cl content of 62 wt %) in 1% solutions in tetrahydrofuran in the presence of potassium hydroxide at 20\u00b0\u0421 for 6 h in accordance with a pub lished procedure [3] as an example are given below. Nanoglobular carbon, which was prepared by the thermal oxidative pyrolysis of heavy catalytic cracking gas oil (TU [Technical Specifications] 38.301 19 87 97, amendments 1\u20133) in a laboratory setup (Fig. 1) [5], was used as a dispersed carbon component. The hydrocarbon raw material as an aerosol was injected with compressed air through a nozzle into a tubular Cl Cl Cl Cl Cl Cl Cl Cl Cl DOI: 10.3103/S0361521912040039 272 SOLID FUEL CHEMISTRY Vol. 46 No. 4 2012 ANIKEEVA et al. quartz reactor. The feed rate was 10 cm3/min. Before the onset of the process, the reactor was heated to a starting temperature of 1200\u00b0\u0421 (the high temperature pyrolysis of hydrocarbons with the formation of glob ular dispersed carbon occurs due to heat released upon the combustion of a portion of the raw material)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003972_j.ifacol.2019.12.330-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003972_j.ifacol.2019.12.330-Figure2-1.png", + "caption": "Fig. 2. Top view of the simulated USV (Sarda et al., 2016).", + "texts": [ + " At each corner the commanded heading angle increases step-wise by +\u03c0/2. To ensure its smoothness, the reference trajectory \u03b7d(t) tracked by the USV is obtained by passing the nominal trajectory through a fourth order linear filter, which is comprised of a cascaded pair of second order filters (Fig. 1). Each second order filter has a damping ratio of \u03be = 1.0 and a natural frequency of \u03c9n = 1.0 rad/sec. The USV simulated is a catamaran with an overall length of 4 m, an overall beam of 2.44 m and a mass of 180 kg (Fig. 2). The manuevering coefficients and physical characteristics of the USV are obtained from Sarda et al. (2016). Exogenous disturbances are modeled as first order Markov processes Du et al. (2016) of the form Tbb\u0307 = \u2212b + anwn, where b \u2208 R3 is a vector of bias forces and moments, Tb \u2208 R3\u00d73 is a diagonal matrix of time constants, wn \u2208 R3 is a vector of zero-mean Guassian white noise, and an \u2208 R3\u00d73 is a diagonal matrix that scales the amplitude of wn. To provide a basis of comparison for the performance of the higher order controller-observer in (7) and (13), simulations were also conducted using a PID controller combined with the NDO developed by Chen et al", + " At each corner the commanded heading angle increases step-wise by +\u03c0/2. To ensure its smoothness, the reference trajectory \u03b7d(t) tracked by the USV is obtained by passing the nominal trajectory through a fourth order linear filter, which is comprised of a cascaded pair of second order filters (Fig. 1). Each second order filter has a damping ratio of \u03be = 1.0 and a natural frequency of \u03c9n = 1.0 rad/sec. The USV simulated is a catamaran with an overall length of 4 m, an overall beam of 2.44 m and a mass of 180 kg (Fig. 2). The manuevering coefficients and physical characteristics of the USV are obtained from Sarda et al. (2016). Exogenous disturbances are modeled as first order Markov processes Du et al. (2016) of the form Tbb\u0307 = \u2212b + anwn, where b \u2208 R3 is a vector of bias forces and moments, Tb \u2208 R3\u00d73 is a diagonal matrix of time constants, wn \u2208 R3 is a vector of zero-mean Guassian white noise, and an \u2208 R3\u00d73 is a diagonal matrix that scales the amplitude of wn. To provide a basis of comparison for the performance of the higher order controller-observer in (7) and (13), simulations were also conducted using a PID controller combined with the NDO developed by Chen et al. (2000). The PID control law has the form \u03c4PID = \u2212d\u0302\u2212 JT (\u03b7) ( Kp\u03b7\u0303 +Kd \u02d9\u0303\u03b7 +Ki \u222b t 0 \u03b7\u0303dt ) , (17) Fig. 2. Top view of the simulated USV (Sarda et al., 2016). where Kp = KT p > 0, Kd = KT d > 0 and Ki = KT i > 0 are diagonal controller design matrices (and elements of Rn\u00d7Rn). The disturbance estimate for the PID controller is computed as d\u0302 = q(t) +K0Mv (18) q\u0307(t) = \u2212K0q(t)\u2212K0 [N(v)v + \u03c4 +K0Mv] , where K0 = KT 0 > 0, K0 \u2208 Rn \u00d7 Rn. Three simulation cases were studied using each controller: (1) no disturbances; (2) continuous disturbances with Tb = 103\u00b716\u00d76, b0 = \u2212an and an = [1.75 2.35 2.35 0.25 1.20 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000276_milcom.2009.5379819-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000276_milcom.2009.5379819-Figure2-1.png", + "caption": "Figure 2: Paparazzi System Components [Source : http : / /paparazzi.enac .fr/ reproduced with courtesy of ENAC]", + "texts": [], + "surrounding_texts": [ + "Paparazzi [15] is an open-source project started in 2003 whose goal is to design and build a cheap autopilot for fixed wing autonomous UAVs. The autopilot component has al ready been used on more than 15 airframes by several teams around the world. Hundreds of hours of autonomous flight have successfully been achieved with it. The Paparazzi system includes the airborne hardware, the airborne autopilot software, a ground control station, the communication protocols linking the different components and a simulation environment. Safety has been one of the main concerns during all the phases of its development : the airborne code has been made as simple and short as possible; software and hardware seg regation ofthe critical code ensures a good level of reliability of the system. The flight plan language (XML-based) allows the user to define complex autonomous missions while taking external events into account. The ground control station is based on a client/server architecture which enables to control one or several aircraft from one or several locations. The ground station operator can control the aircraft with high level com mands (e.g. area coverage patterns)." + ] + }, + { + "image_filename": "designv11_33_0003269_s42417-019-00086-4-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003269_s42417-019-00086-4-Figure2-1.png", + "caption": "Fig. 2 3D geometrical models of (a) the PGJB rotor system and (b) the plain gas film", + "texts": [ + " The calculating process of the FFSIM is realized in an automated manner (2) \ud835\udf15 ( \ud835\udf0cv\u20d7 ) \ud835\udf15t + \u2219 ( \ud835\udf0cv\u20d7\u2297 v\u20d7 ) = \u2212\u2207p + \u2207 \u2219 \ud835\udf0f + \ud835\udf0cv\u20d7 + F\u20d7, (3)\ud835\udf0f = \ud835\udf07 ( \u2207v\u20d7 + ( \u2207v\u20d7 )T \u2212 2 3 \u2207 \u2219 v\u20d7I ) , (4) \ud835\udf15 ( \ud835\udf0c ( h + 1\u22152v\u20d72 )) \ud835\udf15t + \u0394 \u2219 ( \ud835\udf0cv\u20d7 ( h + 1\u22152 | | | v\u20d7 2|| | )) = \u2207 \u2219 (\ud835\udf06\u2207T) + \u2207 \u2219 ( v\u20d7 \u2219 \ud835\udf0f ) + v\u20d7 \u2219 ( \ud835\udf0cv\u20d7 + F\u20d7 ) + S E , (5)[M]{u\u0308} + [C]{u\u0307} + [K]{u} = {F(t)}, in ANSYS. The flow chart for calculating process of the FFSIM is shown in Fig.\u00a01. Geometrical Model of\u00a0the\u00a0PGJB Rotor System and\u00a0the\u00a0Corresponding Boundary Conditions Based on\u00a0the\u00a0FFSIM The PGJB rotor system studied in this paper is a singlespan single-disc rotor supported by two PGJB, as shown in Fig.\u00a02a. The nominal radial clearance of the PGJB is 50\u00a0 m . Diameter and width of the disc are 100\u00a0mm and 30\u00a0mm, respectively. Axial length and diameter of the rotor axis are 200\u00a0mm and 50\u00a0mm, respectively. Thickness, diameter and length of the bearing sleeve are 20, 50.1 and 60\u00a0mm, respectively. Diameter of the pressure released vent is 3\u00a0mm. For simulating the accurate nonlinear dynamic characteristics of the PGJB rotor system shown in Fig.\u00a02, the FFSIM is applied in the commercial software ANSYS (Fig.\u00a01), in which the fluid model is the gas film model and the structural model is the rotor-bearing sleeve model. The shared boundaries for these two models are the outer surface of the rotor shaft and the inner surface of the bearing sleeve. These surfaces act as the free fluid-structure interaction interface to exchange data of pressure loadings and structural deformations between these two models. 1 3 Detailed boundary conditions of the FFSIM are as follows: (1) the rotor shaft and the bearing sleeve are flexible structures; (2) the wall of the gas film model is treated as a flexible wall; (3) the flow condition is laminar and isothermal; (4) the gas inlet is complete self-absorption and the change of the gas inlet and outlet depends on the rotating state of the rotor; (5) the rotor shaft starts its motion from a concentric location within the bearing under action of the gravitational force; and (6) the bearing eccentric location varies with the gravitational force and the rotating speed", + " In the analytical method, the nonlinear dynamic characteristics of the PGJB rotor system are obtained by solving the nonlinear dynamic governing equations through the Runge\u2013Kutta method or the finite difference method. For details of the analytical method, Ref. [50] can be referred. First, bifurcation analysis of the PGJB rotor system is performed by the FFSIM. The rotating speed n and mass mr of the rotor are taken as the bifurcation parameters. Here, only several cases of n and mr are chosen for the bifurcation analysis. Bifurcation diagrams of displacement in the x direction (Fig.\u00a02) against rotating speed n =8000, 10,000, 12,000 and 14,000\u00a0rpm, at mr= 2\u00a0kg, obtained the FFSIM are shown in Fig.\u00a05a. For comparison, at mr= 2\u00a0kg, the bifurcation diagrams of displacement in the x direction against rotating speed 6000 \u2264 n \u2264 14000 rpm by the analytical method are shown in Fig.\u00a05b. From Fig.\u00a05a, b, it can be seen that at n =8000 and 10,000\u00a0 rpm, the dynamic center of the rotor center is the T-periodic motion. At n =12,000 and 14,000\u00a0rpm, the 2T-periodic motion and 3T-periodic motion occur, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002248_s00170-017-1085-4-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002248_s00170-017-1085-4-Figure9-1.png", + "caption": "Fig. 9 Three clamping poses of two coordinated robots", + "texts": [ + " A simulation model is created inMATLAB to compute the deflection of the point C on the workpiece with an external loading based on our developed model. The same system is modeled in ADAMS with the same external load. Then, the two results are compared. In general for industrial robots, the deflection caused by link flexibility is much smaller than that by joint flexibility. So in this simulation, the links are considered rigid and the flexible joints are modeled as torsion springs. According to Guerin\u2019s research [24], the stiffness values are given in Table 1. As shown in Fig. 9, three clamping poses are selected. In this simulation, the distance of two robots bases is set to 2.5 m, and a force is applied at the center of the clamped plate with its value given in Table 2. The simulation is first performed in ADAMS, and at the same time, the theoretical deformation using our model is computed using MATLAB. Figure 10 and Fig. 11 show the results of the two simulations. It can be seen from Figs. 10 and 11 that the theoretical results are very close to the simulated results and the average error between the two results is less than 5%" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003921_b978-0-12-803581-8.11722-9-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003921_b978-0-12-803581-8.11722-9-Figure2-1.png", + "caption": "Fig. 2 Structure of a typical hydrodynamic clutch: 1 \u2013 housing, 2 \u2013 pump, 3 \u2013 turbine, 4 \u2013 sealing ring, 5 \u2013 input shaft, 6 \u2013 output shaft.", + "texts": [ + " In hydrodynamic clutches and brakes the rotors are connected by the fluid\u2019s hydrokinetic impact on the rotor blades. Depending on the build of the rotors, one can distinguish between hydrodynamic clutches and brakes whose rotors contain inner rings, and those whose rotors have no inner rings. There is also a distinction depending on the build of the housing: with a fixed housing or a movable housing. Most commonly used are the hydrodynamic clutches and brakes with rotors with radial-axial flow with flat radial blades, Fig. 2. Hydraulic clutches and brakes are widely used in power transmission systems, such as vehicles, hosting cranes, conveyor belts, pumps, air blowers, mills etc., mainly to connect rotating shafts of various speeds, or as soft start clutches or retarders. The ER fluids used as working fluids in hydraulic clutches and brakes can be categorized as homogeneous or heterogeneous fluids, depending on their composition. Homogeneous fluids are uniform, while heterogeneous fluids are mixtures consisting of solid particles, base fluid and a small amount of additives preventing sedimentation and agglomeration of the solid particles (Dong et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001633_jmems.2010.2100034-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001633_jmems.2010.2100034-Figure1-1.png", + "caption": "Fig. 1. Three-dimensional schematic of a flexible membrane with a partially buried MEMS structure. (a) Fabricated planar MEMS structure is partially buried in a planar but flexible membrane. (b) When the membrane deforms, the whole structure can change to any curvilinear shape following the membrane.", + "texts": [ + " The thickness of the structures and devices implemented by these techniques is up to 2 \u03bcm, which is enough for electronic devices but significantly limits the applications for microelectromechanical systems (MEMS) devices. In this paper, we present a method to fabricate complex structures in a large area on nonplanar surfaces by utilizing a transfer method. Different sizes of 3-D and high-aspect-ratio MEMS structures and devices could be transferred to flexible membranes and could further be transferred to any nonplanar surfaces through this method. Fig. 1 shows the 3-D schematic of the deformation of a flexible polymer membrane with a complex MEMS structure partially buried in it. The square pillar array represents a high-aspect-ratio MEMS structure or device. After the fabrication of the planar MEMS structure, it is partially buried into a polymer membrane, formed surrounding it. Manuscript received October 5, 2010; accepted November 21, 2010. Date of publication January 10, 2011; date of current version February 2, 2011. This work was supported by the U" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002977_icstcc.2018.8540739-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002977_icstcc.2018.8540739-Figure3-1.png", + "caption": "Fig. 3. Rotor flapping approximation [22]", + "texts": [ + " The longitudinal and lateral dihedral derivatives are equal in magnitude and both cause the rotor to flap away from the incoming air. \u03b4b1i \u03b4\u00b5v = \u2212\u03b4a1i \u03b4\u00b5 (17) The upward heave movement of the rotor causes a higher lift on the advancing blade which causes a moment on the rotor hub. The same stabilizer scaling coefficient is used: \u03b4a1i \u03b4\u00b5z = K\u00b5 16\u00b52 ri (1\u2212 \u00b52 ri/2)(8 |\u00b5ri|+ ar\u03c3r) (18) Rotor flapping is the dominant effect on rotor moments. The restraint is approximated using a linear torsional spring with constant stiffness coefficient K\u03b2 . This is illustrated in Figure 3. This results in a longitudinal (pitch) and lateral (roll) moments: Mk,lon = K\u03b2a1i (19) Lk,lat = K\u03b2b1i (20) Once flapping occurs, the rotor thrust vector tilts and contributes to the body moments. Assuming the thrust vector tilts proportionally to the rotor flapping angles, the total rotor pitch and roll moments can be deduced as (N = 4 for a quadcopter): Lr = \u2211N i=1 (K\u03b2 + Thr) b1i (21) Mr = \u2211N i=1 (K\u03b2 + Thr) a1i (22) Nr = \u2211N i=1Qi (23) where hr is the distance between the rotor head and the center of gravity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000971_tie.2011.2158033-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000971_tie.2011.2158033-Figure7-1.png", + "caption": "Fig. 7. Schematic diagram for the magnetic bearing system.", + "texts": [ + " Satisfactory results are obtained for time-varying systems in the presence of model uncertainties and unknown disturbances. In this example, the proposed MLFC-MISO controller is compared with a nonlinear controller for a magnetic bearing system, which is used in various industrial applications. The magnetic bearing system is a typical MISO nonlinear system, which uses magnetic forces to suspend a rotor shaft in the middle of air. A schematic diagram of a single axis of the magnetic bearing system is shown in Fig. 7. Each electromagnet comprises multiple turns of conducting coil around the highly permeable magnetic core, and the mass of the rotor is m. The deviation of the rotor from the centered position is denoted as x, which is referred as \u201cposition variation.\u201d The nominal distance of the rotor away from the magnets is h. The coil currents are denoted as I1 and I2, and the input voltages at the coil terminals are E1 and E2. The magnetic bearing system is modeled as a fourth-order two-input nonlinear system, which can be described in the following state-space form [18]: y\u03071 = y2 y\u03072 = \u2212 k 2m [ I2 1 (h + y1)2 + I2 2 (h \u2212 y1)2 ] I\u03071 = \u2212 R k (h + y1)I1 + (h + y1)\u22121y2I1 + 1 k (h + y1)E1 I\u03072 = \u2212 R k (h \u2212 y1)I2 \u2212 (h \u2212 y1)\u22121y2I2 + 1 k (h \u2212 y1)E2 (25) where the four state variables are x = [x1, x2, x3, x4] = [y1, y2, I1, I2] and y1 is the position variation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000041_cae.20391-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000041_cae.20391-Figure1-1.png", + "caption": "Figure 1 Typical conventional MPS stations on separate platforms.", + "texts": [ + " MPS units involve many automation technologies such as mechanics, Correspondence to R. Yenitepe (ryenitepe@marmara.edu.tr). 2009 Wiley Periodicals Inc. 1 pneumatics, electro-pneumatics, electrical and electronics engineering, sensors, safety and drive technology, PLC technology, industrial communication means and computers [10,11]. However, conventional MPS units have several independently working stations each of which is controlled by its own PLC as shown in Figures 1 and 2. The conventional MPSs shown in Figure 1 are placed on separate platforms. Figure 2 shows a variation of the above MPS stations. Although this type of conventional MPS is placed on a platform, it has still several independently working PLC systems. Obviously, these types of lab setups require more space and high cost. Therefore, these systems do not meet today\u2019s educational approach in terms of cost and space. If a Multitask MPS (MTMPS) can be constructed in one compact platform with a SCADA system, it would overcome cost and space problems leading to more efficient training environment", + " Otherwise OFF label is displayed. Exit: Exit to the main page of the SCADA System. This section deals with the course evaluation which took place in Spring Semester 2008. The evaluation was performed at the end of the Introduction to Mechatronics Course MAK368. This process took 3 weeks time which corresponds to 6 h instruction. The aim of the evaluation was to compare the MTMPS and Conventional MPS and to emphasis the motivation effect on training. The MTMPS (Fig. 8) and a 5 stationed Conventional MPS (shown in Fig. 1, the left picture) were tested by 20 volunteer and enthusiastic students. All the students are supposed to be having almost the same initial background knowledge since they completed the same prerequisite courses successfully. Handouts related to both the MTMPS and Conventional MPS units were provided for students covering all the different scenarios described (emergency stops, fault findings, etc.). The students have been able to implement the programs and make the MTMPS and Conventional MPS work" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000188_arso.2010.5680046-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000188_arso.2010.5680046-Figure1-1.png", + "caption": "Fig. 1. The kinematic model of a car with n off-hooked trailers.", + "texts": [ + " We also propose a kinematic configuration of a trailer system to improve the backward motion control performance. Performances of the proposed control strategy and the kinematic configuration are verified by theoretical verifications and experimental results. This paper is organized as follows. Section II shows the kinematic model of a trailer system. The control strategy is proposed in section III. The kinematic configuration is proposed in Section IV. Section V presents the experimental results. Some concluding remarks are given in Section VI. P 978-1-4244-9123-0/10/$26.00 \u00a92010 IEEE Fig. 1 is shows the kinematic model of the off-hooked trailer system with a car-like mobile robot in [10]. It is shown that the trajectory tracking error can be minimized in following condition. A1) dcar = F = R The pose of the car-like mobile robot [x0,y0, 0] is defined at the center of the rear axle of the robot. The kinematic model shown in Fig.1 can be represented as eq. (1). The velocity of the tractor [v0 , 0] generate the velocity of the nth passive trailer [vn , n]. In this study, only one passive trailer is considered, since the kinematic condition between car-like mobile robot and a passive trailer is same regardless of the number of passive trailers, and the kinematic model of off-hooked trailers can be easily extended to n trailers. A trailer system which was consisted of one off-hooked trailer and car-like mobile robot in Fig. 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002109_1350650117723484-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002109_1350650117723484-Figure2-1.png", + "caption": "Figure 2 depicts an angular contact ball bearing with five displacements x, y, z, x, y of the inner ring with respect to the outer ring. Figure 3 shows the relationship between the ball and the inner ring groove curvature centers with the outer ring under the assumption that the groove curvature center of the outer ring is fixed. Based on these two figures, the relationship between the raceway curvature and ball centers at the jth ball position can be expressed as3", + "texts": [], + "surrounding_texts": [ + "Bearing characteristics, constant preload, bearing stiffness, thermal effects, lubricant film thickness Date received: 5 April 2017; accepted: 26 June 2017" + ] + }, + { + "image_filename": "designv11_33_0002941_icelmach.2018.8506824-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002941_icelmach.2018.8506824-Figure3-1.png", + "caption": "Fig. 3. Time-average iron loss distribution in the stator of the IPM, over one electrical period, as calculated by the transient FE model for (a) 1000 rpm and (b) 4800 rpm. Values are normalized by the maximum value calculated.", + "texts": [ + " For the homogenous case, the average loss distribution, or lumped loss, for each component of interest is obtained according to equation (2): \u00a0lumpedp P V (2) where P is the total loss in the component and V is its respective volume. For the analysis undertaken, a selection of components of interest were defined, where the different ways of transferring loss data were used, namely, the stator region consisting of iron laminations, the rotor region consisting of iron laminations and the rotor permanent magnets. The measured core loss data for the stator and rotor of the machine are shown in Fig. 2. Both stator and rotor are constructed using conventional iron lamination materials. Fig. 3 illustrates the per-unit time-averaged iron loss distribution in the stator iron, as calculated by the transient FE model, for 1000 rpm and 4800 rpm motor operation under Maximum Torque Per Ampere (MTPA) [3]. It can be observed that the region exhibiting the highest iron loss density is the stator teeth, in particular, close to the air-gap and near the slot ends in the stator back iron. This is attributed to the rectangular shape of the slots and coils, resulting in non-uniform tooth width and is quite typical" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003748_10402004.2019.1669755-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003748_10402004.2019.1669755-Figure5-1.png", + "caption": "Figure 5. Simplified diagram of a 3D model of the vertical hydrostatic slide system.", + "texts": [ + " Thus, pr is the pressure acting on the oil pad and guideway and can be computed by where L2 is the distance from the center of the pad to the Z axis and pr0 \u00bc psq0cR psh3 \u00f0cr 1\u00deq0cR : The distribution of the bearing pad pressure is supposed to be practical, as shown in Fig. 4. The maximum value of bearing pad pressure within the pocket is pr and the pressure will decrease linearly along the land from the pocket to the edge of the land until zero. Modeling of hydrostatic slide structure The geometries of pads and guideways, including the contact profiles, should be modeled as realistically as possible because they may influence both the deformation value and the performance of the hydrostatic slide. Figure 5 shows a detailed 3D geometric model of the a vertical hydrostatic slide system that was built on the basis of the pivotal components of the experimental apparatus. The model was initially created in a commercial Computer Aided Engineering (CAE) software, SolidWorks. Guideway II is fastened to the column and guideway I is fastened to guideway II. The hydrostatic block contains oil tubing in order to sustain the oil film for the hydrostatic bearing and ensure kinematic movement of the bearing. One of the guideways in the vertical slide is a column" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001164_sii.2011.6147547-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001164_sii.2011.6147547-Figure1-1.png", + "caption": "Fig. 1. The principle of the framework", + "texts": [ + " In this approach, the follower robot walks by following the walking pattern generated on-line by the interactive force occurred between the leader and the follower robots. However, this interactive force could induce unstable condition, and in the worst case, this leads to falling down of the robots. Therefore, the main goal of this research - 779 - SI International 2011 is to ease the interactive force while moving, and achieve cooperative transportation by the humanoid robots. The principle of the framework is shown in Fig. 1. The leader robot and follower robot are both stationary at first, and the hands position of follower robot in waist coordinates at this moment is d1. Interactive force between the robots arises as the leader robot start to move. Then the follower robot stretch its arm to ease the force. Distance d1 becomes d2, and the follower robot starts to generate walking pattern online to recover the hands position at the same time. Fig. 2 shows the concept of the whole system described above. To achieve this approach, online walking pattern generation and force control are fundamental technologies", + " Howerver, the spring constant of follower robot is set to zero. Therefore, follower robot recover its hand position to the reference position by walking. The walking command generation corresponds to the spring and damper between the object and follower\u2019s leg, and the impedance control of hands corresponds to the damper between the object and follower\u2019s body in Fig. 2 respectively. As the leader robot starts to move, the hands of follower robot start to stretch due to the impedance control, then d1 becomes d2 (Fig. 1). We use PD control to generate walking command that recovers d2 back to d1. The walking command is generated as: pacex = P(d2\u2212d1)+D(d\u03072\u2212 d\u03071) , (14) where pacex is a walking command that is the relative distance between landing placement and support foot, and P and D act as spring and damper in Fig. 2. Besides, only x direction (back and forward) is verified in this research. Reference ZMP is computed with the walking command, then CoM trajectory is calculated by appling preview control theory" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000365_iccaie.2010.5735132-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000365_iccaie.2010.5735132-Figure2-1.png", + "caption": "Fig. 2 Filtered layer gram process", + "texts": [ + " The theory is needed in order to study the inverse of the Radon transform for the sort of data that arise in medical imaging. The notion of weak derivative is also well adapted to measurements defined by averages. The object space will be filtered by the Radon transform, and then back-projected to find the Radon space. When the data collected, it will convert using the Fourier Slice Theorem to become the Fourier Space. Lastly the Fourier Space will inverse the Fourier Transform to get the object space back [6]. The process will vice versa as shown in the Figure 2. This paper will cover the development of image reconstruction algorithm using Filtered Back Projection technique through sinogram approach of a Computed Tomography System using MATLAB. As in figure 3, the phantom is use as an object to test or evaluate the image quality of a CT system. Line integral is the mathematical basis of tomography with non-diffracting sources that will show how one can go about recovering the image of the cross section of an object from the projection data. In ideal situations, line integral is a projection which produces a set of measurements data of the integrated values of some parameter of the object-integrations being along straight lines through the object" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000081_cphc.200900169-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000081_cphc.200900169-Figure1-1.png", + "caption": "Figure 1. Cross-section of the 2D spherical domain available to diffusion limited by a spherical-cap cluster formed at its north pole.", + "texts": [ + " KGaA, Weinheim 1593 Irrespective of the solution adopted in the following, namely, analytical and numerical or by random walk, we consider the 2D surface of a spherical body of radius R on which N0 compa- ratively small entities (molecules, ions, electrons, etc.) are evenly distributed initially, but which are prone to move randomly on it when submitted to a gradient of force created at one pole of the spherical surface. We assume here that such a driving force stems from the particles\u2019 ability to agglomerate with each other, thus forming an expanding cluster (Figure 1).[1, 3\u20135] To assure generality, neither the type of stimulation or the condition that provokes this aggregation nor its nature should be specified. Under such conditions the movement of the particles over the 2D curved surface is equivalent to diffusion with an apparent diffusion coefficient that we denote D. This diffusion coefficient may have a real physical origin (i.e. physical displacement) or represent the probability of exchange times the square of the distance between sites.[5\u20138] Similarly, we recently showed that if the 2D surface is partitioned due to local constraints, an equivalent diffusion law is obeyed, although each particle displacement may be energy-gated so that D obeys an Arrhenius-type law", + " This is further justified by the general observation that such structures tend to regroup to form densely packed arrays to avoid forms with long branches.[14] To avoid any occurrence of such fractal effects in the present simulations, after each time step the total cluster area was determined by accounting for the summed area of all aggregated objects and its shape reorganized symmetrically over the 2D surface at constant surface area. Since the cluster is represented by a spherical cap, one can easily evaluate the zenith angle of the cluster edge qclust\u00f0t\u00de assuming that its centre is located at the north pole of the sphere (see Figure 1) for any number of objects N\u00f0t\u00de that have agglomerated by time t and their known overall area [namely Saggr\u00f0t\u00de \u00bc N\u00f0t\u00deb2, where b2 is the area occupied by a single object in the cluster and may be selected to differ from pr2 0 ] . Then, at time t the area of the cluster is given by Equation (3): Saggr\u00f0t\u00de \u00bc N\u00f0t\u00deb2 \u00bc 2pR2 1 cos qclust\u00f0t\u00de\u00bd \u00f03\u00de Hence, Equation (4) follows: qclust\u00f0t\u00de \u00bc arccos 1 Saggr\u00f0t\u00de 2pR2 \u00bc arccos 1 21 N\u00f0t\u00de N0 \u00f04\u00de where 1 \u00bc N0b2=4pR2 is the fraction of the spherical surface area covered by the cluster at infinite time, that is, when the N0 objects initially present on the surface of the sphere are aggregated", + " Iterative repetition of the above sequence of procedures was achieved for all objects at each time step until the required duration of the experiment texp (or its dimensionless equivalent tmax \u00bc D texp=R2) was reached or up to when all the objects were aggregated. This afforded the number of aggregated objects N\u00f0t\u00de as a function of time, as is represented for example in Figure 3. As for the Brownian simulations, we consider a sphere of radius R whose surface initially contains freely diffusing objects present at a surface concentration of G0 \u00bc N0=4pR2. We suppose that at time t \u00bc 0 the objects start agglomerating to form a spherical-cap cluster centered at the north pole of the spherical surface (Figure 1) in which the concentration of the objects is constant and equal to G0=1. The cluster expansion with time may then be represented by the variations of the zenith angle qclust\u00f0t\u00de which defines the position of its edge according to Equation (4). The diffusion of objects on the spherical surface is described by the following partial differential equation [Eq. (6)]: @G @t \u00bc D R2 @2G @q2 \u00fe 1 tanq @G @q \u00f06\u00de where D is the diffusion coefficient describing the process irrespective of its exact physicochemical nature" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002009_icmsao.2017.7934926-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002009_icmsao.2017.7934926-Figure4-1.png", + "caption": "Fig. 4. Illustration of the three-dimensional finite element model.", + "texts": [ + " 2 2 2 2 2 2 3( / ) 3( / ) 3( / ) 3( / ) 3( / ) 3( / ) 6 3 ( , , ) 6 3 ( , , ) f r x af y b z d f f x a y b z dr r r f Q q x y z e e e a bd f Q q x y z e e e a bd , 2f rQ VI f f (2) where the parameters of a, b, c, f are obtained from the experimental welding pool shape, which is shown in Fig. 2(a). The simulating welding pool shape is presented as the red region of Fig. 2(b). The simulation result is suitable for the experimental result. III. FINITE ELEMENT MODELS The three-dimensional finite element model is composed of 9 clamping, a substrate and single-pass 6 layers component, which is shown in Fig. 4. The sub-clamping is a rigid body which represents the workbench and is set as bearing clamping. Clamping 1, 2, 3, 4 are the corner clamping. Clamping 5, 6 are the transversal clamping and clamping 7, 8 are the longitudinal clamping. Clamping 1-8 are configured as fixing clamping. Four clamping forms are used to investigate the influence caused by clamping. The simulation models of four clamping forms are shown in Fig. 5. Clamping form 1, which contains clamping 1- 4, is to get the effect of corner clamping" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002109_1350650117723484-Figure13-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002109_1350650117723484-Figure13-1.png", + "caption": "Figure 13. Comparison of contact angles and contact forces with and without thermal effect: (a) 10,000 r/min; (b) 15,000 r/min; (c) 18,000 r/min.", + "texts": [ + " The higher the preload, the smaller the absolute axial displacement is. These findings agree with the conclusions in Harris6 and Cao et al.26 Based on the temperature results in Than et al.,20 thermal expansions of bearing parts at different speeds are evaluated and listed in Table 1. After updating the bearing parameters D, ri, ro, dm\u00f0 \u00de, the bearing with thermal effects is analyzed (refer to algorithm in Figure 6). A comparison of contact angles and contact forces with or without the thermal effect at 10,000 r/min, 15,000 r/min, and 18,000 r/min is shown in Figure 13. The figure indicates that both contact angles and contact forces have a tiny change with thermal effects under a constant preload at varying speeds. Explanation for this point is that the two main factors, the preload and bearing parameters, which directly affect the contact angles and contact forces, are hardly altered under a thermal influence. Indeed, the preload is maintained as a constant through the spring preload, which can adjust its position by itself when affected by external forces or the increase in temperature in the spindle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002153_s12206-017-0704-1-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002153_s12206-017-0704-1-Figure1-1.png", + "caption": "Fig. 1. Test specimen: (a) Seal cross section; (b) model specimen.", + "texts": [ + " Such studies will provide an understanding that will assure their safe usage during application and prevent catastrophic events. Therefore, in this paper, a detailed analysis of the stress condition and fracture behavior of D-shaped seal (hereafter \u201cnew seal\u201d) was performed. The new seal was designed to have a ratio of rectangular height (H1) to circular height (H2) of 2.25. The optimal ratio of H1/H2 according to Finite element analysis (FEA) in Sec. 3 was used to fabricate the new seal ring with the cross section as shown in Fig. 1(a). It was subjected to 20 % compression and various pressures. The photoelastic model specimens shown in Fig. 1(b) were cast from a mixture araldite and hardener in the ratio (araldite: hardener) of 10:3. Details of the casting procedure for photoelastic models can be found in Ref. [8]. The new seal ring was installed into a specially designed loading device and stress freezing was done to lock stress fringes in the seal models. Stress freezing was done according to the thermal cycle in Ref. [8]. Small slices were cut from stress frozen seal models and polished to about 1.0 mm thickness. Isochromatic fringes were recorded from the polished samples using a digital camera mounted on transparent photoelastic experimental device Mod" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000188_arso.2010.5680046-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000188_arso.2010.5680046-Figure2-1.png", + "caption": "Fig. 2. The kinematic model of a car with a off-hooked trailer", + "texts": [ + " The kinematic model shown in Fig.1 can be represented as eq. (1). The velocity of the tractor [v0 , 0] generate the velocity of the nth passive trailer [vn , n]. In this study, only one passive trailer is considered, since the kinematic condition between car-like mobile robot and a passive trailer is same regardless of the number of passive trailers, and the kinematic model of off-hooked trailers can be easily extended to n trailers. A trailer system which was consisted of one off-hooked trailer and car-like mobile robot in Fig. 2. The velocity of the car-like mobile robot [v0, 0] generate the velocity of a passive trailer [v1 , 1] by eq. (2). The kinematic condition between the velocity of the car-like mobile robot [v0 , 0] and steering angle [ is shown in eq. (3). TABLE.1 PARAMETERS OF A CAR WITH A OFF-HOOKED TRAILER (2) (3) A. Inverse kinematics In proposed the control strategy, a driver can control the trailer system as a forward motion control instead of a backward motion control directly. Therefore a passive trailer is considered as tractor, and a driver should control the linear and angular velocity of a passive trailer [v1 , 1]", + " One off-hooked passive trailer is attached to rear of front bumper of the car-like mobile robot according to the experiments. Wheel diameter of car-like mobile robot and a passive trailer are 14.8cm, 11.0cm, respectively. The pose of a passive trailer is monitored by a commercially available pose sensor STARGAZER. The degree between car-like mobile robot and a passive trailer is offered by a potentiometer. The maximum steering angle of car-like mobile robot is 22.5 , and it is controlled by a servo motor. The length of wheelbase is 0.305m. The length of the [dcar,] [dtrailer] in Fig. 2 is 0.198m . Therefore condition A1) is satisfied. A commercially available web camera is setup on a passive trailer to offers the information of rear sight. If the conventional trailer system is considered to be compared with proposed strategy and kinematic configuration, web camera is installed at a side of the car-like mobile robot as a side view mirror. We divide the trailer system into C1)-C3) to perform the comparative experiment of backward motion control according to the control strategy and the kinematic configuration" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000715_cca.2010.5611245-Figure2.2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000715_cca.2010.5611245-Figure2.2-1.png", + "caption": "Fig. 2.2. A partially despun projectile coning with a constant angle of attack about its velocity vector. (a) The fuze tip traces a helical path along its trajectory. From head-on, the fuze appears to trace a circular path. (b) From the side view, this appears as an up-down sinusoidal oscillation of the body\u2019s angle with respect to its velocity.", + "texts": [ + " The canards are assumed to be actuated by electrical motors or voice-coils. Coning in projectiles can be described as the tip of the fuze tracing a helical path along the vehicle\u2019s trajectory. When viewed from head on, perpendicular to the velocity vector, the tip of the fuze appears to trace an ellipse throughout each coning cycle. This is a familiar phenomenon that is commonly exhibited when a football is thrown with spin. Even with the fuze of a projectile despun, the projectile may still cone when perturbed. See Fig. 2.2 for an illustration. Spinning the rear-end is desirable because conservation of angular momentum aids to stabilize the spinning vehicle in flight. Despinning the fuze simplifies attitude sensing and control action, because a fuze-fixed body frame does not roll with the rear-body. This frame does however continue to yaw and pitch as the body cones, exhibiting an angle of attack and sideslip. This phenomenon is known as precession or coning. As a partially despun projectile cones, the un-actuated canards move together with the non-rolling fuze-frame in a periodic motion. From a 3D perspective, it is clear that the body maintains an angle of attack that side-slips about the velocity vector, as in Fig. 2.2a. From a side view, this appears as an up-down sinusoidal change in the body\u2019s vertical angle with respect to the velocity vector, as in Fig. 2.2b. The following is a breakdown of three general coning-compensation scenarios. For simplicity it is assumed that the coning is symmetric, such that the coning path perpendicular to the velocity vector forms a perfect circle. Analysis and development will be carried out for a single canard on the fuze; similar analysis and controller design can be performed for multiple canards situated around the fuze circumference. The mathematical formulation in Section III takes these assumptions into account" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002746_s41314-018-0014-0-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002746_s41314-018-0014-0-Figure6-1.png", + "caption": "Fig. 6 SIR-X mesh representation", + "texts": [ + " 11 NT901 mesh representation The SIR-X projectile characteristics are given in Table 1. A series of RW impact tests at three different impact velocities covering the entire impact velocity range have been performed with the SIR-X projectile. The three impact velocities are 30 m/s, 58 m/s and 81 m/s corresponding to low, medium and high velocity range, respectively. The projectile is modelled with hexahedral solid elements. After a parametric mesh study, the body part is made of 2032 elements and the nose of 4336 elements. The mesh is shown in Fig. 6. The constant stress element formulation is used. The experimental results show that the strain rate does not have a great influence and can be neglected. The curve at 88 m/s is discarded as most of the projectiles were broken. The curves at 30 m/s are also discarded as no densification region is present. Therefore, the experimental stress-strain curve at 58 m/s is used as input in the model (Fig. 7). This curve is already smooth and monotonic. The unloading part is internally modelled in LS-DYNA by using the material parameters HU and SHAPE of the model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002793_ilt-07-2017-0208-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002793_ilt-07-2017-0208-Figure4-1.png", + "caption": "Figure 4 Approach to calculate the normal contact load Fc in the closed-form contact model", + "texts": [ + " The components along the x and y axes are: Frix \u00bc Fritcos b ri Friy \u00bc Fritsin b ri Mrix \u00bc Frizrocos b ri Mriy \u00bc Frizrosin b ri : (7) Considering the actual situation of a gas face seal, the tangential rub-impact loads can be ignored. Friz, Mrix and Mriy are retained serving as Fc, Mcx and Mcy in equation (5). By this, Mcx and Mcy are related to Fc and b ri. b ri can be determined by the relative position of the two mated rings. Thus, the remaining difficulty is how to briefly calculate the normal load Fc. Figure 4 provides the approach to calculate Fc in the closed-form contact model. As similar to the GW model, an equivalent rough surface (i.e. the rotor) contacting a smooth plane (i.e. the stator) is used to model the contact between two rough surfaces. There are three solid-line circles in the Figure: 1 The one with dot line relates to the mean surface height of the rotor zma. 2 The one with dot-dash line relates to the mean summit height of the rotor zms. 3 The remaining one relates to the smooth stator", + " Engineering surfaces usually exhibit a Gaussian or Gaussian-like characteristic. Figure 5 shows two Gaussian surfaces obtained from a spiral groove gas face seal with parameters being listed in Table I. The surface and summit parameters are listed in Table II. Note that in the surface analysis, a node will be regarded as a summit if it owns the maximal height within a 3 3 window (Brunetiere, 2010; Minet et al., 2010; Hu et al., 2016a, 2017a, 2017b, 2017c, 2018a, 2018b). As mentioned above, the truncation plane in Figure 4 should be defined as the sum of zms and Nts s. To obtain the correction factors tk and t s, Nt is set to six, yielding that the upper limit of integral 1 in equation (6) equals to zms 1 six s s. In the present study, E is set to 23.65 GPa. Figure 6 shows the relation between the contact pressure and the separation. The exponential curve well fits the numerical solution of equation (6), respecting the conclusion obtained by Greenwood and Williamson (1966) and Brunetiere (2010) that the load-separation relation for a Gaussian distribution approximates to that for an exponential distribution", + " However, it should be emphasized that even in the region with a large deviation, owing to the scaled effect of t ss s g rel tkkriexp c rog rel t ss s ffiffiffiffiffiffiffiffiffiffiffiffiffi 2ro t ss s g rel q in equation (18), the contact loads obtained by the two closed-form contact models are both tiny (i.e. smaller than 10 N). Furthermore, in Figure 7(a), the result obtained by the direct numerical contact model with grooves is smaller than the other three. For a better explanation, Figure 8 illustrates the coordinate schematic of the direct numerical contact model referring to Figure 4. The zone of the local Table I Parameters of a spiral groove gas face seal Parameter Value Outer radius ro (mm) 61.7 Inner radius ri (mm) 51.6 Balance radius rg (mm) 53.1 Spiral angle a (\u00b0) 15 Number of grooves N 12 Groove to land width ratio l 0.5 Groove to dam length ratio b 0.6 Groove depth d g (mm) 6 Figure 6 Relation between the contact pressure and the separation for two Gaussian surfaces D ow nl oa de d by U N IV E R SI T Y O F T O L E D O L IB R A R IE S A t 0 4: 51 1 1 A ug us t 2 01 8 (P T ) face contact always starts from a land in the direct numerical contact model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001841_0278364916679719-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001841_0278364916679719-Figure3-1.png", + "caption": "Fig. 3. The critical contact forces in (a) an SSI cell, and (b) an SSS cell of a three-contact stance with \u03bc = 0.4 at the contacts. The contacts are x1 =( 8, 0, 1), x2 =( 4, 4 \u221a 3, 1.2), x3 =( 0, 0, 1), with contact normals ni = ( sin \u03b1i sin \u03b2i, \u2212cos \u03b1i sin \u03b2i, cos \u03b2i) such that ( \u03b11, \u03b21) = ( 70\u25e6, \u221220\u25e6), ( \u03b12, \u03b22) = ( \u221215\u25e6, 34\u25e6), ( \u03b13, \u03b23) = ( \u221270\u25e6, \u221210\u25e6).", + "texts": [ + " In these equations \u03bd i( \u03c6i) = n\u0304i\u00d7\u03b7i( \u03c6i), where n\u0304i is a unit normal to the plane spanned by the contacts xi, xi+1, xi+2 and ( s\u0304i, t\u0304i) are orthonormal tangent vectors to that plane. The detailed proof of Proposition 3 appears in Appendix dimensionality, are given as follows. In the cell class SSI , the condition given by equation (12) implies that the forces f1 \u2208 S1 and f2 \u2208 S2 are directed such that the tangent planes 1( \u03c61) and 2( \u03c62) to the friction cones C1,C2 are both passing through the contact x3. This condition is illustrated in the three-contact stance of Figure 3(a). Criticality can be viewed by the fact that infinitesimal changes in f1 and f2 within the friction cone tangent planes, accompanied by infinitesimal changes in f3 \u2208 I3, will generate zero net torque about the intersection line of the two tangent planes 1( \u03c61), 2( \u03c62), which passes trough x3. Thus, the image of the critical contact forces under L lies in a five-dimensional linear subspace in B\u2019s wrench space of net wrenches. The condition given by equation (12) gives a finite number of solution pairs, ( \u03c6\u2217 1 , \u03c6\u2217 2 ), which are held fixed while the magnitudes \u03bb1, \u03bb2 > 0 and the three components of f3 vary freely within the SSI cell, thus spanning a five-dimensional sub-manifold of the cell. In the cell class SSS, the condition given by equation (13) implies that the three forces f1, f2, f3 are directed such that the intersection point z of the friction cone tangent planes 1( \u03c61), 2( \u03c62), 3( \u03c63) lies on the plane spanned by x1, x2 and x3. As an example, this condition is illustrated in the three-contact stance of Figure 3(b). Criticality of these contact forces can be explained as follows. Small changes in fi along its friction cone tangent plane can be partitioned into changes along the vector z\u2212xi, and changes along the orthogonal complement ( z\u2212xi) \u00d7\u03b7i, where \u03b7i is the tangent plane\u2019s normal. The changes in f1,f2,f3 along the z \u2212 xi components only span a two-dimensional subspace of net wrenches, V1, since they generate forces in the plane spanned by the contacts x1, x2, x3, while contributing zero torque about a line normal to that plane which passes through z" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002793_ilt-07-2017-0208-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002793_ilt-07-2017-0208-Figure3-1.png", + "caption": "Figure 3 Closed-form contact model used in the dynamics of a gas face seal", + "texts": [ + " This added assumption is logical because the closedform contact model is proposed for the opened operation rather than the startup or shutdown operation. Based on the assumptions, the axial rub-impact model of rotor dynamics Contact model for gas face seals Songtao Hu et al. Industrial Lubrication and Tribology D ow nl oa de d by U N IV E R SI T Y O F T O L E D O L IB R A R IE S A t 0 4: 51 1 1 A ug us t 2 01 8 (P T ) (Yuan et al., 2007, 2008) can be revised regarding a gas-faceseal system, as illustrated in Figure 3. Friz is the axial rubimpact load, and Frit is the tangential rub-impact load. The components along the x and y axes are: Frix \u00bc Fritcos b ri Friy \u00bc Fritsin b ri Mrix \u00bc Frizrocos b ri Mriy \u00bc Frizrosin b ri : (7) Considering the actual situation of a gas face seal, the tangential rub-impact loads can be ignored. Friz, Mrix and Mriy are retained serving as Fc, Mcx and Mcy in equation (5). By this, Mcx and Mcy are related to Fc and b ri. b ri can be determined by the relative position of the two mated rings" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002853_j.jmatprotec.2018.09.003-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002853_j.jmatprotec.2018.09.003-Figure3-1.png", + "caption": "Fig. 3. Experimental setup (a) side view, and (b) top view.", + "texts": [ + " In this paper, stability of LBW of with an off-axis wire feed in leading direction (allowing seam tracking close to the process) is investigated. Morphology of the welds and events forming irregularities are analysed using surface images, HSI and streak images. 2mm thick austenitic stainless steel sheets (304 L) were welded together in a close-to-zero-gap butt joint configuration, with added filler wire (1mm in diameter) to bridge the gap and ensure good mechanical properties. The LBW process had an off-axis wire feed in leading direction, Fig. 3. The chemical composition of both the base material and filler wire are presented in Table 1. In all experiments, the setup shown in Fig. 3 remained the same. An IPG fibre laser with maximum power 15 kW, wavelength 1070 \u00b1 5 nm, 400 \u03bcm output fibre diameter and beam parameter product 14.6mm.mrad was used in continuous wave (CW) mode. The optics used was Precitec YW50, with 150mm collimator lens and a 250mm focussing lens creating to create a focal spot diameter of 0.67mm. The focus was positioned 8mm above the surface of the plates, producing a theoretical Gaussian spot size 1mm width at the surface, so that the laser beam irradiated the gap and the full width of the wire", + " The HSI was carried out during welding at 10 000 fps using a Photron Fastcam mini UX100 camera with a Nikon 200mm f/4 macro lens and an additional 67 mm macro magnifying lens. Laser illumination was provided by a Cavilux HF pulsed diode laser with maximum effect of 500W and wavelength 808 nm manufactured by Cavitar Ltd. A narrow band pass filter matching the wavelength of the illumination was used to block process light. The camera and illumination were angled at 30\u00b0 to the horizontal plane, as shown in Fig. 3a, yielding results similar to Fig. 2a. Stability evaluations were made by tracing the edges (from above) on surface photographs for both weld cap and root of all welds, on the inside of any undercuts. The trace lines were then analysed by a code made in MATLAB. Weld width throughout the weld length and its variations at small intervals could then efficiently be derived. Three cross sections were made 4 cm apart on each weld, not necessarily at specific topographic locations. They were photographed under a stereo microscope for tracing the weld cap geometry" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002897_irsec.2017.8477423-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002897_irsec.2017.8477423-Figure7-1.png", + "caption": "Figure 7. (a) Mesh generation ; (b) flux paths; (c) Flux density |B| of SRM", + "texts": [ + " The M36 steel with nonlinear B-H characteristic were chosen for the stator and rotor, the boundary conditions is the type \u2018\u2019PrescribdA\u2019\u2019and the phase circuit is the copper, with 260 number of turns in windings. Once the preprocessor phase is finished, begins the processing phase. The FEMM generates mesh for the geometry, a triangulation is made, dividing the problem into triangles to solve the problem by finite elements analysis. Then, simulations can be made and results can be deduced. The machine divided into triangles, the flux density |B| and flux paths of the SRM are shown in Fig. 7 when only phase A is excited. The characteristics of inductance and torque are essential for creating the model of the SRM. For this, many simulations have to be launched for different values of the phase current and different positions of the rotor, so the using of opensource scripting language called LUA in the development of this study was important to automate this process. The script calculates the flux using \u201cmo_getcircuitproperties\u201d function and torque using \u201cmo_ blockintegral (22)\u201d function, for 20 current values from 1,5 A to 20 A, and for 31 position values from 0\u00ba to 30\u00ba using two interactive loops" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001580_iros.2011.6094712-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001580_iros.2011.6094712-Figure3-1.png", + "caption": "Fig. 3. Rigid-link model described in section IV.", + "texts": [ + ", when Fl(Xi)Fr(Xi) = 0 is not satisfied for all i, the following relation holds: J\u2032 = J\u2212 2 N N\u2212n \u2211 i=1 Fl(Xi)Fr(Xi) < J. (21) Thus, the force distribution given by (19) and (20) does not minimize the cost function J, which implies that the cocontraction of antagonistic muscles does not contribute to efficient locomotion. In this section, we propose a decentralized control scheme for a snake-like robot with multi-articular muscles in order to realize efficient locomotion; this scheme is based on the theoretical result derived in the previous section. The adopted model is shown in Fig. 3. The backbone consists of N rigid links concatenated one-dimensionally, where the length of each link is \u2206s and the joint angle between the i\u22121th link and the ith link is denoted by \u03c6i. A rotational spring is embedded in each joint. A rigid rectilinear link of length 2r perpendicularly bisects each backbone link. The tips of these links on the ipsilateral side are connected by n-articular muscles. The forces generated by the muscles that connect the ith and i+nth link on the right and left side are denoted by F\u0302r,i and F\u0302l,i, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000188_arso.2010.5680046-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000188_arso.2010.5680046-Figure4-1.png", + "caption": "Fig. 4 Presentation of feasible velocity region of the passive trailer attached to the rear side of the car", + "texts": [ + " Desired velocity of the car-like mobile robot [v0 0] is converted to control input of car-like mobile robot [v0 , ]. STEP3 A electronic control input device enter the control input [v0, ] into the car-like mobile robot, and the velocity of the car-like mobile robot [v0, 0] generate desired velocity of a trailer [v1, 1]. A driver secures the motion of a passive trailer by rear view display, and iterates STEP1-STEP3 to control the backward motion of a trailer system. Table.2 Notations for deriving feasible velocity region In Fig. 4, the feasible-velocity region at the revolute joint is presented when a passive trailer is attached rear bumper of the car-like mobile robot. The notation of Fig. 4 is presented in Table. 2. in Fig (5), denotes the admissible velocity direction at the revolute joint. of a trailer system has the kinematic configuration in Fig. 4 is presented by Eq. (5). Therefore the feasible velocity region is proportional to maximum steerable angle of car-like mobile robot. If the angle between the car-like mobile robot and a passive trailer increase more than , a trailer system control is failed, since the control input velocity decreasing does not exist. (5) max is the maximum angle between the car-like mobile robot and a passive trailer. Therefore In backward motion, a trailer system has the minimum turning radius when a trailer system maintaining max" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000908_0020-7403(65)90021-4-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000908_0020-7403(65)90021-4-Figure7-1.png", + "caption": "Fig. 7. The rigid ver t ical member and the sloping elastic member are both of length L.", + "texts": [ + " The theory of elast ica in the non-l inear behaviour of plane f rameworks 675 E x a m p l e (2) Eigenvalue buckl ing of por ta l - type f rameworks is s tudied on the model f rame in Fzo. 7. Example 2 ; elastic buckl ing of a por ta l - type frame, '\u00a3he elastic member is inclined to the horizontal at angle Y and its flexural r igidi ty is E l . The critical load in this case is g iven by W e L ~ w \u00b0 . . . . 3(1 + s iny) 2 (23) E I and the pa th by AWL 2 A w = E 1 ~ c o s y ( l + s i n y ) v n + . . . (24) ] 'he slope takes the largest value when $ = 30 \u00b0. Pa ths for different values of 7 are shown plot ted in Fig. 7. The analysis is similar to tha t of the last example. A X I A L L Y H Y P E R S T A T I C R I G I D L Y J O I N T E D B R A C E D F R A M E W O R K S I n cer ta in cases buckling of such frameworks m a y be p reven ted al together . An example is g iven in Fig. 8. Assuming the members are axial ly rigid and t h a t flexural buckl ing occurs, jo in t A mus t lie in the area bounded by the circles a and b. This, however , is impossible since it mus t also lie wi th in the are~ bounded by the temgential circles a ' and b'" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003138_cdc.2018.8618889-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003138_cdc.2018.8618889-Figure3-1.png", + "caption": "Fig. 3: Solution of the nonlinear diffusion model at different times t", + "texts": [ + " We also validate the coverage approach based on nonlinear diffusion by numerically solving the corresponding PDE (3). We use the finite volume method [16] to solve the PDE. The manifold M defined in this example is the 2- dimensional torus. Solving the PDE on the torus is equivalent to solving it on a rectangular domain with periodic boundary conditions. We set \u03b1 = 2 in the PDE. The target swarm density is shown in Fig. 2. Snapshots of the swarm density under diffusion at different times (i.e., the solution of the PDE) are shown in Fig. 3. As predicted by the previously stated asymptotic stability results, the solution is close to the target density for large enough times. In this paper, we presented two generalizations of a diffusion-based multi-agent coverage strategy. First, we extended the theory to general compact Riemannian manifolds without boundary. Then, we considered coverage based on nonlinear diffusion models. Numerical simulations verified the validity of the approaches. In future work, we will investigate the extension of this approach to manifolds with boundary" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002414_s11668-018-0398-4-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002414_s11668-018-0398-4-Figure1-1.png", + "caption": "Fig. 1 Bearing geometry describing rotation mechanism", + "texts": [ + " As the fault frequency model for the bearing surface roughness faults has not been derived yet, thus, these noninvasive techniques could not be used for the diagnosis of bearing roughness faults [25, 26]. The bearing consists of five elements known as outer race, inner race, ball, cage and the rotor on which bearing has been installed. The rotor motion (fr), ball motion (fb), motion of the inner race (fi), motion of the outer race (fo) and cage motion (fc) could be used to describe the dynamics of the bearing. The frequencies of the various elements of the bearing are shown in Fig. 1. The balls and cage of the bearing rotate with the rotation of the inner race. The rotational frequency of the bearing cage could be calculated using Eqs 1\u20133 [31]. Vc \u00bc Vo \u00fe Vi 2 \u00f0Eq 1\u00de fc \u00bc Vc rc \u00bc Vo \u00fe Vi Dc \u00bc foro \u00fe firi Dc \u00f0Eq 2\u00de fc \u00bc 1 Dc fo Dc \u00fe Db cos a 2 \u00fe fi Dc Db cos a 2 \u00f0Eq 3\u00de where Vo is the bearing outer race velocity; Vi is the bearing inner race velocity; Db is the bearing ball diameter (6 mm); Dc is the pitch diameter of bearing (25 mm); a is the ball contact angle; fi is the fault frequency of the inner race; f0 is the fault frequency of the outer race; ri is the bearing inner race radius; and ro is the bearing outer race radius" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001257_cm.20459-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001257_cm.20459-Figure2-1.png", + "caption": "Fig. 2. Schematic model of OAD and estimation of the HC stalk tilting angle. (a) Schematic model of OAD along the microtubules in the axonemes. The AAA-rings of a-, b-, and c-HC are colored in orange, yellow, and blue, respectively, while the stalks are in green. The stalk length is indicated by l. The circumferential and radial directions of flagella are indicated by the x- and y-axis, respectively, while the longitudinal direction of microtubule from the minus end to the plus end is the z-axis. The axis of y0 indicates the viewing direction suggested by Ueno et al. [2008] and is tilted by 20 against the radial direction (y). The protofilaments indexes correspond to those shown by Sui and Downing [2006]. The protofilaments indicated in white index belong to B-tubules, while those in black belong to A-tubules. (b) The distance between the HC ring and the adjacent B-tubule and the stalk tilting angle to be measured were indicated by lp and u, respectively. (c) Schematic diagrams of OAD and MT doublets according to Ueno et al. [2008]. On the assumption that the stalk length is l \u00bc 15 nm, the stalk angle (y) with respect to MTs was estimated from lp and u in (b) (see Table I for values).", + "texts": [ + " After all combinations between mutants and wild-types were tested, the best combinations between mutant and wild-type were selected so as to show little difference in the doublet regions and the other backgrounds except for the OAD regions. Images of wild-type were sharpened by unsharp mask (radius: 70 pixels, contrast weight: 500%), so that the edge of rings was enhanced while stalk regions were eliminated as faint and fine textures (Photoshop ver. 7). The contour lines of the rings were then traced at a middle level of thresholds. The tilting angles of density differences extending from a ring of OAD to the adjacent B-tubule (u in Fig. 2) and the distance between a ring of OAD and the adjacent B-tubule (lp in Fig. 2) were measured as follows. By assuming the stalks were l \u00bc 15 nm in length, the tilting angle against the axis parallel to the B-tubule was calculated (y in Fig. 2). In addition, the binding position of the MT binding domain (MTBD) on the adjacent B-tubule was identified as the position where the wild-type protofilament showed significantly larger values than that of the mutants and/or the projections considered to be stalks could be linearly extrapolated. Thin cross sections of wild-type C. reinhardtii and mutant (oda11, oda4-s7, and oda2-t) axonemes were prepared and examined by electron microscopy (Figs. 1a\u20131d). It is known that only eight doublet MTs in wild-type and oda11 flagella have OADs among the total of nine doublet MTs present, while doublet-1 does not contain OAD [Hoops and Witman, 1983; Sakakibara et al", + "9 to 90%, protrusions could be observed extending from each AAA-ring of OAD to the adjacent B-tubule (as indicated by the arrowheads in Fig. 4) and portions of the protrusion tips were also detected on the adjacent B-tubules. These regions likely originated from the stalk and MTBDs, respectively, as the flagella were in the rigor state under the conditions used for imaging, which would allow the MTBDs to strongly bind the B-tubules through the stalk region. Each protrusion was 12 nm in length and the determined stalk angles against the long axis of the MT doublet cross sections (u in Fig. 2b) were 40 (Table I). Sequence analyses have shown that the stalks of all three OAD HCs are almost identical in length (Yagi T., personal communication). When the stalk n 470 Takazaki et al. CYTOSKELETON length was assumed to be 15 nm [Goodenough and Heuser, 1984], the stalk angles against the axis along the MTs (y in Fig. 2c) were estimated to be 47\u201357 (Table I), which is quite similar to the angles reported by Ueno et al. [2008]. From the imaging analyses in the stalk region, the tubulin protofilaments to which the MTBDs of each HC bind could be assigned (Table I and Supporting Information Fig. S4). The stalks of a-HC and c-HC were most often connected to B6 and B3, respectively, of the adjacent B-tubule, while the stalk of b-HC was most frequently associated with B4. Although B6 represented another peak in the binding site distribution of b-HC (Supporting Information Fig", + " In addition, as the axonemes were 200 nm in diameter (Supporting Information Table S1), which is consistent to the diameter reported by Bui et al. [2009], this indicates they were largely unaffected using plastic-embedded thin sections for electron microscopic observations. Thus, the calculated protrusion length is plausible, which indicates this preparation and visualization method is suitable for examining the detailed structures of OAD. In order to estimate the stalk angle, it was assumed that the stalk and MTBD are a total of 15 nm in length. According to the notation of Fig. 2, it was estimated that stalks are tilted at an angle of 50 against MTs (Table I). These results are close to the angle of 54.0 6 8.7 reported for sea urchin OAD in the rigor state [Ueno et al., 2008] and suggest that the protrusions extending from OAD to adjacent B-tubules are indeed stalks and the determined binding positions are plausible. Thus, the present study also supports that the stalk of OAD is tilted against MTs under rigor conditions in the axonemes as well as in the reconstructed dyneinMT complexes of plastic-embedded specimens described by Ueno et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000345_6.2009-5890-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000345_6.2009-5890-Figure4-1.png", + "caption": "Figure 4. Collision of NextGen wing structure with high sweep and high chi.", + "texts": [ + " Exceeding these ranges can result in mechanical damage as the drive system would collide with the end-stops, most likely resulting in motor burnout and damage to the structural components. The four designated wing configurations chosen for specific flight conditions are Loiter, Cruise, High Lift, and High-Speed Maneuver. Two of these configurations (Loiter and Cruise) are on the boundary of the independent mechanical constraints and form two corners of the \u201cmorphable\u201d region. In addition to the independent mechanical constraints, the maximum chi and sweep angles cannot be achieved simultaneously, as wing damage may result when wing components collide with one another (Figure 4) or with other structural components within the aircraft body. The constraint boundary created by this limitation can be estimated based on the geometry of the wing design. The resulting constrained morphing region is depicted in Figure 5. During implementation and hardware-in-the-loop testing, a more conservative morphable region was used for the purposes of risk reduction. This region is confined by the four designated wing configurations (chi, sweep): [(30, 30) (35, 75) (50, 65) (60, 30)]. A more convenient space, however, is the 2 x 2 square contained by: [(\u22121, 1) (1, 1) (1,\u22121) (\u22121,\u22121)]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001216_icc.2012.6363714-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001216_icc.2012.6363714-Figure2-1.png", + "caption": "Fig. 2. Combining cooperative transmission with multi-hop routing.", + "texts": [ + " It is important to make the distinction between being a cooperative AF relay and serving as a relay along a multi-hop route through the WSN. It is assumed that the sensor network in this study employs an ad-hoc routing algorithm where nodes can serve as relays along a multi-hop routing path. However, nodes can also be selected to serve as AF relays by the medium access control layer of the protocol stack. Thus, the existence of an AF relay is completely transparent to the routing layer. The routing layer simply sees increased network connectivity that is made possible by AF transmission. Fig. 2 shows an example of a multi-hop network with cooperative transmission. The figure shows a multi-hop route starting with node 1 and ending at node 5. Note that cooperative transmission is utilized on the link between nodes 1 and 2 and the link between nodes 4 and 5. The routing layer would be completely unaware that nodes 6 and 7 had participated in relaying the transmitted data. It is assumed that each node must be able to transmit data to any other node in the network. This is a high reliability scenario where nodes are inclined to share data with multiple destinations rather than trying to consolidate all data with a single destination" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000227_sav-2010-0518-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000227_sav-2010-0518-Figure1-1.png", + "caption": "Fig. 1. Typical example of a tooth bending fatigue crack.", + "texts": [ + " A gear set may exhibit a variety of failure modes affecting either all gear tooth surfaces (e.g. scuffing, pitting, plastic flow, abrasive wear) or a single tooth (or a few teeth) on a gear (e.g. tooth fracture due mainly to bending fatigue, gear rim failure) [2]. Of these faults, tooth fracture, which is one of the most dangerous failures, generally initiates either at the root or on the surface of a tooth [3], spreads rapidly into the gear body, and finally results in either complete or partial tooth breakage as shown in Fig. 1. The surprising feature is that tooth breakage can occur and may not be noticed until a routine strip down. In the end, the major hazard may become inevitable if the broken tooth attempts to go through the mesh and may jam the drive. Over the last two decades, vibration analysis is extensively used as the basis for fault detection in gearboxes and condition indicating information is obtained by employing different techniques including: conventional signature analysis in time and frequency domains [4\u20137], signal demodulation [8\u201310], and cepstrum [11,12] techniques" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001039_09544119jeim730-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001039_09544119jeim730-Figure4-1.png", + "caption": "Fig. 4 Working space of the bidirectional steerable catheter", + "texts": [ + " Here, mi j and eij are used to denote the position and direction vectors of the sensor, where i (i5 1, 2) refers to sensor 1 or sensor 2, and j (j5 1,\u2026,n) refers to the sampling time. The bidirectional steerable catheter is capable of bending in a plane parallel to the plane formed by two tracking sensors. If these sensors are well integrated into the centre of the catheter, these two planes overlap. The bending angle is denoted by h. Both theoretical and experimental results showed that the deflection curve of the steerable catheter can be treated as a circular arc [16]. Figure 4 shows the workspace of the catheter. In the local coordinate system SOxy built on the tip of sensor 2, it can be expressed as xt~L 1{cos h h zls sin h yt~L 1{cos h h zls sin h \u00f01\u00de where h [ \u00bd0 p , L is the length of the bending segment, and ls is the length of the non-bending segment of the distal end. The movement of the tip on advancing is determined by the contact between catheter and vessels. However, the catheter is flexible and the contact site with the vessels is uncertain. Thus, there is no definite relationship between the advancement of Proc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002158_gt2017-64151-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002158_gt2017-64151-Figure2-1.png", + "caption": "Figure 2. Forces acting on ball bearing elements.", + "texts": [ + " The studied deep-groove ball bearing had a bore size of 70 mm, and it had a PTFE cage. The inner race, outer race, and balls were assumed to be made of SUJ2 steel and coated with a soft metal (i.e., silver). It was assumed that the shapes of the ball bearings were ideal spheres, and whirling of the inner race caused by the rotor whirling motion is ignored to reduce computational time. The outer guidance clearances of the studied cages were 1.14, 1.04, 0.94, 0.84, and 0.74 mm and the ball\u2013pocket clearances were 0.62, 0.92, 1.22, 1.52, and 1.82 mm. Figure 2 shows the principal forces acting on the bearing balls and cage. The forces can be categorized largely by the mechanical interaction between the ball bearing elements and the drag force by the viscosity effect of lubrication [15]. In some studies, the hydraulic force generated between the cage and race land is considered using short bearing theory (laminar flow). However, after passing through the seal in the liquid turbo-pump, cryogenic fluid passes quickly between the cage and the race land in the axial direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001381_kem.490.237-Figure9-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001381_kem.490.237-Figure9-1.png", + "caption": "Fig. 9. Flank tooth surfaces of the pinion", + "texts": [], + "surrounding_texts": [ + "Methods of numerical modeling of cutting, which was presented above, is of considerable practical importance. We can speed up the process thanks to its introduction into the production of the new transmission, obtaining at the design stage of technology, appropriate landscaping the lateral surfaces of the teeth, which will trace the desired mesh. Through numerical simulation, we can quickly determine appropriate values for machine settings and check the effect of treatment, setting the received transmission units in constructional gear pair. Financial support of Structural Funds in the Operational Programme - Innovative Economy (IE OP) financed from the European Regional Development Fund - Project No POIG.0101.02-00-015/08 is gratefully acknowledged." + ] + }, + { + "image_filename": "designv11_33_0002106_msec2017-2792-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002106_msec2017-2792-Figure5-1.png", + "caption": "FIGURE 5 GENERAL HYPER-ELASTIC MATERIALS ENGINEERING/NOMINAL BEHAVIOR AND WEIGHT FUNCTION REPRESENTATION.", + "texts": [ + " (12), the WFB model assumes that the strain energy is function only in the first principal stretch value (\ud835\udf061). Equation (12) can be calculated numerically or be solved directly after choosing a proper formula of the weight function F(\u03bb1). The author avoided giving a closed form solution for Eq. (12) as it depends on the weight function F(\u03bb) which is, according to the proposed model, a user pre-defined function and it can take any mathematical form as long as it satisfies the weight function constrains and conditions; (Figure 5-b). \ud835\udc4a = \ud835\udc4a(\ud835\udf061) = \u222b {\ud835\udc39(\ud835\udf061) \ud835\udc34(\ud835\udf061\ud835\udc52\u2212\ud835\udc35 \ud835\udc3c1) + \ud835\udc36(\ud835\udf061\ud835\udc3c1 \u2212\ud835\udc37)} (\ud835\udf061 \u2212 1 \ud835\udf061 ) \ud835\udc51\ud835\udf061 \ud835\udc3f\ud835\udc53 1 (12) Where (Lf) is the fracture stretch value. By substituting Eq. (12) into Eq. (11), the first Cauchy principal stress equation for uni-axial tension (\ud835\udf061 = \ud835\udf06), can now be represented by the following formula, \ud835\udf0e1 = {\ud835\udc39(\ud835\udf06) \ud835\udc34(\ud835\udf06\ud835\udc52\u2212\ud835\udc35 \ud835\udc3c1) + \ud835\udc36(\ud835\udf06\ud835\udc3c1 \u2212\ud835\udc37)} (\ud835\udf062 \u2212 1 \ud835\udf06 ) (13) where A, B, C and D are the proposed WFB model parameters. Upon several experimental observations, the authors found that the accuracy of the proposed model is related to the stressstretch curve slope transition point (TP). TP represents the stretch value of the minimum slope on the stress-stretch curve for hyper-elastic material; see Fig.5. Physically, TP is the stretch value after which the hyper-elastic material stops resisting the applied load. From Eq. (13), the exponential term (\ud835\udf06\ud835\udc52\u2212\ud835\udc35 \ud835\udc3c1) is multiplied by a weight function to restrict the effect of the exponential term after a specific stretch value. By observing the least square error (LSE) compared to experimental data, it was found to be minimum by forcing the weight function F(\u03bb) to be zero for \u03bb greater than a specific ratio from the TP value (H*TP); see Figure 5-b and Figure 6, where H is a transition point factor and it is a material constant that can vary from material to another. H has been chosen to be 1.409 for the tested natural rubber materials; see Fig.2, as it gives the minimum error compared to the tested data; see Eq. (28) for error calculation. Figure 7 represents a sample of a weight function for the tested natural rubber specimen. the authors used Eq. (14) to fit the weight function curve, \ud835\udc39(\ud835\udf06) = \ud835\udc39\ud835\udc431(\ud835\udf062 + \ud835\udc39\ud835\udc432)\u2212\ud835\udc39\ud835\udc433 (14) where the weight function parameters FP1, FP2 and FP3 can be evaluated using optimization tool from Matlab (Lsqcurvefit) or any other linear or non-linear optimization fitting tool, and they found to be 2.378E8, 15.5128 and 7.0574 respectively. Equation (14) is not a must for F(\u03bb) fitting, the user is free to use any smooth piecewise function as long as it fits the main weight function assumptions and shape; see Fig. 5-b. 4 Copyright \u00a9 2017 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 08/25/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use The WFB Model for Bi-axial Loading: For Bi-axial stress state (\ud835\udf0e1 = \ud835\udf0e2 = \ud835\udf0e and \ud835\udf0e3 = 0); see Fig. 3, the principal stretches, Cauchy\u2019s principal stress (\ud835\udf0e) and the arbitrary pressure (p) can be represented as follows, \ud835\udf061 = \ud835\udf062 = \ud835\udf06 \ud835\udc4e\ud835\udc5b\ud835\udc51 \ud835\udf063 = \ud835\udf06\u22122 (15) p = 2 \ud835\udf064 \ud835\udf15\ud835\udc4a \ud835\udf15\ud835\udc3c1 + 2 \ud835\udf062 \ud835\udf15\ud835\udc4a \ud835\udf15\ud835\udc3c2 (16) \ud835\udf0e1 = \ud835\udf0e2 = [2 \ud835\udf15\ud835\udc4a \ud835\udf15\ud835\udc3c1 + \ud835\udf062 \ud835\udf15\ud835\udc4a \ud835\udf15\ud835\udc3c2 ] (\ud835\udf062 \u2212 1 \ud835\udf064) (17) By using the same equation (Eq", + " Numerically, the Ogden model should be more accurate as it has more flexible formulation compared to the Yeoh and the WFB models, however it takes a very long time to extract its parameters. By comparing the processing time, the proposed WFB model reduced the processing time compared to the Ogden models by 99.82% and 99.96% for N=3 and N=4 respectively. According to the WFB methodology, the weight function is related to the TP value for reaching the minimum LSE compared to the test data. By observing the LSE variation with (H); see Fig. 5, the LSE was found to me minimized when H = 1.964 where TP is 2.4 for Treloar\u2019s vulcanized rubber material; see Fig. 8. TABLE 1 HYPER-ELASTIC MODELS LSE FOR UNI-AXIAL LOADING Model LSE Processing time (sec) Yeoh 0.0173 0.0156 Ogden (N=3) 0.0100 16.9219 Ogden (N=4) 0.0105 76.3906 WFB 0.0104 0.0313 TABLE 2 HYPERELASTIC MODELS PARAMETERS FOR FITTING TRELOAR\u2019S UNI-AXIAL TEST DATA USING LSQCURVFITT OPTIMIZATION TOOL IN MATLAB. Yeoh Model: \ud835\udf0e = [2\ud835\udc36\ud835\udc4c1 + 4\ud835\udc36\ud835\udc4c2(\ud835\udc3c1 \u2212 3) + 6\ud835\udc36\ud835\udc4c3(\ud835\udc3c1 \u2212 3)2](\ud835\udf06 \u2212 \ud835\udf06\u22122) CY1 = 0.1784 MPa CY2 = -2", + "066 compared to 0.0036, 0.0040 and 0.0263 for Yeoh and Ogden (N=3) and (N=4) respectively; see Table 4. However, the Yeoh model is the least in processing time with 0.01563 sec compared to 0.0.03125 sec for all the WFB and Ogden (N=3) and (N=4) respectively. Although the WFB is not the optimum tool for fitting Treloar\u2019s Bi-axial data, it showed a better fitting that reduced the LSE by 74.9% compared to the Ogden model for (N=4) for the same processing time. By observing the LSE variation with (H) in Fig. 5, it was found to me minimized when H equals to 2.0 where TP is 1.943 for Treloar\u2019s vulcanized rubber material; see Fig. 10. TABLE 3 HYPER-ELASTIC MODELS PARAMETERS FOR FITTING TRELOAR\u2019S BI-AXIAL DATA USING LSQCURVFITT OPTIMIZATION TOOL IN MATLAB. Yeoh Model: \ud835\udf0e = [2\ud835\udc36\ud835\udc4c1 + 4\ud835\udc36\ud835\udc4c2(\ud835\udc3c1 \u2212 3) + 6\ud835\udc36\ud835\udc4c3(\ud835\udc3c1 \u2212 3)2] (\ud835\udf06 \u2212 \ud835\udf06\u22125) CY1 = 0.1857 MPa CY2 = -16.0075E-3 MPa CY3 = 3.2246E-3 MPa Ogden Model (N=3): \ud835\udf0e = \u2211 \ud835\udf07 \ud835\udc41 [\ud835\udf06\ud835\udefc\ud835\udc41\u22121 \u2212 \ud835\udf06\u22122 \ud835\udefc\ud835\udc41\u22121] \u03bc1 = 387.7853 MPa \u03bc2 = 6.2636E-2 MPa \u03bc3 = -1.0625E-6 MPa \u03b11 = 1.9945E-3 \u03b12 = 2.4213 \u03b13 = -9", + "5774 MPa \u03bc3 = -8.4648E-11 MPa \u03bc4 = -1086.4000 MPa \u03b11 = 47.8172E-3 \u03b12 = 1.5143 \u03b13 = -14.1835 \u03b14 = 0.1668 WFB Model: \ud835\udf0e = {\ud835\udc39(\ud835\udf06) \ud835\udc34(\ud835\udf06\ud835\udc52\u2212\ud835\udc35 \ud835\udc3c1) + \ud835\udc36(\ud835\udf06\ud835\udc3c1 \u2212\ud835\udc37)} (\ud835\udf062 \u2212 \ud835\udf06\u22122) A = 1.2656 MPa B = 0.5894 C = 0.2512 MPa D = 0.3884 Although the accuracy of the Ogden model should increase by increasing the number intervals, this is not the case in the pure shear results, as the model reached the maximum number of iterations of (2.0E6) for N=4 without reaching the optimum solution. By observing the LSE variation with (H); see Fig. 5, the LSE was found to me minimalized when H = 2.0 where TP is 2.402 for Treloar\u2019s vulcanized rubber material; see Fig. 11. TABLE 6 HYPER-ELASTIC MODELS LSE FOR TRELOAR\u2019S PURE SHEAR DATA. Model LSE Processing time (sec) Yeoh 0.0027 0.01562 Ogden (N=3) 0.0256 429.87500 Ogden (N=4) 0.0017 2.28125 WFB 0.0071 0.04688 8 Copyright \u00a9 2017 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 08/25/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use The WFB model has been applied to the tested natural rubber specimen" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002015_tmech.2017.2713397-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002015_tmech.2017.2713397-Figure5-1.png", + "caption": "Fig. 5 Experimental testbed for duplex cutting of thin-wall WP", + "texts": [ + " The twin-SB behaves as a (Z, , ) fine-motion 3 4 5e e eq q q manipulator [25, 26]. III. RESULTS AND DISCUSSIONS In this section, numerical and experimental results are presented to illustrate and validate the F/T model for multi-DOF load compensation, and evaluate its effectiveness for minimizing MBR in a machining application. The investigation was conducted numerically on a Direct-Drive Dual-Disc (D4) spindle motor as illustrated in the CAD model in Fig. 4, and experimentally on a testbed (Fig. 5) for duplex lathe cutting of a thin-wall disk-like workpiece (WP). The D4-motor has a Twin-SB configuration (Fig. 4), each set consists of 48 stator-EMs and 64 rotor-PMs (satisfying the assumption that both NE and NP are multiple of 4) for six-DOF actuation discussed in Section II.C. The EMs are air-cored rectangular coils wrapped with 0.5mm-diameter copper wires. The (30\u00d750\u00d710 mm) coil has a square cross-sectional area of 100mm2. Each EM is individually driven by a linear current amplifier (capable of delivering a maximum current of 4A) with a closed-loop current-control implemented on an onboard circuit (Fig. 5c) to assure precise 1083-4435 (c) 2016 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. > REPLACE THIS LINE WITH YOUR PAPER IDENTIFICATION NUMBER (DOUBLE-CLICK HERE TO EDIT) < 6 and smooth current tracking. This design, which enables parallel processing of distributed drivers with low on-board power, provides an essential testbed for investigating the parametric effects of different motor designs/control methods on external-load compensation", + " 4, any external load exerted on the rotor can be described by (23, 24) where fX, fY and fZ are the components of Fc in the radial (feed), tangential (cutting) and normal (depth-of-cut) directions respectively: T cl a X Y a Zm f f m g f F F g (23) T c cand l Y c X c Z c Y cf h f h f r f r T r F (24) The effects of the MBR compensation on the bearing condition and cutting process were investigated by measuring the bearing sound, vibrations and spindle speed using the microphone, accelerometer and encoder as shown in Fig. 5 (d). The microphone was installed far from the cutter so that the noise from the cutting interface can be isolated for monitoring the bearing conditions. A. D4-Motor System Parameters The D4-motor control law requires the characteristic force coefficients (kjR, kjT, kjZ) and an estimate of the static friction torque bo. The former can be numerically determined once the motor structure and the EM/PM parameters are defined, whereas the latter must be obtained experimentally. A.1 Numerical computed characteristic force coefficients For a given EM/PM pole-pair, the kernel functions ( , , R T Z ) which provide a basis for computing (kjT, kjZ) from (9d) can be determined experimentally [27] or analytically computed using the Lorentz force equation, energy method or Maxwell stress tensor [28~30]", + " The setup also provides a means to verify experimentally the computed tangential-force coefficient kjT that contributes to the spinning torque e (9c). For these objectives, a tension load fs was applied on the rotor in the XY plane and statically maintain the rotor at an angular displacement \u03b8 against a motor torque \u03c4e with an adjustable ball-screw micrometer: T sin cos 0l s s af f m g F (27a) T 0 0 sinl s ff r T (27b) 1 1 2 sin where tan cos L f f w r h r h (27c) The tension fs and displacement \u03b8 were measured by a force sensor through a pulley (Fig. 7a) and an encoder (Fig. 5a). Identical current inputs were supplied to the EMs 3 2ju of Phase A (single-phase operation with one-third of the 2\u00d7NE EMs). The combined torque \u03c4e from the twin-SB motors to maintain an equilibrium at a specified \u03b8 is computed from However, additional X and Y component forces must be generated to compensate for the tension load and rotor weight. As illustrated in Section C.2, the D4-motor can be operated as Mode 1 (u+=u) planar actuator that provides a means to supplement the uniform currents in Phase A that contributes the pure spinning torque to reduce the MBR: T T T T 1 sin cos 2 s a s X Y X X Y Y f m g f u u K K K K K K (29) Fig", + " The current input vector u+ (22b) given by (32a, b) is composed of speed regulation ur and feedforward MBR compensation uc: T0 T MBR compensator 0 No 2 With MBR compensator b r c Y Z cZ Z r f u P uP P (32a) T T T T 1 2 2 where 2 X a Y c X Y X X Y Y f m g f u K K K K K K (32b) Equation (32a) compares two methods, where the scalar PID output \u03c4r can be derived from (4d) with the (motor, sensors, control and friction) parameters listed in Tables 2 and 3 and Fig. 7(b). In (32a~b), the cutting forces (fX, fY) are low-pass-filtered measurements from the dynamometer (Fig. 5d). In Table 3, M1 and M2 refer to the two control methods, without and with MBR compensation, respectively. Sensor specifications: Dynamometer (Kistler 9122AA): 8.1pC/N (FX, FZ), 4.1pC/N (FY) Accelerometer (PCB 356A16): 100mv/g Microphone (GRAS Array): 50 mV/Pa Encoder (HEDSS ISC5208): 10,000 pulses/round Surface profiler (Talyrond 595) The effectiveness of the MBR compensation on the multi-DOF D4-motor and the cutting process were experimentally investigated. The results are presented in Fig. 9 and Table 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000274_icara.2000.4803910-Figure10-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000274_icara.2000.4803910-Figure10-1.png", + "caption": "Fig. 10. Experimental Environments", + "texts": [ + " a) Mobile Robot \u201dPenguin II\u201d: The right of Fig. 5 shows our mobile robot \u201dPenguin II\u201d. It is 40 cm long\uff0c45 cm wide and 42 cm high without wireless antenna. Its body shape is a quadrangular prism cut off at the edges. Penguin II has two drive wheeled in the front and two caster wheeled in the back, and can turn pinwheeling within a 50 centimeters radius. This robot is equipped with a Pentium-M 2.0 GHz (FSB400 MHz) processor, 1 GB memory, RT-Linux for OS and 1 [ms] PD controller for the motor. b) Experimental Overview: Fig. 10 shows the layout for this experiment. The two 1.2 meter long partition boards are placed 2 meters apart in an 8 meter square area. The start is located 2 meters from one board and the planning task is to pass through the two boards. For the purpose of comparison with our planner, we implemented the shortest path search using non-informed A* [9] with efficient optimal search [7] on the real wheeled robot and conducted an experiment in the same situation. We ran a simulation using our planner and the shortest path search as referred to above for the same map configuration" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002455_1.4039395-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002455_1.4039395-Figure1-1.png", + "caption": "Fig. 1 The redundant robot in consideration is an omnidirectional mobile platform equipped with an industrial robot arm", + "texts": [ + " The damped Newton\u2013Raphson method is then used to track the solutions as k varies from 0 to 1 as follows: xl\u00fe1 \u00bc xl a H kl; xl\u00f0 \u00de H0 kl; xl\u00f0 \u00de (6) where a is the damping factor. Our motion planning process is a two-stage procedure where we need to first initialize and then sequentially solve for the optimal solutions. This requires defining a performance criteria and formulating the task constraints based on the robot kinematics for use in an optimization routine. In this section, we describe the first stage. As an example, we consider the optimal motion planning for a redundant mobile manipulator as shown in Fig. 1. This involves the transformation and formulation of an optimal robot position problem into the expanded Lagrangian homotopy to obtain the globally optimum starting point for motion planning. 3.1 Objective Function. Without loss of generality, the manipulability measure [33] is used in this work as the robot performance criteria to be optimized. The robot manipulability can be calculated using 031010-2 / Vol. 10, JUNE 2018 Transactions of the ASME Downloaded From: http://mechanismsrobotics.asmedigitalcollection" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002636_j.mechmachtheory.2018.05.003-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002636_j.mechmachtheory.2018.05.003-Figure5-1.png", + "caption": "Fig. 5. Seal dynamics - the elastic contact approach.", + "texts": [ + " Secondly, all forces acting on the seal, rotor, and housing are considered two-dimensional, which means that none of the forces is acting along the axial direction (perpendicular to the page). It is also important to point out that the angular and linear variables of the seal cannot be geometrically related; hence, to find those variables, one must solve systems of differential equations. 4. A model for seal dynamics In this section, the contacts between the apex-seal and seal-groove, the seal and machine-housing, and the seal and sealspring are modelled as spring-mass-damper systems as detailed in Fig. 5 . In Fig. 5 , the groove-attached XY frame and its origin, O g , is defined at the initial position of a seal-attached frame, X s Y s whose origin, C , is the centre of gravity of the seal. During the rotational motion of the rotor, the origin, C , of the X s Y s frame is displaced from its initial position, O g , to a new position, ( x, y ), measured in the groove frame, XY . As such, the position of the seal centre of gravity, C , can be described by the position vector rc as follows, rc = [ x y ] (10) Also, the seal frame X s Y s undergoes an angular displacement, \u03d5, measured from the groove frame in a right-handed sense", + " The acceleration of the seal\u2019s centre of gravity is expressed as follows: a 1 c = X\u0308 1 \u02c6 i + Y\u0308 1 \u02c6 j (14) where the acceleration components, X\u0308 1 and Y\u0308 1 , are given as follows; { X\u0308 1 = x\u0308 \u2212 4 r \u03c9 2 sin \u03b8 \u2212 \u03c9 2 L i \u2212 \u03c9 2 x \u2212 2 \u03c9 \u0307 y Y\u0308 1 = y\u0308 + 4 r \u03c9 2 cos \u03b8 \u2212 \u03c9 2 y + 2 \u03c9 \u0307 x (15) The seal axial length, d s , is defined as the combination of d s 1 and d s 2 as follows: d s = d s 1 + d s 2 (16) where d s 1 and d s 2 are measured from the seal centre of gravity to the outside and inside ends of the seal, respectively, as shown in Fig. 5 , which also shows the seal circular-arc face of radius r o and its centre, C o . The angle \u03b2o is measured from the seal centre line to the radial line C o c 3 (or C o c 1 ) while the angle \u03b2 is measured from the seal centre line to the radial line C o s . These angles, \u03b2 and \u03b2o , will later on in this paper help determining the locations of points s , s , and s . 5 5 8 9 Y It is essential to determine the special points on the seal onto which the forces and pressures from the seal-groove, seal spring, housing wall, and the working chambers would act. The expressions describe the positions of those points will depend on the relative positions, x and y , of the seal centre of gravity, C , and the seal angular position, \u03d5. Such expressions will be determined in the below sections. 4.1. Vector representation of the critical contact points As shown in Fig. 5 , the contact points on the seal groove are labelled from g 1 to g 4 , in which points g 1 and g 3 are fixed on the groove and located near the tip of the rotor apex while the actual positions of points g 2 and g 4 vary with the seal\u2019s motion. The position vectors from O g to g 1 and to g 3 can be expressed, respectively, as follows: rg 1 = { x g 1 = d s 1 \u2212 \u03b60 \u2212 \u03b4w y g 1 = \u2212W g 2 (17) and rg 3 = { x g 3 = d s 1 \u2212 \u03b60 \u2212 \u03b4w y g 3 = W g 2 (18) where the constant \u03b6 0 is the initial value of the seal protrusion outside the apex groove", + " The position vectors of these points with respect to the moving origin, C , can be expressed respectively as, r s c 2 = [ \u2212d s 2 \u0302 X s \u2212W s 2 \u02c6 Y s ] and r s c 4 = [ \u2212d s 1 \u0302 X s W s 2 \u02c6 Y s ] (19) The contact points on the seal correspond to points g 1 and g 3 are i 1 and i 2 , respectively; the position vectors of these points in the XY -frame are shown as follows: ri 1 = [ ( l 2 \u2212 d s 2 ) cos \u03d5 + W s 2 sin \u03d5 ] \u02c6 X + [ ( l 2 \u2212 d s 2 ) sin \u03d5 \u2212 W s 2 cos \u03d5 ] \u02c6 Y (20) ri 2 = [ ( l 4 \u2212 d s 2 ) cos \u03d5 \u2212 W s 2 sin \u03d5 ] \u02c6 X + [ ( l 4 \u2212 d s 2 ) sin \u03d5 + W s 2 cos \u03d5 ] \u02c6 Y (21) Where l 2 and l 4 are given by the following expressions: l 2 = \u2212x cos \u03d5 \u2212 y sin \u03d5 + d s 2 + ( d s 1 \u2212 \u03b60 \u2212 \u03b4w ) cos \u03d5 \u2212 W g 2 sin \u03d5 (22) and l 4 = \u2212x cos \u03d5 \u2212 y sin \u03d5 + d s 2 + ( d s 1 \u2212 \u03b60 \u2212 \u03b4w ) cos \u03d5 + W g 2 sin \u03d5 (23) Fig. 5 depicts the contact point s 5 between the seal tip and the machine housing; the vector representation of this point will be defined below. As the seal slides on the machine housing, point s 5 moves between the points c 1 and c 3 on the seal tip. The positions of c 1 and c 3 with respect to the origin, C , of the moving frame X s Y s are as follows: r s c 1 = [ d s 0 \u0302 X s \u2212W s 2 \u02c6 Y s ] and r s c 3 = [ d s 0 \u0302 X s W s 2 \u02c6 Y s ] (24) Consequently, the position of s 5 in frame XY can be found as: rs 5 = rc + ( d s 1 \u2212 a o ) \u0302 X s + b 5 \u0302 Y s (25) which can be re-written as, rs 5 = ( x + ( d s 1 \u2212 a o ) cos \u03d5 \u2212 b 5 sin \u03d5 ) \u0302 X + ( y + ( d s 1 \u2212 a o ) sin \u03d5 + b 5 cos \u03d5 ) \u0302 Y (26) where a o and b 5 are shown in Fig. 5 and can, respectively, be calculated as, a o = r o ( 1 \u2212 cos \u03b2) (27) and b 5 = r o sin \u03b2 (28) Since point s 5 always stays attached to X -axis, the Y component of point s 5 has to be zero. Hence, the coefficient of the \u02c6 component of Eq. (26) can be written as: y + ( d s 1 \u2212 a o ) sin \u03d5 sin \u03d5 + b 5 cos \u03d5 = 0 (29) which can be substituted and rearranged to give y + d s 1 sin \u03d5 \u2212 r o ( 1 \u2212 cos \u03b2) sin \u03d5 + r o sin \u03b2 cos \u03d5 = 0 (30) Value of angle \u03b2 can then be found by solving Eq. (30) as: \u03b2 = sin \u22121 [( 1 \u2212 d s 1 r o ) sin \u03d5 \u2212 y r o ] \u2212 \u03d5 (31) Of note is that if the radius of curvature, r o , of the seal tip is large, the value of \u03b2 will approach zero ( \u03b2\u223c= 0), the pressure forces on the seal tip will be acting parallel to the X s axis", + " This difference, x s 5 , can be calculated as follows: x s 5 = x + ( d s 1 \u2212 a o ) cos \u03d5 \u2212 b 5 sin \u03d5 \u2212 d s 1 + \u03b4w (33) where x s 5 determines the elastic contact force between the seal and the housing as follows, F w = { k w x s 5 i f x s 5 > 0 0 i f x s 5 \u2264 0 (34) It has been assumed that sliding contact exists between the seal and the seal spring. When the seal is at its initial position, the seal-spring contact point is s 7 . Point s 7 is on the surface of the seal on the negative side of the X s -axis as shown in Fig. 5 , this point can be expressed as rs 7 = [ d s 2 cos \u03d5 \u2212 x \u2212 ( W s 2 \u2212 b 6 ) sin \u03d5 ] \u02c6 X + [ ( W s 2 \u2212 b 6 ) cos \u03d5 ] \u02c6 Y (35) During the motion of the apex seal, the seal-spring contact point shifts away from s 7 , this new contact point is labelled s 6 . The distance b 6 measured from c 2 to s 6 can be expressed as b 6 = W s 2 + b 5 \u2212 ( d s 1 \u2212 a o ) tan \u03d5 \u2212 d s 2 tan \u03d5 (36) The distance from the seal centre of gravity to point s 6 can be calculated as: x s 6 = ( d s 1 \u2212 a o ) + d s 2 cos \u03d5 \u2212 x s 5 (37) The position of point s 6 when compared to its initial position, d s 2 \u2212 \u03b4w , will determine the spring deflection, x s 6 , and this will in turn be used to calculate the force exerted by the spring, F sp , on to the inner end of the seal. The spring deflection is given as follows: x s 6 = x s 6 \u2212 ( d s 2 \u2212 \u03b4w ) = ( d s 1 \u2212 a o ) + d s 2 cos \u03d5 \u2212 x s 5 \u2212 ( d s 2 \u2212 \u03b4w ) (38) The following spring force exits only if x s 6 > 0, F sp = k s ( x s 6 + \u03b4s ) (39) It is now essential to determine the distances between the points g 1 , g 2 , g 3 , and g 4 and their corresponding contact points on the seal \u2013 shown as b 1 , b 2 , b 3 , and b 4 in Fig. 5 . If the distances are positive, there are no forces between the seal and the groove; otherwise, contact forces will exist. The distances b 1 , b 2 , b 3 , and b 4 can be calculated as follows: b 1 = ( rc + rs c 2 \u2212 rg 1 ) \u00b7 \u02c6 Y s = \u2212x sin \u03d5 + y cos \u03d5 \u2212 W s 2 + ( d s 1 \u2212 \u03b60 \u2212 \u03b4w ) sin \u03d5 + W g 2 cos \u03d5 (40) b 2 = ( rc + rs c 2 \u2212 rg 1 ) \u00b7 \u02c6 Y = y \u2212 d s 2 sin \u03d5 \u2212 W s 2 cos \u03d5 + W g 2 (41) b 3 = \u2212 ( rc + rs c 4 \u2212 rg 3 ) \u00b7 \u02c6 Y s = x sin \u03d5 \u2212 y cos \u03d5 \u2212 W s 2 \u2212 ( d s 1 \u2212 \u03b60 \u2212 \u03b4w ) sin \u03d5 + W g 2 cos \u03d5 (42) b 4 = \u2212 ( rc + rs c 4 \u2212 rg 3 ) \u00b7 \u02c6 Y = \u2212y + d s 2 sin \u03d5 \u2212 W s 2 cos \u03d5 + W g 2 (43) The contact forces between the seal and the seal groove at points g 1 , g 2 , g 3 , and g 4 are denoted by F g 1 , F g 2 , F g 3 , and F g 4 , respectively. The forces F g 2 and F g 4 act along the lines that are perpendicular to the walls of the seal groove at points g 2 and g 4 , respectively; while the forces F g and F g act along the lines that are perpendicular to the seal side surfaces at points i 1 1 3 and i 2 , respectively. Those lines of action are shown in Fig. 5 . The conditions from which the magnitudes and directions of the seal and seal-groove forces can be determined are shown in Algorithm 1 below: Algorithm 1 Conditions to determine the magnitude and direction of the seal and seal-groove forces. 1 Begin 2 if b 1 < 0 then 3 F g 1 = | b 1 | k g 1 ( \u2212\u03b31 \u03bc2 \u0302 X s + \u0302 Y s ) 4 else F g 1 = 0 5 end if 6 if b 2 < 0 then 7 F g 2 = | b 2 | k g 2 ( \u2212\u03b32 \u03bc2 \u0302 X + \u0302 Y ) 8 else F g 2 = 0 9 end if 10 if b 3 < 0 then 11 F g 3 = | b 3 | k g 3 ( \u2212\u03b33 \u03bc2 \u0302 X s \u2212 \u02c6 Y s ) 12 else F g 3 = 0 13 end if 14 if b 4 < 0 then 15 F g 4 = | b 4 | k g 4 ( \u2212\u03b34 \u03bc2 \u0302 X \u2212 \u02c6 Y ) 16 else F g 4 = 0 17 end if 18 End where k g i is the stiffness of the groove material at location g i and \u03b3 i signifies the sign of the relative velocity at the same location ( i = 1 , 2 , 3 , 4 ) ; \u03bc2 is the friction coefficient between the seal and seal-groove. When the value of b i is greater than zero ( b i > 0) pressures from the working chambers are able to enter the cavity between the seal and seal groove, which will affect the seal dynamics. These forces will act at points s 1 , s 2 , s 3 , and s 4 as shown in Fig. 5 , where s 1 , s 2 , s 3 , and s 4 are the mid-points between i 1 \u2212 c 1 , c 2 \u2212 i 1 , i 2 \u2212 c 3 , and c 4 \u2212 i 2 , respectively. The positions of such points can be expressed in the XY frame as follows: rs 1 = [( d s 1 \u2212 l 1 2 ) cos \u03d5 + W s 2 sin \u03d5 ] \u02c6 X + [( d s 1 \u2212 l 1 2 ) sin \u03d5 \u2212 W s 2 cos \u03d5 ] \u02c6 Y (44) rs 2 = [( l 2 2 \u2212 d s 2 ) cos \u03d5 + W s 2 sin \u03d5 ] \u02c6 X + [( l 2 2 \u2212 d s 2 ) sin \u03d5 \u2212 W s 2 cos \u03d5 ] \u02c6 Y (45) rs 3 = [( d s 1 \u2212 l 3 2 ) cos \u03d5 \u2212 W s 2 sin \u03d5 ] \u02c6 X + [( d s 1 \u2212 l 3 2 ) sin \u03d5 + W s 2 cos \u03d5 ] \u02c6 Y (46) rs 4 = [( l 4 2 \u2212 d s 2 ) cos \u03d5 \u2212 W s 2 sin \u03d5 ] \u02c6 X + [( l 4 2 \u2212 d s 2 ) sin \u03d5 + W s 2 cos \u03d5 ] \u02c6 Y (47) where l 1 = d s 2 + d s 0 \u2212 l 2 and l 3 = d s 2 + d s 0 \u2212 l 4 It worthy of note here that the distance d s 0 is measured along X s from the centre of gravity of the seal to the point where the curved end starts" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003466_j.mechmachtheory.2019.06.017-Figure15-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003466_j.mechmachtheory.2019.06.017-Figure15-1.png", + "caption": "Fig. 15. Illustration of the tapered involute pinion with double-crowned tooth surfaces.", + "texts": [ + " Under the same parameters, the maximal tooth width of the tapered involute pinion could be much greater than the face-gear ( Fig. 14 ). For the purpose of shifting the contact pattern to an expected position on the face-gear surface, and avoiding edge contact of the face-gear drive ( Section 6 ), another type of geometry of the tapered involute pinion is proposed. Both profile modification and longitudinal crowning are performed for the pinion of the second geometry. This section covers the doublecrowned tapered involute pinion generation by a generating worm. As shown in Fig. 15 , a tapered involute pinion with tooth surface including profile crowning and longitudinal crowning is presented. Assuming that the three related surfaces ri , s and w are meshing simultaneously during the generation of pinion surfaces (where w is the tooth surface of the generating worm). Two processes are implemented: (i) worm thread surface generation by a rack-cutter, and (ii) generation of the pinion surface by the worm ( Fig. 16 ). The worm thread surface w is generated by ri of the rack-cutter" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003281_j.triboint.2019.04.005-Figure25-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003281_j.triboint.2019.04.005-Figure25-1.png", + "caption": "Fig. 25. Schematic diagram", + "texts": [ + "4 Pe ak o pe ra tio n pr es su re /M Pa Compression ratio/% Fig. 24. Maximum operation pressure 5. Dynamic oil injection test A dedicated test tooling is developed according to the specific load spectrum to validate the curve in Fig. 24. Meanwhile, colored gear oil is injected into the sealing mechanism through the test mandrel with specified pressure. After the test, the sealing performance is validated by the presence of the colored gear oil in the bearing cavity. 5.1 Test rig The schematic diagram of the dynamic oil injection test is shown in Fig. 25. Fig. 26 shows the test rig, which consists of a motor, a rotating shaft, a tested bearing, a hydraulic system with colored gear oil and a load device. The in-house-made test tooling is shown in Fig. 27. M ANUSCRIP ACCEPTE D Fig. 26. Test rig and partial detail Fig. 27. Test tooling The test is carried out in three cycles (Table II.) to 5L10 (L10 is the base service life of the tapered roller bearing in hours.). Fr, Fa are the radial and axial load, respectively. The ambient temperature is 22~28\u2103 and the humidity is 34~58% RH" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001550_oceans.2010.5664384-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001550_oceans.2010.5664384-Figure1-1.png", + "caption": "Fig. 1 Transform schematic diagram of depth and height", + "texts": [ + " Problem Formulation During the task of bottom following, the main purpose is to keep the vehicle at a constant height from the seafloor in order to assess the oil pipes below, or to scan a given area by side scan sonar (SSS). In order to keep the AUV at a constant height above the seafloor, bottom following control can be used to keep the AUV operating at a fixed vertical point in relation to the seafloor, diving control can be used to achieve this using measurements from the altimeter and pressure sensors to assess the AUV\u2019s distance from the seafloor and surface, respectively. The transform schematic diagram is shown in Fig. 1, and the SNAME notation will be used for denoting the vehicle model and motion. 0 iP P\u22c5 \u22c5\u22c5 denote the positions of the AUV. 0h is the desired height satisfying 0 0h = . At it t= , the height and depth of vehicle are measured as ih and iz . Then the diving command can be calculated as 978-1-4244-4333-8/10/$25.00 \u00a92010 IEEE In fact, the final objective is to realize the bottom following control in three dimensions. However, according to practical operational applications in AUVs, the six degrees of freedom (DOF) nonlinear equations of motion can be separated into three lightly interacting subsystems, including diving, steering and speed control" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000176_tia.2009.2013568-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000176_tia.2009.2013568-Figure1-1.png", + "caption": "Fig. 1. Winding arrangement of an elementary two-pole three-phase IM.", + "texts": [ + " According to the new proposed approach, detailed analytical calculations of the zero-sequence components are performed, respectively, of the air-gap flux and the stator voltage. The final mathematical expressions confirm the results of the experimental tests, performed on a laboratory IM drive described later in this paper, showing the harmonic ripples of the flux and voltage in presence of additional highfrequency currents. II. FLUX CALCULATION IN PRESENCE OF SATURATION AND VARIATION OF SATURATION Calculation of the flux linked with the stator is performed with reference to the elementary two-pole three-phase IM shown in Fig. 1. The stator flux linkage equation for phase a, written in terms of the stator and rotor currents and the self- and mutual inductances, is \u03bbas =Llsias + Lasas \u00b7 ias + Lasbs \u00b7 ibs + Lascs \u00b7 ics + Lasar1 \u00b7 iar1 + Lasbr1 \u00b7 ibr1 + Lascr1 \u00b7 icr1 + Lasar3 \u00b7 iar3 + Lasbr3 \u00b7 ibr3 + Lascr3 \u00b7 icr3 (1) where Lisjs = \u03c0\u222b \u2212\u03c0 \u03bc0 r \u00b7 l g [Nis(\u03d5) \u00b7 Njs(\u03d5)] \u00b7 d\u03d5, i, j = a, b, c Lisjr1 = \u03c0\u222b \u2212\u03c0 \u03bc0 r \u00b7 l g [Nis(\u03d5) \u00b7 Njr1(\u03d5 \u2212 \u03b8r)] \u00b7 d\u03d5, i, j = a, b, c Lisjr3 = \u03c0\u222b \u2212\u03c0 \u03bc0 r \u00b7 l g [Nis(\u03d5) \u00b7 Njr3(3\u03d5 \u2212 3\u03b8r)] \u00b7 d\u03d5, i, j = a, b, c" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000682_s12555-012-0508-0-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000682_s12555-012-0508-0-Figure4-1.png", + "caption": "Fig. 4. The curves C and C and the rays r and r of Proposition 1.", + "texts": [ + "E S\u2229 Proof: Firstly, it is clear that the boundary of E1\u2229 S2 is given by two of the following curves: { : ( ) 0}x F x\u2208 =U { : det( ( ), ) 0} for 1, 2 i x f x x i\u2208 = =U Let denote these curves by C and .C As these curves are convex or concave, we know that the tangent line at the origin is contained in E1\u2229 S2 or in its complementary set. Hence, we have the following cases. If the tangent lines at the origin to C and ,C respectively, are contained in E1\u2229 S2, we can choose r and r as the rays given by these lines (see Fig. 4(a)). Carmen P\u00e9rez and Francisco Ben\u00edtez 924 If the tangent line to C is not contained in E1\u2229 S2 and the tangent line to C does, r will be the tangent line to C and r a ray whose slope is between the slope of the tangent line to C and the slope of the tangent to r (see Fig. 4(b)). In the case where the tangent line to C is contained in the intersection and the tangent line to C is not contained in it, we define the rays in the same way (see Fig. 4(c)). Finally, if the tangent lines at the origin of C and C are not contained in E1\u2229 S2, we choose as r and r the rays starting from the origin and passing through the intersection points of C and C with the set U (see Fig. 4(d)). Remark 4: It is possible to define the same result for 41 ,E S\u2229 \u2260 \u2205 2 1 E S\u2229 \u2260\u2205 and 2 3 .E S\u2229 \u2260 \u2205 Definition 4: Let (1) be a switched system that satisfies the conditions in the previous proposition and 0 C(0, , )x r r\u2208 \u2229 U a state. We say that a switching law is of type II if we switch whenever the trajectory intersects r or .r In this section, we use the previous switching laws in order to present a method that ensures the switched convergence of a switched nonlinear system in the plane. Consider the switched nonlinear system given by (1)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003238_s40684-019-00095-4-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003238_s40684-019-00095-4-Figure4-1.png", + "caption": "Fig. 4 a Diagram of single-link manipulator, where lc, l, m and J denote the centroid location, length, mass and moment of inertia with respect to centroid the link. b Diagram of two-link revolute manipulator, where lc1, lc2, l1, l2, m1, m2 and J1, J2 denote the centroid location, length, mass and moment of inertia with respect to centroid of the", + "texts": [ + " Although assumption ||z\u0304|| \u2264 zM is a relatively strong assumption, function \u03d5 (35) Vj(t \u2212 s (j + 1)) \u2264 Vj(t + s (j))e\u2212 (ts(j+1)\u2212ts(j)) + \u2215 (1 \u2212 e\u2212 (ts(j+1)\u2212ts(j))). (36) Vj+1(t + s (j + 1)) \u2264 Vj(t + s (j))e\u2212 (ts(j+1)\u2212ts(j)) + \u2215 (1 \u2212 e\u2212 (ts(j+1)\u2212ts(j))) + 2\u0394. (37) Vj(t) \u2264 V(0)e\u2212 (t\u2212t0) + \u2215 (1 \u2212 e\u2212 (t\u2212t0)) +2\u0394 j\u2211 k=0 e\u2212 (t\u2212ts(k)) \u2212 2\u0394e\u2212 t, (38)eTe \u2264 2V(0)e\u2212 t + 2 \u2215 (1 \u2212 e\u2212 t) + 4\u0394 M . can be designed to bound the input of neural network, e.g. \ud835\udf19 = ||e|| + \ud835\udf19M(||z\u0304||\u2215zM)k where k > 1. Two simulation examples based on a single-link manipulator and a two-link revolute manipulator whose diagrammatic sketches are shown in Fig.\u00a04 are given in this section. 4.1 Example 1 In the first example, comparisons between the control performance of the ESC and the NNC under different values of k6 are given. All units are in SI, and dynamics of the system is given by (1) with M = \u03b81, C = 0 and G = \u03b82\u00a0sin\u00a0(q), where \u03b81 = mlc2 + J and \u03b82 = mglc. In the simulation, we have m = 200\u00a0kg, l = 0.2\u00a0m, lc = l/2, J = mlc 2/3, g = 9.8\u00a0m/s2, leading signal q0 = \u03c0/4\u00a0 -\u00a0 \u03c0/4\u00a0 cos\u00a0 (2\u03c0t/5)rad and disturbance f = 20 sin (q + 2q\u0307)N \u22c5 m \u22c5 rad\u22121 . Control parameters are chosen as K1 = 5, K2 = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003616_j.promfg.2019.08.022-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003616_j.promfg.2019.08.022-Figure2-1.png", + "caption": "Fig. 2. Scheme of cutting samples for mechanical tests for tensile and impact bending from blanks obtained by laser deposition of Inconel 625 alloy on steel 15Cr11MoW", + "texts": [ + " Alloying elements of the deposited metal were distribute along the vertical axis of the deposited layers, and they were determined using a scanning electron microscope (Phenom ProX and Mira Tescan) locally in increments 0.25 mm. Samples for tensile tests were manufactured in accordance with GOST 6996-66 (type II) in such a way, that that the fusion zone of the base metal and the cladding is located in the middle of the working part of the samples, while one half of the sample consisted of the base metal, and the second consisted of the weld (Fig. 2). 166 M. Kuznetsov et al. / Procedia Manufacturing 36 (2019) 163\u2013175 4 Author name / Procedia Manufacturing 00 (2019) 000\u2013000 Part of the blanks after surfacing was subjected to heat treatment in form of tempering at the temperature of 680 \u00b0C with an exposure of 2.5 hours. The tests were carried out at room temperature according to DIN EN ISO 6892-1: 2009 using a Zwick / Roell Z250 tensile testing machine. To determine the values of impact toughness of welded joints, blanks were cut from witness samples, which made it possible to manufacture impact specimens with a V-notch located in the zone of fusion of the base metal and cladding" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003106_j.inoche.2019.01.015-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003106_j.inoche.2019.01.015-Figure2-1.png", + "caption": "Fig. 2. Scheme of the compression and hot extrusion process.", + "texts": [ + " Results in Physics 12 (2019) 718\u2013724 2211-3797/ \u00a9 2018 The Authors. Published by Elsevier B.V. This is an open access article under the CC BY-NC-ND license (http://creativecommons.org/licenses/BY-NC-ND/4.0/). T 40 \u00b1 0.3 g, and the backfilling chips were heated to 523 K in the mold. Then the mold with chips was quickly placed on a plate of the pressure machine, and the chips were subsequently compressed into a billet by unilateral pressing of 640MPa and a compacting time of 45 s, as illustrated in Fig. 2(a) and (b). Following the above treatment, the same mold with the three formed billets was heated to the extrusion temperature and held for 40min. The average heating rate was 0.5 K/s, and the holding temperature was controlled to\u00b10.5 K by a PID controller, then hot extrusion (Fig. 2(c)) was performed in air at an extrusion temperature of 673 K. Although the quality of production improved with an increasing extrusion ratio [16], energy savings [5] and recrystallization grain size are inversely proportional to the deformation degree [17]. Thus an extrusion ratio of 25:1 was adopted, and the extrusion pressure was 864MPa. In addition, the as-cast ingot was also employed and formed by hot extrusion for comparison, and the extruded sample labelling used in this work is summarized in Table 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000793_17452759.2010.527010-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000793_17452759.2010.527010-Figure4-1.png", + "caption": "Figure 4. Uniform slicing vs. adaptive slicing of a sample part.", + "texts": [ + "2 Implementation of the slicing concept In this section an example of an adaptively sliced part will be presented. In order to visualise the potentials of adaptive slicing the number of facets of the illustrative part is limited. As long as an examined part has regions with vertical facets the mentioned benefits will occur. The sample part initially has been sliced uniformly with a layer thickness of 0.1 mm and consists of 60 layers. With the adaptive slicing approach it is possible to significantly reduce the staircase effect at cost of only 16 additional layers (Figure 4). The average layer thickness is 0.08 mm whereas the limits are tmin 0.04 mm and tmax 0.2 mm. In order to improve the economic efficiency minimum layer thickness can also be kept constant whereas the number of layers can be reduced to a certain extent. Figure 5 illustrates the potential reduction of build time by adaptively slicing an engine block. In this case the Figure 3. Inclusion relations between original and layer model. D ow nl oa de d by [ D al ho us ie U ni ve rs ity ] at 0 6: 43 3 1 D ec em be r 20 14 adaptive concept permits a reduction of the number of layers of about 29%" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002349_s12289-017-1388-x-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002349_s12289-017-1388-x-Figure1-1.png", + "caption": "Fig. 1 Radial-axial ring rolling (schematic representation)", + "texts": [ + "eywords Ring rolling . Contact geometry . CAD-analytical approach . Force prediction Thanks to the wide range of applications, as well as for its important features such as short cycle time, high energy-efficiency, favorable grain orientation, and versatility [1], radialaxial ring rolling, Fig. 1, has become a highly widespread metal forming process. The ring rolling process design phase, as well that of all manufacturing processes, is a critical stage where the mutual interaction among workpiece blank dimensions, material properties, process conditions, and parameters must be taken into account simultaneously. During the design phase, process engineers should make some preliminary calculations in order to identify the ring rolling mill to utilize for the realization of the ring. To this aim, one of the most important parameter for a proper choice of the ring rolling machine is the maximum forming force required during the process", + " Ring initial geometry [mm] Ring final geometry [mm] Mandrel feeding speed [mm/s] Literature models for the calculation of the projection of the contact arc length Previous estimations of the contact geometry between ring and tools in the mandrel-main roll deformation gap are based on simplifications, especially concerning the estimation of the projection of the contact arc between ring and tools. Consequently, as it will be shown in the result section, in some case, the utilization of previous literature models lead to an underestimation of the radial forming force. In the literature, many contributions concerning the contact geometry estimation in the ring rolling process, in terms of formulation for the estimation of the projection of the contact arc between ring and tools in the mandrel-main roll deformation gap, Fig. 1, have been formulated. The proposed equations: Yang et al. [9], Qian et al. [10] and Parvizi et al. [6] are reported in Eqs. (4)\u2013(6) respectively. The terminology is listed and explained in the following Table 1. The variation of thickness between the onset and the exit of the deformation gap is defined as\u0394s = s0 \u2212 sF whereas the average radius of the ring as Ravg = (Ro + Ri)/2. Lc\u2212Y \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi Rs 2\u2212 Ravg \u00fe s0=2 2\u2212 Rs \u00fe s1=2\u00fe Ravg 2\u2212Rs 2 2 Rs \u00fe s1=2\u00fe Ravg \" #2 vuut \u00f04\u00de Lc\u2212Q \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2\u0394s 1=Rs\u00f0 \u00de \u00fe 1=Rm\u00f0 \u00de \u00fe 1=Ro\u00f0 \u00de\u2212 1=Ri\u00f0 \u00de s \u00f05\u00de Lc\u2212P \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi Rs 2\u2212 Rs \u00fe Rm\u2212\u0394s\u00f0 \u00de2 \u00fe Rs 2\u2212Rm 2 2 Rs \u00fe Rm\u2212\u0394s\u00f0 \u00de \" #2 vuut \u00f06\u00de In the result section, these three models as well authors combined CAD-analytical model are applied to three different study cases in order to show that: i) the results obtainable with these three models are very similar to each other and ii) that the developed approach allows overcoming their limitations, resulting in a better estimation of the projection of the contact arc between ring and tools, confirmed by a more precise calculation of the radial forming force" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000841_mmar.2012.6347891-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000841_mmar.2012.6347891-Figure1-1.png", + "caption": "Fig. 1. Kinematics Frames of an underwater vehicle.", + "texts": [ + " After modelling the entire problem we will finally simulate some thruster faults in different time, and show results of this work in some graphs. In modelling the dynamics of underwater vehicles, it is convenient to work with two reference systems, one called inertial frame, which is located on land identified by (\ud835\udc4b0\ud835\udc4c0\ud835\udc4d0) and other called body-attached frame, located at the center of mass of the vehicle (XYZ), the formalism of Newton-Euler is very useful to describe the complex motion of an underwater vehicle with all six degrees of freedom possible, as seen in Fig. 1. Let be u,v,w linear velocity and p,q,r as angular velocity of the vehicle, described at body-attached frame, which are particularly well known as surge, sway, heave, roll, pitch and yaw, respectively. To transforms those velocities at bodyattached frame to inertial frame, we can use a transformation matrix . 978-1-4673-2124-2/12/$31.00 \u00a92012 IEEE 184 \ud835\udc3d1(\ud835\udf021) = \u23a1 \u23a3 \ud835\udc503\ud835\udc501 \u2212\ud835\udc603\ud835\udc502 + \ud835\udc503\ud835\udc601\ud835\udc602 \ud835\udc603\ud835\udc602 + \ud835\udc503\ud835\udc502\ud835\udc601 \ud835\udc603\ud835\udc501 \ud835\udc503\ud835\udc502 \u2212 \ud835\udc602\ud835\udc601\ud835\udc603 \u2212\ud835\udc503\ud835\udc602 + \ud835\udc601\ud835\udc603\ud835\udc502 \u2212\ud835\udc601 \ud835\udc501\ud835\udc602 \ud835\udc501\ud835\udc502 \u23a4 \u23a6 . (1) The transformation matrix of angular velocity is given by: \ud835\udc3d2(\ud835\udf022) = \u23a1 \u23a3 1 \ud835\udc602\ud835\udc611 \ud835\udc502\ud835\udc611 0 \ud835\udc501 \ud835\udc602 0 \ud835\udc602 \ud835\udc501 \ud835\udc502 \ud835\udc501 \u23a4 \u23a6 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001973_10402004.2017.1323146-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001973_10402004.2017.1323146-Figure3-1.png", + "caption": "Figure 3. Diagram of high-speed test rig used for the current study. The glass disc and the steel ball both have surface speeds of up to 20 m/s.", + "texts": [ + " We also investigated the film composition using oil-soluble and water-soluble fluorescent dyes and measured film temperatures and traction (friction) by infrared microscopy. The high-speed test rig used here is a modified EHL rig (manufactured by PCS Instruments, Acton, UK) as described in detail in an earlier publication (Hili, et al. (23)) and is, in essence, a higher speed version of those used in many previous investigations into emulsions (Ratoi-Salagean (6); Wan, et al. (11); Ratoi-Salagean, et al. (16), (17); Yang, et al. (18)). This is shown in Fig. 3. The test contact is that between a glass disc and a 19.05-mm-diameter steel ball, each driven by a separate electric motor, the latter being capable of 20,000 rpm (Fig. 4), so that, in pure rolling, the entrainment speed approached 20 m/s. The contact was illuminated with white light and the interference image viewed through an optical microscope. This allowed film thickness to be measured by an ultra-thin-film interferometry technique (Johnston, et al. (24)). It should be noted that the surface energy of the silica-coated disc used for these measurements was approximately 300 mmJ/m2, which is close to that of the glass disc used for the fluorescence tests" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002110_whc.2017.7989929-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002110_whc.2017.7989929-Figure1-1.png", + "caption": "Figure 1. Force displays developed using granular-jamming technology; (a) wearable force display, (b) glove-type display, (c) and (d) tabletop-type displays, and (e) handheld-type display", + "texts": [ + " Instead of using conventional brakes, a granular-jamming technology is used in passive force displays, which allows for the variable-stiffness control of a particle-containing bag by evacuating the air inside it. This paper refers to the mechanical element that uses granular jamming as the particle mechanical constraint (PMC). The PMC comprises a soft plastic bag that contains particles such as polystyrene foam beads and coffee beans. Hence, it is soft, lightweight, and suitable for systems that interact with the human body. Mitsuda et al. [3] developed a wearable force display wherein a cylindrical PMC fixed on an operator\u2019s trunk below the arm helps in constraining the hand motion (Fig. 1 (a)). This display allows operators to move their hands freely in any direction with the help of the soft, lightweight mechanism. PMCs have also been used in other types of force displays. Glove-type force displays help in constraining the finger motion using PMCs fixed on or under the glove (Fig. 1 (b)) [4\u20136]. In tabletop-type force displays, the hand or finger motion is constrained using PMCs that help connect the body of the operator to the ground (Figs. 1 (c) and (d)) [7\u201310]. In handheld-type force displays, a force sensation is directly produced via the stiffness of the PMC, which the operator grips with his/her hand (Fig. 1 (e)) [11]. PMCs have been also *Research supported by MEXT-Supported Program for the Strategic Research Foundation. Takashi Mitsuda is a member of the College of Information Science and Engineering, Ritsumeikan University, Kusatsu, Shiga, 525-8577 Japan (phone/fax: +71-77-561-5068; e-mail: mitsuda@is.ritsumei.ac.jp). used for controlling the stiffness of soft robotic arms [12\u201316] or deformable robots [17\u201319]. PMCs can be deformed into various shapes, which is one of the reasons for their wide application", + " The increase in stiffness was largely proportional to the number of fabric layers, but not when the number of layers was 48 and 96, which requires further analysis. C. Inner vacuum pressure Fig. 11 shows the elongation stiffnesses of the FJSs obtained using 16 fabric layers at different inner pressures. The figure shows that the stiffness increases with the increase in the interior vacuum pressure, which is in accordance with the PMCs. A tabletop-type force display was developed using an FJS with 32 piled fabrics, as shown in Fig. 1 (d). The FJS was fixed on a cylinder with a diameter of 88 mm. The reaction force obtained by pushing the center of the FJS with a 15-mm-diameter rod was measured, as shown in Fig. 12. Fig. 13 shows the results, which indicate that the reaction force can be controlled via the vacuum pressure inside the FJS. When an FJS with 96 piled fabrics (as opposed to 32 piled fabrics) was used in the force display, the reaction forces at a displacement of 10 mm were 26.8 and 1.5 N for inner pressures of 80 kPa and atmospheric pressure, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001718_s11431-011-4499-5-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001718_s11431-011-4499-5-Figure1-1.png", + "caption": "Figure 1 ED\uff1dFG=HK, ED//FG//HK.", + "texts": [ + " In order to demonstrate the meaning and derivative process of mobility formula expressed in terms of constraint X jm of virtual loop and prove the correction of the method of analyzing mechanism mobility through virtual loop from anther perspective, several concepts related to virtual loop have to be introduced briefly. A link group is regarded as the combination of non-coincident links between any independent loop and its adjacent loop in a spacially multi-loop mechanism. The mobility of a link group can be equal to 0, or less than 0, or more than 0. The links of a link group can be driven link or driving link . The link group defined here is a generalized group, not the Assur group with zero freedom. For the mechanism shown in Figure 1, it contains four independent loops, and has four link groups, ABCD, EFG, HK and PRM, as shown in Figure 2. To describe the motion transmission manner of two ad- jacent loops by a unifying concept, the concept of virtual kinematic pair is defined here. Assume that links M and J are responsible for the motion transmission between two adjacent loops. No matter whether they are adjacent, it is supposed that they are connected by a kinematic pair, which is regarded as a virtual kinematic pair, short for virtual pair, denoted as , 1 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000873_iembs.2010.5627736-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000873_iembs.2010.5627736-Figure4-1.png", + "caption": "Fig. 4. The powered transfemoral prosthesis.", + "texts": [ + " Once intent to flex or extend the knee is known, the magnitude of the angular velocity for the impedance set-point is obtained by projecting the corresponding data point onto its principal axis via PCA. A representative example of the corresponding PCA projections for subject 1 is shown in Fig. 3, where the xp axis corresponds to the PCA projection of the flexion and extension data along the principal component of that data. The volitional knee controller was implemented on each of the three amputee subjects with the powered transfemoral prosthesis shown in Fig. 4 and described in detail in [1]. Note that the prosthesis used in these experiments also contains a powered ankle, although the ankle was not explicitly commanded in these experiments, but rather remained in a \u201cneutral\u201d configuration. In order to characterize the effectiveness of the volitional controller for purposes of moving the knee joint, an experiment was conducted in which each subject was instructed to track various types of knee joint angle movements. Aside from the QDA and PCA parameters extracted from the EMG intent database, all subjects utilized the same set of volitional control parameters for the powered prosthesis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000945_ssd.2010.5585533-Figure16-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000945_ssd.2010.5585533-Figure16-1.png", + "caption": "Figure 16. 3-D flux path trough the stator lamination and the two claws.", + "texts": [ + " These are described as follows: \u2022 axially in the stator yoke (see figure 12), \u2022 radially down through the collector and crossing the air gap down to the magnetic ring (see figure 12), \u2022 radially then axially in the magnetic ring (see figure 12), \u2022 axially-radially in the claws (see figure 13), \u2022 radially up in air gap facing the stator laminations (see figure 14), \u2022 radially up in the stator teeth and circumferentially in the stator core back (see figure 15), \u2022 radially down in stator teeth facing the two adjacent claws (see figure 16), \u2022 radially-axially in the adjacent claws (see figure 16), \u2022 axially then radially in the magnetic ring (see figure 17), \u2022 radially up through the air gap and the mag netic collector on the other side of the machine (see figure 17), \u2022 axially in the other side of the stator yoke (see figure 18). 3.2.2. Leakage Flux Paths Not all the flux produced by the excitation winding and the armature contributes to the EMF generation. Different leakage fluxes have been distinguished in the CPAES. The two main ones are \u2022 the leakage flux linking adjacent claws, \u2022 tow dimensional flux paths which flow through the magnetic circuit as follows: - axially in the stator yoke, - radially down through the magnetic collector and air gap el, - radially then axially in the magnetic rings holding the claws, - axially in the claw, - radially in air gap e2 and in the stator teeth" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003552_1350650119866037-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003552_1350650119866037-Figure1-1.png", + "caption": "Figure 1. Schematic diagram of contact between single asperity and plane.", + "texts": [ + " Therefore, according to the integral idea, the contact between the asperity and the cylindrical surface can be simplified as the contact between the asperity and the plane when the contact of a single asperity is considered. The widely used ZMC model18 is adopted in this paper to consider the elastic\u2013plastic contact of a single asperity. In this model, the deformation of a single asperity is divided into three stages, namely, completely elastic stage, elastic\u2013plastic stage, and completely plastic stage, and the boundary condition is the normal deformation of the asperity !. As shown in Figure 1, z is the height of the asperity and d is the distance between the plane after contact deformation and the average height line of the asperities, also known as the contact distance ! \u00bc z d \u00f01\u00de The tooth contact can be equivalent to the contact of two cylinders by the meshing point, so the line contact between rough cylindrical surfaces is considered. According to the Hertz contact theory and the idea of Greenwood and Williamson,21 the contact of two rough cylindrical surfaces can be simplified into the contact between a smooth curved surface and a nominally flat rough plane" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002284_ssrr.2017.8088167-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002284_ssrr.2017.8088167-Figure3-1.png", + "caption": "Fig. 3 Normalized energy stability margin (NESM)", + "texts": [ + " Basically, the stability index mentioned in Chapter I targets robot walking motion. The crawling motion described above can be regarded as walking motion of five feet if the torso is regarded as a foot and it is considered that the index of stability proposed for the conventional walking motion can be introduced. In this research, since it is applicable in rough terrain and it is easy to grasp the margin to the falling state, we apply the normalized energy stability margin proposed by Hirose et al. [8] to measure stability and apply it to crawling motion. As shown in Fig. 3, the normalized energy stability margin is an evaluation method that uses the value of energy required to generate a rolling motion around the support point from the current posture of the robot as an index of stability margin. Therefore, if the value of this energy increases, it can be regarded that the robot is in a stable state. In this section, we describe a gait generation method for a crawling motion with NESM as the stability index. Specifically, a method for calculating the stability margin and a method for determining the movement of the end-effector and torso will be described" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002015_tmech.2017.2713397-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002015_tmech.2017.2713397-Figure4-1.png", + "caption": "Fig. 4 CAD model illustrating D4 motor for duplex cutting", + "texts": [ + " Mode 2: u u results in self-canceling the 1st, 2nd and 6th terms in (22a). The twin-SB behaves as a (Z, , ) fine-motion 3 4 5e e eq q q manipulator [25, 26]. III. RESULTS AND DISCUSSIONS In this section, numerical and experimental results are presented to illustrate and validate the F/T model for multi-DOF load compensation, and evaluate its effectiveness for minimizing MBR in a machining application. The investigation was conducted numerically on a Direct-Drive Dual-Disc (D4) spindle motor as illustrated in the CAD model in Fig. 4, and experimentally on a testbed (Fig. 5) for duplex lathe cutting of a thin-wall disk-like workpiece (WP). The D4-motor has a Twin-SB configuration (Fig. 4), each set consists of 48 stator-EMs and 64 rotor-PMs (satisfying the assumption that both NE and NP are multiple of 4) for six-DOF actuation discussed in Section II.C. The EMs are air-cored rectangular coils wrapped with 0.5mm-diameter copper wires. The (30\u00d750\u00d710 mm) coil has a square cross-sectional area of 100mm2. Each EM is individually driven by a linear current amplifier (capable of delivering a maximum current of 4A) with a closed-loop current-control implemented on an onboard circuit (Fig", + " The EM/PM layout of the quarter D4-motor, along with the parameters characterizing the rotor dynamics, is summarized in Table 2. The disk-like WP is clamped at its outer perimeter and spins about the Z-axis with the rotor supported on a bearing set. The rotor/WP assembly (modeled as a rigid body with mass ma and moment of inertia J) is concentric with the stator. The cutting tool (CT), fed radially at rc along the X axis with feedrate v), exerts a resultant force Fc (dictated by conditions such as depth-of-cut, feedrate and spindle speed). Using the geometric parameters defined in Fig. 4, any external load exerted on the rotor can be described by (23, 24) where fX, fY and fZ are the components of Fc in the radial (feed), tangential (cutting) and normal (depth-of-cut) directions respectively: T cl a X Y a Zm f f m g f F F g (23) T c cand l Y c X c Z c Y cf h f h f r f r T r F (24) The effects of the MBR compensation on the bearing condition and cutting process were investigated by measuring the bearing sound, vibrations and spindle speed using the microphone, accelerometer and encoder as shown in Fig", + " The equations for calculating the parametric values in the LT are summarized in (30): sin / 2 1,2,4,5,6 cos / 2 3 kj P kj kj kj P kj a N c k u a N c k (30) 1 ( 3) 2 3 20.4sin 1, 3.14 where 20.4cos 2, 3.14 11.7 3, 4.18 j k j k kj j k a a k k (3 3 ) 2 ( 1) 1,2,3,6 and 1,2,3 3 k j p p c k p 1 1 32 tan 9.5tan 4 and 32 tan 9.5cot 5 j j kj j j k c k For illustrating and verifying the inverse solutions, the cutting forces (computed using a commercial software AdvantEdge to simulate the cutting condition of a 0.5mm depth-of-cut and a 0.15mm/revolution feedrate) were numerically applied (at rc and hc defined in Fig. 4) only on one-side of the D4-motor. The current inputs at different rotor displacements for generating the desired F/T vector Qe to compensate for the cutting load (23~24) were computed using three different solutions to the inverse model; optimal solution (17), current decomposition (30) and orthogonal actuation (22b). The results along with the simulation parameters are presented in Fig. 8(a, b, c). Fig. 8(a) compares the three solutions of the current inputs to 1st, 2nd and 3rd EMs. The inverse solutions are verified in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002886_j.matpr.2018.06.311-Figure5.5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002886_j.matpr.2018.06.311-Figure5.5-1.png", + "caption": "Figure 5.5. FIG: At compression ratio 18.5 Heat transfer analysis", + "texts": [], + "surrounding_texts": [ + "Fig (a) Structural Analysis: Vonmisses Stresses Fig (b) Total deformation Fig 5.4 At compression ratio 17.5 Structural analysis For compression ratio 18.5 based on thermal analysis: 19504 K. Satayanarayana et al. / Materials Today: Proceedings 5 (2018) 19497\u201319506" + ] + }, + { + "image_filename": "designv11_33_0000694_j.bbapap.2010.02.014-Figure7-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000694_j.bbapap.2010.02.014-Figure7-1.png", + "caption": "Fig. 7. Diagrammatic representations for the differential influence of Qds+ and polylysine o polylysine, the binding of Qds+ to hCA XII alters the microenvironment of the enzyme's activ the inhibitory effect of DNSA. Since such changes are not manifested upon binding of polyl", + "texts": [ + " In view of these arguments, we propose that the Qds+ mediated changes (via long range interaction) at the active site pocket of the enzyme abolishes the interaction of the Zn2+ cofactor with the sulfonamide nitrogen of DNSA. This occurs presumably due to creation of new binding site of DNSA that is somewhat removed from its original binding site at the active site pocket of the enzyme. As a consequence, Zn2+\u2013OH\u2212 is generated at the active site, and it serves as the Lewis acid\u2013base pair in facilitating the hydration of CO2 (see the cartoon of Fig. 7). Being small molecules, CO2 and H2O can easily diffuse in the vicinity of the Zn2+ cofactor, promoting the carbonic anhydrase catalyzed reaction of CO2+H2O\u21ccHCO3 \u2212+H+. This is in marked contrast to the miniscule influence of polylysine on the configuration of DNSA within the enzyme's active site pocket, and thus it does not facilitate the enzyme catalysis. The effects of differently functionalized nanoparticles (as well as their shapes and sizes) on the structural\u2013functional features of enzymes has been reviewed by Wu et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001616_ut.2011.5774088-Figure10-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001616_ut.2011.5774088-Figure10-1.png", + "caption": "Figure 10 Magnetic fluid based axis mechanism", + "texts": [], + "surrounding_texts": [ + "The vehicle should run on-land such as a desert as well as in underwater by its tracks. Their axis seal should operate these different conditions with high reliability. Conventional underwater-purpose axis sealing such as mechanical sealing, oil-filled sealing are not proper for onland purpose for a long time. Usual air-purpose watertight axis sealing methods are not proper for the professional underwater use. As shown in Figs.9-10, we proposed the MFS(MagneticFluid Seal)based axis sealing mechanism[12]. The seal blocks between the chamber A and B. The chamber A\u2019s air pressure is controlled depending on the water pressure in underwater. This air barrier prevents the sea water contaminates the MFS. In theory, this MFS has no limit to last the pressure. In proportion to the numbers of the magnetic fluid\u2019s layers, the maximum pressure\u2019s limit is increased. Usually, single axis layer could last 30-50 m depth. We carried out the MFS modeling for the simulation using the magetostatic analysis. This approach concerns the ferrofluid that is under the influence of a magnetic field that arises from a permanent magnet. TABLE I PARAMETER FOR SIMULATION Magnet relative permeability 1.05 Magnetic fluid density 1400 [kg/m3] Magnetic fluid viscosity 0.04 [Pa s] Magnetization of magnetic fluid 48000 [A/m] Magnetic susceptibility of magnetic fluid 0.03 Remanent magnetic flux density 0.8 [T] Figure 12 Test result, inner pressure changes" + ] + }, + { + "image_filename": "designv11_33_0000978_icelmach.2010.5607841-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000978_icelmach.2010.5607841-Figure4-1.png", + "caption": "Fig. 4. 3D stator and rotor end winding modeled", + "texts": [ + " Finally the matrix at strand level 11 12 13 14 21 22 23 24 31 32 33 34 41 42 43 44 s M M M M M M M M M M M M M M M M M \u23a1 \u23a4 \u23a2 \u23a5 \u23a2 \u23a5= \u23a2 \u23a5 \u23a2 \u23a5 \u23a3 \u23a6 (5) is formed, where the elements of the first row are exemplary determined as follows: ( ) ( ) ( ) 11 g,11 g,14 12 g,13 g,16 13 g,12 g,15 14 ,17 0.5 0.5 0.5 g M M M M M M M M M M M = \u22c5 \u2212 = \u22c5 \u2212 = \u2212 \u22c5 \u2212 = (6) According to the connection scheme of the strands of a machine the formulas in (6) have to be adapted. Here a R/-T/S scheme is assumed. IV. NUMERIC INDUCTANCE CALCULATION In order to classify the performance of the analytic results the end winding is also modeled in the 3D finite element calculation program ANSYS. Here the stator and rotor bars are modeled not as filaments but as solid conductors as shown in Fig. 4. Because of the enormous calculation effort and the available hardware the inductance calculation is done for single bars and not for complete coil groups or the entire rotor winding. After the determination of the inductance matrix at turn level the connection between the single bars is done via the transformations introduced before in order to get the inductances at strand level. For a first comparison between the two winding components are neglected an condition at the end face is set to paralle stator turns are loaded with direct current The differential inductances are calculated is automatically done by ANSYS: diff , d ij i j M i i H B V = \u0394 \u0394 \u0394 \u0394\u222b Neglecting iron saturation the differen (7) passes into the static one as needed her The further analysis and transformation matrices are carried out via the same me within the analytic method in section III", + " The coupling ases with increasing angular e inductance even becomes he mutual inductances recurs of the two methods are very calculated with the analytic ning the absolute values. etected regarding all other atrix wM as shown in the een the single rotor turns in dent on the rotor position. In ctances concerning rotor turn s of rotor turn number 74 concerning turns mutual inductances between larily displayed for one rotor n With increasing rotor turns (from turn 82 to 75) the inductances logically becom between turn 74 and 75 is based on the dif the bar sections that connect forward and as can be seen in Fig. 4. According to (3) the mutual inductances be rotor turns are dependent on the rotor posit the other inductances. For one rotor position the mutual induc turn to all stator turns are shown in Fig. 8 of rotor turn 74. This characteristic can also be found these mutual inductances against the rotor Therefore all elements of the inductan determined and the transformations accord can be done. Because of the calculation method only one rotor position has been comparison of the two methods. The tran in matrix sM in Table I" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001459_icetect.2011.5760158-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001459_icetect.2011.5760158-Figure1-1.png", + "caption": "Fig. 1 Shielded enclosure with aperture", + "texts": [ + " There must be access covers, doors, holes for cables, ventilation, switches, displays, and joints and seams. All of these apertures considerably reduce the effectiveness of the shield. As a practical matter, at high frequency, the intrinsic shielding effectiveness of the shield material is of less concern than the leakage through the apertures. Apertures have more effect on the magnetic field leakage than on the electric field leakage [3]. Accordingly, greater emphasis is given to methods of minimizing the magnetic field leakage. The shielded enclosure with multiple apertures is shown in figure 1. The incident wave enters the aperture with angle \u03b8. The amount of leakage from an aperture depends mainly on the maximum linear dimension, not area of the aperture, the wave impedance of the electromagnetic field and the frequency of the field. To reduce the aperture effect many small number of apertures can be used instead of a large single aperture. The rectangular slots in the shielded enclosure form slot antennas. Maximum radiation from a slot antenna occurs when the maximum linear dimension is equal to 1/2 wavelength" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002035_s12206-017-0515-4-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002035_s12206-017-0515-4-Figure6-1.png", + "caption": "Fig. 6. Mode shapes and frequencies of the LSR of the analytical models: (a) 1st mode; (b) 2nd mode; (c) 3rd mode; (d) 4th mode; (e) 5th mode.", + "texts": [ + " The material properties for vibration analysis are shown in Table 3. First, modal analysis is performed using the 3D-FEM, in which the mode shapes and their resonant frequencies can be obtained using the following equation [7, 8]: ( ){ } { }2[ ] [ ] 0i i K Mw- F = , (5) where [K], [M], \u03c9i and { }i F are the stiffness matrix, the mass matrix, the i-th natural frequency, and the eigenvector of the i-th resonance frequency, respectively. The modal shapes of the LSR of the two models and the corresponding resonant frequencies are shown in Fig. 6. A slight difference in the resonance frequency at each mode is observed. Next, vibration analysis is performed. The vibrations due to the electromagnetic force are calculated by the mode superposition method as follows [9]: { } { } { } { } 1 1 1 [ ] [ ] [ ] N N N i i i i i i i i i M y C y K y F = = = F + F + F =\u00e5 \u00e5 \u00e5&& & , (6) where [C], N, { },F yi and {f} are the damping matrix, the number of modes considered, the i-th mode shape, the displacement in modal coordinate, and the applied radial force, respectively", + " 7 shows the radial deformation spectra of the analytical models on points A and B, which are in the same location on the LSR surface shown in Fig. 5 when the speed of the LSR is 600 rpm. The rotation frequency is 10 Hz because the rotational speed of the LSR is 600 rpm. Therefore, the electromagnetic force frequencies are multiples of 260 Hz using Eq. (4). Accordingly, radial deformation occurs in the component of a multiple of 260 Hz, as shown in A and C of Fig. 7. Furthermore, resonance occurs in B and D of Fig. 7, with a slightly small excitation force of 12480 Hz, which is close to the resonance frequency of the 5th mode in Fig. 6(e). Consequently, the deformations of the SPM_CMG and the FC_CMG are generated at 12480 Hz. At this frequency, the magnitude of the generated deformation of the FC_CMG is 1.37 times higher than that of the SPM_CMG. On the basis of the result of the vibration analysis on the outside surface of the LSR, the acoustic noise is calculated by the 3D-FEM. The vibration velocities at each frequency for the harmonics of the force are imported to each node on the outside surface of the LSR. Fig. 9 shows the acoustic FEM models for the SPM_CMG and FC_CMG" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002575_2018-01-0401-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002575_2018-01-0401-Figure1-1.png", + "caption": "FIGURE 1 (a) Setup presented in [1], (b) geometry used for the CFD simulations", + "texts": [ + " Duplicating the hardware shown in [1], the CFD geometry is simplified to a rotating disk, mounted on a shaft and centered in a box with the same basic dimensions as the experimental setup. The disk is of 224.67\u00a0mm diameter and 28.5\u00a0mm width. The mesh domain includes the volume inside the box and the walls of the rotating disk and shaft. The initial condition has this rotating disk partially submerged in oil. Therefore, there is an oil/air interface on the mesh volume. Careful attention was given to the mesh around the rotating walls as described in a later section. Figure 1(a) shows the setup as presented in [1], and (b) the geometry used for the CFD model. Within Fluent, there are 2 basic approaches to solving the rotating disk submerged in fluid CFD problem. This can be done via the sliding mesh or the multiple reference frame (MRF) approach. The sliding mesh is theoretically the most accurate method for computing this type of unsteady flow. It should correctly solve the complete transient process. However, it is very computationally demanding, requiring good mesh resolution (high mesh count) at the interface and very small time steps" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001135_1077546309353917-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001135_1077546309353917-Figure3-1.png", + "caption": "Figure 3. Relationship between rigid body configuration and flexible links.", + "texts": [ + " Cubic polynomial to approximate transverse deflection: N\u00bd \u00bc 1 3x2 l2 \u00fe 2x3 l3 , x 2x2 l \u00fe x3 l2 , 3x2 l2 2x3 l3 , x2 l \u00fe x3 l2 \u00f09\u00de f eg \u00bc v1 1 v2 2 T \u00f010\u00de II. Quintic polynomial to approximate transverse deflection: N\u00bd \u00bc \" 1 10x3 l 3 \u00fe 15x4 l 4 6x5 l 5 , x 6x3 l 2 \u00fe 8x4 l 3 3x5 l 4 , x2 2 3x3 2l 2 \u00fe 3x4 2l 3 x5 2l 4 , 10x3 l 3 15x4 l 4 \u00fe 6x5 l 5 , 4x3 l 2 \u00fe 7x4 l 3 3x5 l 4 , x3 2l x4 l 2 \u00fe x5 2l 3 # \u00f011\u00de f eg \u00bc v1 1 m1 v2 2 m2 T \u00f012\u00de The relationship between the rigid body configuration and the flexible links is shown in Figure 3 (Cleghorn et al., 1981, 1984). Point A has longitudinal and transverse deflections of the coupler denoted by X1 and X2, respectively. Point B has transverse deflections denoted by X3 and X4 for the coupler and the output link, respectively. The four deflections can be expressed in terms of the transverse deflections of point A for the input link as follows (Cleghorn et al., 1981, 1984): X1 S1 2 \u00bc sin 2 3\u00f0 \u00de 2 \u00f013\u00de X2 S2 2 \u00bc cos 2 3\u00f0 \u00de 2 \u00f014\u00de X3 S3 2 \u00bc X1= tan 4 3\u00f0 \u00de \u00f015\u00de X4 S4 2 \u00bc X1= sin 4 3\u00f0 \u00de \u00f016\u00de where 2 is the transverse deflection of point A from the rigid configuration of the input link; 2, 3 and 4 are the angles of the input link, the coupler, and the output link with respect to the inertial axis X", + " Substituting equation (20) into (17)\u2013(19), the finite element displacement, rotation and curvature can be expressed as linear combinations of all the nodal curvatures { }, and they are \u00f0i\u00devxx \u00f0i\u00devx \u00f0i\u00dev 8< : 9= ; \u00bc \u00f0i\u00deN\u00f0C\u00de \u00f0i\u00deN\u00f0R\u00de \u00f0i\u00deN\u00f0D\u00de 8< : 9= ; f g \u00bc \u00f0i\u00deN f g \u00f021\u00de where the subscripts (C), (R) and (D) refer to curvature, rotation and displacement, respectively, and the subscripts x and xx represent the first and second partial derivatives of the displacement. To simply demonstrate the derivation of shape functions, the mechanism is considered as each link discretized as two equal two-node elements, and the curvature distribution of each element is approximated as a linear function, i.e., \u00f0i\u00deN\u00f0C\u00de \u00bc 1 x l , x l h i \u00f022\u00de ef g \u00bc m1 m2 T \u00f023\u00de where m1 and m2 represent the curvature at the end nodes for a beam element. The input link can be treated as a cantilever, i.e., there is no transverse deflection and the rotation at point O (Figure 3), which can be utilized to determine the integration constants. The coupler can be treated as a pinned-pinned beam, but there are transverse deflections X2 and X3 (Figure 3) appearing at both pins. Based on Figure 2, the occurrence of the transverse deflection X2 and X3 is due to the transverse deflection of point A for the input link (see equations (14)\u2013(15)). Similar to the coupler, the output link can be treated as a pinned-pinned beam, but there is a transverse deflection X4 appearing at point B. Also, there is no bending stress at points A, B and D, i.e., the curvatures are zero at these locations, which leads the number of nodal variables for the mechanism as four", + "comDownloaded from \u00f05\u00deN\u00f0D\u00de \u00bc 5l22 6 S4x R4 , l22 2 S4x R3 , 0, l4x 2 \u00fe x3 6l4 \u00f028\u00de \u00f06\u00deN\u00f0D\u00de \u00bc 5l22 6 S4x R4 , l22 2 S4x R3 , 0, l24 3 \u00fe l4x 3 \u00fe x2 2 x3 6l4 \u00f029\u00de f g \u00bc m1 m2 m3 m4 T \u00f030\u00de where the superscripts in \u00bd\u00f0i\u00deN\u00f0D\u00de represent the ith element in the mechanism (i \u00bc 1, 2 refer to the input link, i \u00bc 3, 4 refer to the coupler, and i \u00bc 5, 6 refer to the output link);R3 and R4 are the lengths of the coupler and the output link, respectively; l2 and l3 are the lengths of the beam elements for the input link and the coupler, respectively; m1 is the curvature at point O; m2, m3 and m4 refer to the curvatures at the interelement nodes for the input link, the coupler, and the output link, respectively. The longitudinal deflections of the input link and the output link are zero. The transverse deflection 2 (see Figure 3) of the input link causes the longitudinal deflection (see equation (13)), which is given as u \u00bc 5l22 6 S1, l22 2 S1, 0, 0 f g \u00f031\u00de Flexible links O A f4 f1 f6 \u2013f1 \u2013f5 f2 f3 B D Reference configuration Figure 4. Global variables of a flexible four-bar mechanism for the DFE method. Link Modulus of elasticity, E, (GN/m2) Mass density, ,(g/cm3) Link height, h,(cm) Link width, b,(cm) Link length, L, (cm) Input 69 2.70 0.424 2.54 10.2 Coupler 69 2.92 0.160 2.54 27.9 Output 69 2.92 0.160 2.54 26.7 Length of base link\u00bc 25" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003884_itsc.2019.8917468-Figure6-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003884_itsc.2019.8917468-Figure6-1.png", + "caption": "Fig. 6 In simulation environment Vrep, we establish the kinematics simulation model of the tracked vehicle and conducts simulation experiments.", + "texts": [ + " SCENE PARAMETERS AND TEST LIMIT SPEED Parameter First Second Maximum Speed a 9m 12m 5m/s b 6m 6m 7.5m/s Through the experimental data analysis, the steering degree control sequence D when facing the curve with different turning radii can be obtained, as shown in the fig. 5. By Eq.7, it can be seen that there are six MPC weight coefficients P affecting the trajectory tracking effect. In order to determine the impact of all variable parameters on the final tracking effect, the vehicle kinematics simulation model established as shown in fig 6. We design simulation experiments and build simulation scenarios in Vrep simulation software. We establish a large steering radius test trajectory, r=5.5m, and a small steering radius test trajectory r=2.5m. During the test, the simulation experiments of the two sets of steering radii are completed with the MPC control strategy under different weight coefficients. In order to verify the comparison with the control sequences in the real test scenario, each test was completed at V=5 m/s and V=7.5 m/s, which is a set of tests" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003587_1350650119868910-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003587_1350650119868910-Figure2-1.png", + "caption": "Figure 2. Schematic of textures: (a) geometrical sizes of simple texture, and (b) geometrical sizes of compound texture.", + "texts": [ + " The center deviation is evaluated through the eccentricity distance e and attitude angle . The lubricant is filled in the clearance c between the shaft and bearing to balance the load Fz. The clearance can be computed according to the relation c \u00bc R1 r, whose largest value hmax \u00bc c\u00fe e and whose smallest value hmin \u00bc c e. The symbol \u2019 represents the circumferential angular position, which is calculated from the vertical position of zero angle along the circumferential direction. The used simple texture is a cuboid groove structure shown in Figure 2(a), whose width, length and depth along the circumferential, axial and radial directions of the bearing are expressed by W, L and D, respectively. Meanwhile, based on the simple texture, the compound texture with a rectangle-semicircle secondary texture structure is used and shown in Figure 2(b). For the compound texture, its width, length and depth of the first layer texture are separately expressed by W, L and D, and its radius of the second layer texture is expressed by R. The reason to choose this type of compound texture is due to its being easily manufactured and good hydrodynamic action. The squeezed lubrication film can produce the hydrodynamic action, which can be explained with the associated governing equations. According to the used parameters in Tables 1 and 2, the real Reynolds number of the lubricant (Re \u00bc 2 nr c=\u00f060 \u00de) is about 1949" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002579_s0001925900002341-Figure16-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002579_s0001925900002341-Figure16-1.png", + "caption": "FIGURE 16. Loading system for the torsion test.", + "texts": [ + "1017/S0001925900002341 Downloaded from https://www.cambridge.org/core. Tufts Univ, on 21 Jul 2017 at 13:52:02, subject to the Cambridge Core terms of use, available at S T R E S S C O N C E N T R A T I O N I N S H O U L D E R E D S H A F T S casting for the small diameter part of the shouldered shaft. The two castings were rough machined and annealed before being joined together. A conical joint, cemented with Araldite D cold-setting resin glue, was used to ensure the best possible junction between the two castings (see Fig. 16). The glue was allowed to set by maintaining the model at room temperature for 24 hours, and was cured by a further two hours heating at 80\u00b0C. The concentricity of the large and small parts of the shaft and the accurate blending of the shoulder fillet radius were maintained by carrying out the final machining after the model had been assembled. The minimum practicable diameter of the small part of the shouldered shaft and minimum shoulder radius which could be produced accurately were 0-5 in. and 005 in. respectively, giving the parameters of the insert model as r/d=0-l, d/D=005. Three models of this type were manufactured, one being tested under each of the three loading conditions: torsion, tension and bending. A sketch of the loading system used to provide torsion is shown in Fig. 16. Similar arrangements of pulleys and dead weights were arranged to provide tension and pure bending. The loading rig was attached to a plate fitted in the oven shelf runners, while the model stood on a similar strengthened base plate. Correct alignment of the model and the loading pulleys was preserved by the use of a dowel pin fitted in the base plate and locating the axis of the shouldered shaft. The slicing procedure for each of the three loading conditions was similar to that carried out on the corresponding standard models" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000451_978-3-642-29329-0_6-Figure6.5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000451_978-3-642-29329-0_6-Figure6.5-1.png", + "caption": "Fig. 6.5 Lateral slips", + "texts": [ + "1) The non linearities of such a second order autonomous system can be found in the expressions of the forces Fy1 and Fy2, which depend on the lateral slipsa1 anda2, which in turn can be expressed as function of the state variables. The lateral slip a of a wheel is the angle between the line directed as the meridian plane of the wheel and the velocity vector at contact point of the wheel on the ground. In general a vehicle model like the one depicted in Fig. 6.1 has four different lateral slips, one per each wheel (see Fig. 6.5). We have assumed that the vehicle can be described by one single rigid body. Thus the speed of the contact point on the ground of each wheel can be derived as a function of the speed of the centre of gravity vG and of the yaw velocity r. For a conventional two wheel steer vehicle, the following relationships hold: \u2022 Front left wheel: v11 \u00bc \u00f0u 1 2 rt; v\u00fe ra\u00de \u00bc) tan\u00f0d a11\u00de \u00bc v\u00fe ra u 1 2 rt \u2022 Front right wheel: v12 \u00bc \u00f0u\u00fe 1 2 rt; v\u00fe ra\u00de \u00bc) tan\u00f0d a12\u00de \u00bc v\u00fe ra u\u00fe 1 2 rt \u2022 Rear left wheel: v21 \u00bc \u00f0u 1 2 rt; v rb\u00de \u00bc) tan\u00f0 a12\u00de \u00bc v rb u 1 2 rt \u2022 Rear right wheel: v22 \u00bc \u00f0u\u00fe 1 2 rt; v rb\u00de \u00bc) tan\u00f0 a22\u00de \u00bc v rb u\u00fe 1 2 rt In general, the radius of the path is much bigger than the track (r t) so that, recalling that u \u00bc |r| R, we have u rj jt=2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001755_ijnsns.2010.11.3.211-Figure5-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001755_ijnsns.2010.11.3.211-Figure5-1.png", + "caption": "Fig. 5: Graphical presentation of the heavy mass particle normal pressures (FN, \u03b4. material particle vibro-impact motion path along the rough circle with two radially moving limiters and with both sides limited angular elongations, caused by programmed control of the impact limiters positions \"on\" or \"off' presented in Figure 5. Concluding Remarks For all three cases of the heavy mass particle motion along the rough circle line we can identify a member in the double differential equation proportional to the square of the generalized coordinate derivation with respect to time (or parameter) by which a double differential equation of the motion is expressed. This corresponds to the known case of turbulent damping [231]. Changes of the friction force direction as an alternation is a discontinuity. This discontinuity is expressed in the alternation of the friction force direction, and in the double alternate equilibrium position as consequence of the discontinuity of the friction force direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0002321_j.compstruc.2017.11.003-Figure3-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002321_j.compstruc.2017.11.003-Figure3-1.png", + "caption": "Fig. 3. 25th mode with bending (1F) and twisting blades (1T).", + "texts": [ + " 1 illustrates the stator vane cluster of a high-pressure compressor; it is composed of eleven blades linked by both inner and outer platforms. A modal analysis has been conducted on the tuned case. Fig. 2 presents the evolution in the normalized eigenfrequencies vs. the mode number. Results have been normalized with respect to the 20th eigenfrequency. Fig. 2 indicates a high modal density area, with 15 modes being clustered within a 20% frequency range. This area is primarily composed of first flexural (1F) and torsional (1T) blade modes (Fig. 3). High modal density areas are extremely sensitive to mistuning because energy localization depends on the mistuning pattern. The eigenmodes of a stator vane cluster may therefore be completely different from one to the next. For this reason, the dynamic behavior of a stator vane cluster is difficult to predict in these areas. This paper thus seeks to statistically determine the vibratory response envelope of a mistuned stator vane cluster. For this purpose, a parametric stochastic approach [12] has been chosen to model mistuning in connection with geometric and material uncertainties" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0000105_amm.37-38.1462-Figure1-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0000105_amm.37-38.1462-Figure1-1.png", + "caption": "Fig. 1 Geometry model (a) and finite element model (b) of the piston", + "texts": [ + " Therefore, quantifying the effect of cooling gallery position on the temperature and thermal stress fields of piston can help find the suitable cooling gallery position. Taking example for the piston used in germen diesel engine DEUTZ 1015, a finite element model of piston is developed using Pro/E software and finite element analyses are achieved through ANSYS code. Numerical simulations are performed to assess the effects of cooling gallery on the temperature and thermal stress of piston top, first ring groove, pin seat, and cooling gallery surface. The study can be a reference for piston design. Geometry and Finite Element Models of Piston. Fig. 1(a) shows the geometry of studied piston. The study will focus on the outer and inner surfaces of piston crest, first ring groove, pin seat, and cooling gallery surface. The piston and the ring carrier of first ring are assembled by pressure casting. Tetrahedron element with 10 nodes is used in messing the geometry model of piston. Messing density is 5 mm for piston and 4 mm for ring carrier. In order to reduce the influence of cooling gallery change in position, the messing density in vicinity of cooling gallery is set to be 1 mm. The finite element model of piston is shown in Fig. 1(b). The finite element model includes 102,724 elements and 155,200 nodes. Material Parameters and Boundary Condition. Schematic of heat transfer boundary condition is shown in Fig. 2, and the heat transfer coefficient is given in Table 1 [6-7]. The mechanical and physical properties of material are sensitive to its temperature. The studied material is ZL8(12Si-2.5Ni-1Mg-1Cu). This metal are commonly used for diesel engine piston, pulley, strap wheel, and other parts which require high heat-durability, low thermal expansion All rights reserved" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001140_amm.110-116.4977-Figure4-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001140_amm.110-116.4977-Figure4-1.png", + "caption": "Figure 4. Intersection area of Hourglass shapes for the whole simulation time", + "texts": [], + "surrounding_texts": [ + "Maximum Angular Momentum 12Nms Maximum Angular Velocity 280 rad/sec Inertia Moment 0.00043 kgm2 Nominal Velocity 100 d/sec" + ] + }, + { + "image_filename": "designv11_33_0002404_6.2018-1965-Figure12-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0002404_6.2018-1965-Figure12-1.png", + "caption": "Figure 12: A decentralized estimation strategy with (a) directed sensor network and (b) undirected communication network with a communication graph involving two loops", + "texts": [ + "2 01 8- 19 65 \u221220 0 20 0 20 40 \u221220 \u221210 0 10 20 x (km) x 21 y (km) z (k m ) (a) \u221220 \u221210 0 10 \u221280 \u221260 \u221240 \u221220 \u221220 \u221210 0 10 x (km) x 32 y (km) z (k m ) (b) \u2212100 0 100 \u2212100 \u221250 0 50 100 \u2212100 \u221250 0 50 x (km) x 13 y (km) z (k m ) (c) Figure 10: True relative orbits (red) and estimated relative orbits using decentralized estimation strategy with (green) and without (blue) consensus feedback (mixed measurements combination 1) \u221220 0 200 20 40 \u221210 0 10 20 x (km) x 21 y (km) z (k m ) (a) \u2212100 10 \u221280 \u221260 \u221240 \u221220 \u221220 \u221210 0 10 x (km) x 32 y (km) z (k m ) (b) \u2212100 0 100 \u2212100 \u221250 0 50 100 \u2212100 \u221250 0 50 x (km) x 13 y (km) z (k m ) (c) Figure 11: True relative orbits (red) and estimated relative orbits using decentralized estimation strategy with (green) and without (blue) consensus feedback (mixed measurements combination 2) In the forgoing analysis, the observability of the decentralized estimation strategy based on a simple triangular loop is studied. However, more complicated estimation networks might be considered. One particular interesting type of network with multiple loops in the communication graph is the one in which two loops share a same measurement link (e.g. Figure 12(a)). For this type of network, the observability of one loop can transmitted to another neighboring loop with a shared measurement link. To illustrate these results, we simulate the following four types of sensor networks with varied measure- 12 of 15 American Institute of Aeronautics and Astronautics D ow nl oa de d by U N IV E R SI T Y O F A D E L A ID E - IN T E R N E T o n Ja nu ar y 16 , 2 01 8 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /6 .2 01 8- 19 65 ment types in Figure 13. The numerical results show that whenever one of the loops is guaranteed to be observable, the neighboring loop also becomes observable", + " 13 of 15 American Institute of Aeronautics and Astronautics D ow nl oa de d by U N IV E R SI T Y O F A D E L A ID E - IN T E R N E T o n Ja nu ar y 16 , 2 01 8 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /6 .2 01 8- 19 65 \u221210 0 10 20 30 \u221280 \u221260 \u221240 \u221220 0 \u221220 \u221210 0 10 20 x (km) x 24 y (km) z (k m ) (a) \u2212100 1020 40 60 80 100 120 140 \u221260 \u221240 \u221220 0 20 40 x (km) x 43 y (km) z (k m ) (b) Figure 14: True relative orbits (red) and estimated relative orbits using decentralized estimation strategy without (blue) and with (green) consensus feedback (sensor topology given by Figure 13(a) and communication topology given by Figure 12(b)) 2. For range-only measurements in a triangular loop, the strategy can help to avoid six ambiguous orbits though it allows the persistence of the type (c) mirror ambiguous orbits. 3. For mixed measurements in a triangular loop, the strategy can guarantee the uniqueness of the estimated orbits and thus ensure the observability. These conclusions are tested and veri ed by mutiple numerical examples. Furthermore, the notion that the observability of estimation loops with shared measurement link can transmit from one to another is proposed and illustrated through numerical results" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001616_ut.2011.5774088-Figure8-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001616_ut.2011.5774088-Figure8-1.png", + "caption": "Figure 8 Dimension of the experimental cargo unit", + "texts": [ + " For the sealing, two pieces are put together and decompressed the chamber\u2019s inner air with vacuum pump. Since the chamber is depressed, its O-ring is pressed. This secures the chamber\u2019s sealing. After the vehicle\u2019s diving, outer water pressure presses the O-ring. Deeper depth gives more reliable sealing due to heavy water pressure. Over 18 years, Twin-burger has successfully achieved numerous tests with this sealing method. In order to demonstrate this sealing function, a box shape experimental cargo unit was built as shown in Fig. 7. Fig.8 shows its dimensions. This polycarbonate box could use for a shallow water (less than 10m) test. The hatch opens and closes as \u2018falling-down and stand-up\u2019. Inside of the box, 2 electric winches were attached. When they wind the wire, the hatch is closed. The box has double O-rings for the reliable hatch sealing. After closing the hatch, the winches\u2019 winding makes the hatch plate to press the O-rings. Then, the inner micro vacuum pump depresses the air. It confirms the sealing. Conventional buoyancy control method uses an air tank" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0001595_tmag.2009.2032862-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0001595_tmag.2009.2032862-Figure2-1.png", + "caption": "Fig. 2. a) Magnetic flux density lines in the closed slots machine. b) Magnetic flux density lines in the open slots machine.", + "texts": [ + " With regard to the material, stator coils are made of copper, rotor bars of aluminum and the stator core of laminated Fe-Sc. The stator core presents 36 slots, which means, for a 2-pole pairs machines, 18 slots per pole pair. No slotting harmonics due to the stator are therefore to be expected since, if the number of stator (or rotor slots) per pole pair is a multiple of 3, the corresponding current harmonic does not exist. There are, however, 28 bars in the rotor, which means, for a 2-pole pairs machines, 14 slots for pole. Fig. 2(a) and (b) shows the magnetic flux density lines obtained with the FEA, respectively, in case of closed and open rotor slots, when the machine runs at steady state at no-load supplied at fundamental frequency of 1 Hz, at the rated rotor flux linkage of 1 Wb, and with a voltage carrier frequency at 1500 Hz. It can be observed that, from the contour analysis, there is no appreciable difference in the magnetic behavior of the two machines. To find the difference in the magnetic behavior of the two machines, the analysis in Section IV is thus proposed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_33_0003341_s42417-019-00128-x-Figure2-1.png", + "original_path": "designv11-33/openalex_figure/designv11_33_0003341_s42417-019-00128-x-Figure2-1.png", + "caption": "Fig. 2 Two-dimensional control volume grid", + "texts": [ + " The resolved component of the bearing reaction force along X and Y direction can be written as [7] (3)h\u0304 = 1 \u2212 \ud835\udf00X cos(\ud835\udf03) \u2212 \ud835\udf00Y sin(\ud835\udf03) + \ud835\udeff(\ud835\udf03, z\u0304). (4)\ud835\udeff(\ud835\udf03, z\u0304) = \ud835\udefc(p\u0304(\ud835\udf03, z\u0304) \u2212 1). (5) = 2paSo CE ( lo tb )3( 1 \u2212 2 ) . (6)\ud835\udeff\ud835\udf0f = CB CB + KB\u0394\ud835\udf0f \ud835\udeff\ud835\udf0f\u2212\u0394\ud835\udf0f + \u0394\ud835\udf0f CB + KB\u0394\ud835\udf0f p\u0304\ud835\udf0f(\ud835\udf03, z\u0304). (7)KB = 1 , CB = ( ) KB. (8) p\u0304 \ufffd \ud835\udf03, z\u0304 = \u00b1 L 2R \ufffd = 1 p\u0304(\ud835\udf03 = \ud835\udf0b, z\u0304) = 1 p\u0304(\ud835\udf03 = 0, z\u0304) = p\u0304(\ud835\udf03 = 2\ud835\udf0b, z\u0304) \u23ab \u23aa\u23ac\u23aa\u23ad . (9)fX = \u2212paR 2 2\ud835\udf0b \u222b 0 L\u2215R \u222b 0 p\u0304(\ud835\udf03, z\u0304) cos(\ud835\udf03)d\ud835\udf03dz\u0304, (10)fY = \u2212paR 2 2\ud835\udf0b \u222b 0 L\u2215R \u222b 0 p\u0304(\ud835\udf03, z\u0304) sin(\ud835\udf03)d\ud835\udf03dz\u0304. 1 3 Control Volume Discretization A control volume 2D grid [11] with shaded region is shown in Fig.\u00a02. The shaded area has two points each along and Z direction, respectively. After following the nomenclature in Fig.\u00a02, Eq.\u00a0(1) becomes, In Eq.\u00a0(11), the flux J and Jz is discretized as, After integration over control volume, Eq.\u00a0(11) yields, where Using neighboring coefficients of control volume, Eq.\u00a0(13) can be written as where Peclet number Pe is the ratio of convection and diffusion terms. It is written as follows, (11) \ud835\udf15J\ud835\udf03 \ud835\udf15\ud835\udf03 + \ud835\udf15Jz \ud835\udf15z\u0304 + 2\ud835\udeec\ud835\udf08 \ud835\udf15 \ud835\udf15\ud835\udf0f ( p\u0304h\u0304 ) = 0. (12)J\ud835\udf03 = ( \ud835\udeecp\u0304h\u0304 \u2212 p\u0304h\u03043 \ud835\udf15p\u0304 \ud835\udf15\ud835\udf03 ) , Jz = ( \u2212p\u0304h\u03043 \ud835\udf15p\u0304 \ud835\udf15z\u0304 ) . (13)Je \u2212 Jw + Jn \u2212 Js + SP = 0, Je = ( \ud835\udeecp\u0304h\u0304 \u2212 p\u0304h\u03043 \ud835\udf15p\u0304 \ud835\udf15\ud835\udf03 ) e \u0394z\u0304, Jw = ( \ud835\udeecp\u0304h\u0304 \u2212 p\u0304h\u03043 \ud835\udf15p\u0304 \ud835\udf15\ud835\udf03 ) w \u0394z\u0304, Jn = ( \u2212p\u0304h\u03043 \ud835\udf15p\u0304 \ud835\udf15z\u0304 ) n \u0394\ud835\udf03, Js = ( \u2212p\u0304h\u03043 \ud835\udf15p\u0304 \ud835\udf15z\u0304 ) s \u0394\ud835\udf03, SP = 2\ud835\udeec\ud835\udf08\u0394\ud835\udf03\u0394z\u0304 \u0394\ud835\udf0f (( p\u0304Ph\u0304P )\ud835\udf0f \u2212 ( p\u0304Ph\u0304P )\ud835\udf0f\u2212\u0394\ud835\udf0f) " + ], + "surrounding_texts": [] + } +] \ No newline at end of file